Text

Plant-Specific Analysis in Response to NRC Bulletin 88-011, Pressurizer Surge Line Thermal Stratification, Davis-Besse Nuclear Power Station,Unit 1
	ML20085D147
Person / Time
Site:	Davis Besse
Issue date:	09/30/1991
From:	Chandler C, Costa D, Moore R; BABCOCK & WILCOX CO.
To:	;
Shared Package
ML20085D144	List: ML20085D143; ML20085D147;
References
	BAW-2127-S01, BAW-2127-S1, IEB-88-011, IEB-88-11, NUDOCS 9110150282
	Download: ML20085D147 (119)
	v • d • e

BAW 2127 Supplement 1 September 1991 I.

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Plant Specific Analysis in Response to Nuclear Regulatory Commission Bulletin 8811

" Pressurizer Surge Line Thermal Stratification" Davis Besse Nuclear Power Station Unit 1 B& W NUCLEAR SERT / ICE COMPANY TMf'TM J!.2%

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IIAW 2127 Supplement 1 g

September 1991 I

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I l'lANT-SI'l!Cil:lC ANAL.YSIS l

IN llliSPONSli TO NUCl..liAlt liliGUI.ATol(Y COMMISSION llUI.LIiTIN 8811 "PilliSSUltlZIllt SUl(GII 1 INii TlllilthiAl. STi(ATil;lCATION" Prepared for Toledo lidison Company Prepared by (see Section 10 for document signatures)

It W. Moore C. K. Chandler il li. Costa G. l.. Weatherly 11&W Nuclear Service Company P.O.1. lox 10935 l-ynchburg, Virginia 24506-0935

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l EXECUTIVE SUMt.RY Or. December 20, 1988 the Nuclear Regulatory Commission issued NRC Bulletin 88-11.

The bulletin addressed technical concerns associated with thermal stratification in the pressurizer surge line and required utilities to establish and implement I

a program to ensure the structural integrity of the surge line. The B&W Owners Group (B&WOG) has developed a comprehensive program to address the requirements of the bulletin.

This program and its results were summarized in BAW-2127,

" Final Submittal for Nuclear Reculatory Commission.Bulletin 83-11 "Pressur.iZE Surae Line Thtrmal Sir _iti_1fication" for the B&W-designed lowered-loop plants.

Davis-Besse Nuclear Power Station Unit 1 (DB-1), the only operating B&WOG raised-loop plant, requires a plant-spe'ific eve %ation which is summerized in this report supplement. As with the lowered-roep plant evaluation, the plant-specific I

evaluation for Davis-Besse Unit 1 invo'we., comprehensive instrumentation.

The evaluation also involved assessment ci operating practices and procedures, collection and review of historical p ant data, and develvpment of new design basis transient conditione for the surge line to conservatively account for

. thermal cycling, therm,.I uratification, and thermal striping.

The evaluation of thermal striping incarporated the best available data to characterize this phenomenon as it may occur in the surge line.

The structural analysis for the new design conditions has shown that the Davis-I Besse Unit i urge line can meet its 40-year design life given the completion of procedural and design modifications. Detailed finite element analyses have been performed on the pressurizer surge nozzle, on the surge line to hot leg nozzle, and on the limiting portions (the elbows) of the pressurizer surge line piping.

l At all points in the surge line and the associated nozzles, the cumulative fatigue usage factor remains less than one for the design life.

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I MRLQf CON 1ENTS b

EXECU11VE

SUMMARY

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INTRCDUC110N 1-1 1.1 Dack round.

1-2 s e 3 1.2 Conc usion 1-3 2.

OVERVIEW Of BW ORif s W/.T PROGRAM 2-1 2.1 Development of New Design Bats Conditions 2-1 1

2.2 Stress Analysis 2-3 3.

PLANT-SPEClfl0 APPROACH TOR DAVIS-BLSSE UNil 1 3-1 1

3.1 Comparison of Configurations 3-1 3.2 Plant Operations 3-2 3.3 Conclusion 3-2 4.

DEVELOPMEN1 Of NEW DESIGN BASIS FOR SURGE LINE 4-1 4.1 Instrumentation of Davis-Besse Unit 1 Surge Line 4-1 4.2 Correlation of Surge Line Temperatures 4-2 1

4.3 Thermal Striping 4-7 4.4 Review of Operational History 4-7 4.5 Development of Revised Design Basis Transients 4-10 4.5.1 Heatup Transients 4-13 I

4.5.2 Cooldown 1ransients 4-22 4.5.3 Other Design Transients 4-26 4.6 Design Transients Summary 4-29 5.

PIPING ANALYSIS 5-1 5.1 Structural Loading Analysis 5-1 I

5.1.1 Mathematical Model 5-1 5.1.2 Non-Linear Temperature Profile 5-2 5.1.3 Verification Run for Displacements 5-3 5.1.4 Structural loading Analysis for the Thermal Stratification Conditions 5-3 5.2 Generic Stress indices for the Surge Line Elbows 5-4 5.3 Verification of NB-3600 Equations (Equations 12 and 13, and Thermal Stress Ratcheting) 5-5 5.4 Development of Peak Stress 5-7 5.4.1 Peak Stresses Due to fluid flow 5-8 5.4.2 Peak Stresses Due to lhermal Striping 5-8 5.4.3 Peak Stresses Due to the Non-Linearity of the Temperature Profile 5-9 5.5 f atigue Analysis of the Surge Line 5-9 5.6 f atigue saalysis Results for the Surge Line 5-11

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F (DlRUllS (Cont'd)

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N0ZZLE ANALYSES 6-1 6.1 Pressurlier Surge Nozzle 6-1 6.1.1 Methodology Hodifications f rom Lowered-loop Analysis 6-1 6.1.2 Sunnary of Results and for.:lusion,

6.i 6.1 Hot leg Surge Nozzle 6-4 6.2.1 Geometry 6-4

'I 6.2.2 Description of Loadings 65 6.2.3 Discussion of Analysis 6-6 6.2.4 List of Assumptions / Inputs Used in Analysis 6-9 6.2.5 Thermal Analysis of Axisynnetric loads 6-10 6.2.6 Stress Analysis of Axisynnetric Loads 6-13 6.2.7 Stress Analysis of Non-Axisynnetric Loads 6-14 6.2.8 ASME Code Calculations 6-15 6.2.9 Sunnary of Results and Conclusion 6-16 i

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SUMMARY

Of RESULTS 7-1 8.

BASES FOR THE DAVIS-BESSE 1 ANALYSIS 8-1 9.

REFERENCES 9-1 10.

DOCUMENT SIGNATURES 10-1 APPENDIX A A-1 I

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Ll51 DF 1ADLLS IL Table Page 4-1 Surge Line Design Basis irarsieat List 4-31 4-2 Design Transients - Number of Events 4-35 4-3 Events Afft.'. ting Surge t ine flew for Plant Heatup and Cooldown 4-37 4-4 Summary of Results for Thermal Transient Parameters

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(Lower Horizontal) 4-41 5-1 Measured and Equivalent Linear Temperature Profiles 5-12 5-2 Total fatigue Usage factors for the Davis-Besse Unit 1 Surge Line 5-13 6-5 Stresses and fatigue for a lypical PV at the Lnd of the 1

Nozzle Taper 6-17 7-1 Usage factor Results for Davis-Besse Unit 1 7-2 A-1 Signal identification A-5 I

LISlALL10@iS figure Page 4-1 Design Heatup Transient Temperatures 4-43 4-2 Design Cooldown Transient Temperatures 4-44 5-1 Surge Line Mathematical Model 5-14 5-2 (omparison of Surge Line Displacements (X-Direction) 5-15 5-3 Comparison of Surge Line Displacements (Y-Direction) 5-16 I

5-4 Comparison of Surge Line Displacements (Z-Direction) 5-17 6-8 Geometry of Ilot Leg Surge Nozzle 6-18 6-9 Finite Element Model of Hot Leg Surge Nozzle 6-19 6-11 Location of Delta-T Values 6-20 I

6-12 Hot Leg Surge Nozzle Temperature Contours (f) for a Typical PV 6-21 6-13 Location of Stress Classification Lines 6-22 8-1 Surge Line Operational Limit 8-5 I

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Davis-Besse Data Acquisition Hardware Configuration A-9 A-2 Instrumentation Locations at Davis-Besse Unit 1 A-10 A-3 Thermocouple Positions A-ll

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INTRODUCTION r

This report is a supplement to [inal Suhmittal for Nuclear Reaulatory Commission

[ht}ltt in 88-11 " Pressurizer Surae line Thermal Stratification" (Reference 1) and summarizes the B&W Owners Group program addressing the technical issues described in NRC Bulletin 88-11 (Reference 2) for the Davis-Besso Nuclear Power Station Unit 1.

The analyses described in this report confirm that all surge line pressure boundary components (including all nozzles) will satisfy applicable code stress allowables for the Davis-Besse Unit 1.

The introduction briefly reiterates the background for the thermal stratification, striping and cycling issues, and a summary of the plant-specific surge line fatigue analysis results and conclusions for Davis-Besse Unit 1.

The remaining sections of the report are as follows:

a Section 2 reviews the technical approach which has been developed by the B&W Owners Group, a

section 3 discusses the plant-specific approach for Davis-Besse Unit 1, a

Section 4 describes the development of the new design basis thermal-hydraulic conditions for the Davis-Besse Unit I surge line, a

Sections 5 and 6 describe the stress and fatigue analyses performed for the Davis-Besse linit I surge line piping and its nozzles, a

Section 7 provides conclusions resulting from the Davis-Besse Unit 1 plant-specific analysis with regard to new design basis transients which I

represent surge line thermal conditions and the structural integrity of the surge line, a

Section 8 states the conditions which form the basis for the analysis, a

Section 9 lists all references, and s

Appendix A provides a supplement to Section 3.1 and contains a detailed discussion of the data acquisition for the Davis-Besse Unit 1 surge line.

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LL_farlstand The surge 'ine in B&W 17i fuel assembb (177 FA) plants, including Davis-Besse Unit 1, contains approxiniately 50 feet of piping which connects the pressurizer lower head and the reactor coolant hot leg piping. During plant operation, the reactor coolant system (RCS) is pressurized with a steam bubble in the pressurizer. Thus, the pressuriter contains saturated fluid while the remainder of the RCS is subcooled with temperatures cooler than the pressurizer fluid by I

43*f or more.

The surge line provides the means by which the pressurizer accotr.mdates changes in RCS inventory.

The reactor conlant flows through the surge 'ine during arges into and out of the pressurizer. During reactor coolant pump cperations, ther

's normally a small outflow from the pret. Jrizer due to l

continuous minimum pressurizer spray flow.

Due to differences in density, the reactor coolant can stratify in the horizontal pipir.g tactions whereby the fluid temperature varies f rom top to bottom with the warmer fluid located above the denser (cooler) fluid. This phenomenon, known as I

thermal stratification, is most pronounced during outsurges from the pressurizer.

During an insurge or outsurge under.tratitied conditions, thermal striping may occur at the fluid layer interface.

Thermal striping is a rapid oscillation of the thermal boundary interf ace caused by interfacial waves and turbulence l

effects. The original surge line fatigue analyses performed f or the R&W 177-fA plants did not account for thermal stratification which causes additional bending moments in the piping, nor did the analyses account for thermal striping which affects the fatigue usage at the inner surface of the pipe.

I In order to confirm pressurizer surge line integrity, the Nuclear Regulatory Commission issued NRC Bulletin Number 88-11, Prenitrller_Strne line_lhermal Eintification (December 20, 1988).

This bulletin requires certain actions of licensees of all operating pressurized water reactors (PWRs).

The applicable actions are paraphrased below:

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At the first available cold shutdown after receipt of the bulletin, and l

which exceeds seven days, conduct a visual inspection of the pressurizer surge line.

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Within four mMths of receipt of the bullet %, licensees of plants in operation over ten years are requested to demonstrate that the pvessurizer surge line meet 5 the applicable design codes and other TSAR and regulatory connitments for the itcensed life of the plant, considering thermal stratification and thermal striping in the fatigue and stress evaluations; or provide the staff with a justification for continued operation while a detailed analysis of the surge line is performed that implements items Ic and Id below.

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If necessary, obtain plant-specific surge lit.e thermal and displacement data.

Data can be obtained through collective eff orts if suf ficient similarities in geometry and operation can be demonstrated.

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Update the f atigue and stress analyses to ensure compliance with the I

applicable Code and Regulatory requirements within two years of receipt of the Bulletin or submit a justification for continued operation and a description of the proposed correttive actions for ef fecting long-term resolution.

A portion of the B&W Owners Group program was presented to the Nuclear Regulatory Commission Staff on September 29, 1988 and April 7,1989. An interim evaluatiori, BAW-2005, dated May 1989, provided the staff with a justification for near term operation for all of the operating BiW 177-fA plants (Reference 3).

The NRC I

I concluded that sufficient information had been provided to justify near term operation for B&W plants until the final repot t sould be completed (Reference 4).

The final report for the lowered-loop plants was completed and submitted to the NRC in December 1990. This supplement summarizes the plant-specific evaluation for the Davis-Desse Unit I raised-loop plant, and documents compliance with s

actio items Ib, Ic, and Id of NRC Bulletir. W ii 12 Enn u l19a Given the completion of the procedural and design modifications described in Section S, the surge line for Davis-Besse Unit 1 is shown to fulfill the 40-year 1-3 l

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licensed plant life.

The structural analysis of the surge line and associated nozzles has accounted f or thermal conditions (thermal stratification, thermal striping, and thermal cycling) existing during the life of the plant.

The highest cumulative usage factor for 40 years of operation (?40 heatup/cooldown cycles) has a value of 0.93 and occurs in the nozzle-to-head corntr oi the I

pressurtzer norzle. The se:ond highest cumulative usage factor for the 40 years of operation occurs in the nozzle-to-hot leg corner of the hot leg nozzle and has a value of 0.76.

Within the surge line proper, the highest cumulative usage f actor is 0.62 and occurs in the straight pipe in the lower horizontal run just past the first elbow (elbow A on figure 5-1).

The cumulative usage factor ar the snubber stanchion (a welded attachment) is 0.08.

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OVERVIEW Of B&W OWNER'S GROUP PROGRAM The B&W Owner's Group Haterials Committee report, hereafter referred to as the main report (Reference 1), includes a detailed discussion of the program developed to address the technical concerns identified in NRC Dulletin 88-11.

The discussion of the program will not be repeated in this supplement, but will be reviewed briefly. The program is divided into two casic sections: the design basis thermal transients and structural analyses required to assess the integrity I

of the surge line and associated nozzles for the balance of the design life of each of the plants. This supplement is a summary of the program that addresses I

the plant-specific evaluation for the Davis-Berse Unit I raised-loop plant. The key elements of the program are as shown in figure 2-1 of the main report except that the surge line data is taken at Davis-Besse Unit 1.

21 De v e_l o nmenLnLutw_Eq11gn_LuiLCamillions The thermal-hydraulic phenomena which must be accounted for in the surge line are l

thermal stratification, thermal striping, and thermal cycling.

As these phenomena occur to some degree in almnst all modes of plant operation, the surge line conditions must be carefully considered from cold shutdown through heatup, I

power escalation, normal power operation, and cooldown.

Thermal cycling i; associated with coolant mass and temperature changes in the reactor coolant system (RCS), Thermal stratification can occur in the surge line l

only during moderate to low flow rates through the surge line and may exist in a steady state as wel' as in a transient condition.

Thermal striping requires l

the existence of thermal stratification.

The main report considers the requirements for a quantitative treatment of these phenomena in Subsection 2.1.

Davis-Besse Unit I was instrumented to record the thermal transients in the surge line for plant heatup, power escalation, full power operation, and plant 2-1

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cooldown, Surge line displacement instrumentation was also added to the surge litie. The Davis-Besse Unit 1 data collection process, described in more detail g

5 in Appendix A, provided circumferential temperature measurements at several axial locations along the surge line in addition to displacement measurements for each major uisplacemant axis.

As performed for the lowered-loop plants, a review of the operating procedures provided a better undcrstanding of those plant evolutions likely to cause surge I

line upsets.

The operating procedure review and consideration of the hourly j'

heatup and cooldown data over the plant operating history provided the bases for i

generating design basis surge line transients for plant heatup, cooldown, and power operation.

.oii 2.1 of the mair report, an important part of the t

As described in Sub:

operating procedure ceview dealt with the potential upper bound for the

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pressurizer to hot leg temperature difference as described in the lowered-Ictp analysis.

For Davis-Besse Unit 1, the maximum calculated surge line top-to-bottom temperature difference is 358F for the design transients, as discussed in Subsect ion 4.5.1.2.

The maximum surge line top-to-bottom temperature difference measured at Davis-Besse Unit I was 253F.

The relationship of thermal striping amplitude and frequency to the pipe fluid conditions for Davis-Besse Unit I are based on the Battelle data as discussed in Subsections 2.1 ard 4.3 of the main report, lhe thermal striping data correlation permits the determination of striping characteristics for any given surge line flow rate and imposed top-to-bottom temperature difference.

The product of the thermal-hydraulic program is a revised set of curge line design I

basis transient descriptions that account for thermal

cycling, thermal stratification, and thermai striping.

Design transients censidered in the previous design basis for the surge line were aodified to account for all three thermal phenomena. All design basis transients involving surges were censid' 'd in the evaluatioi..

Results of the thermal-hydraulics, part of the program consists of the input for the stress analysis of the surge line itself, the associated nozzles at each end, 2-2

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I horizontal run.

The stress analysis portion of the program is described in the next subsection.

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The stress analysis procedure is essentially the same as that used 'or the lowered-loop plants.

The first phase of the stress analysis involved building

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a structural mathematical model containing the pressurizer, the surge line, the hot leg, the reactor vessel and the steam generator.

This structural l

mathematical model was verified by using the measured surge line temperature data from the Davi s-11e s se Unit I heatup of June 1990 to predict surge line displacements.

These predicted surge linc displacements agree well with the measured surge line displacements (see Sul ection 5.1.4 and figures 5-2, 5-3. and 5-4).

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The structural loading analysis was performed using the new thermal-hyd aulic l

design basis and considering potential surge line whip restraint interference with the gaps set at the as-measured (onditions for each of the eight whip restraints. The internal forces and moments were generated from the structural l

loading analysis and were used for the stress analysis of the surge lins and the nozzles associated with its endpoints.

Load.ng cases were developed for each period of history using the measured restraint gap data representative of the period.

The applicable piping code is the 1986 Edition of ASML Lode NB-3600, in accordance with NRC Bulletin 88-11 which states:

" fatigue analysis should be performed in accordance with the latest ASME Section til requirements incorporating high cycle fatigue." A Code reconciliat%n was performed wita a raview of the surge line stress report for Davis-Besse Unit 1.

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Using finite element analysis for the elbows and simplified equations elsewhere in the surge line, all stress intensity values (Equations 12 and 13, and Thermal I

Stress Ratchet) were found to be within the allowables (taken from ASME Sectiou III, Appendix I). Therefore, the elasto-plastic fatigue analysis was perf ormed g

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F in accordance with Nth 3653.6(c). lo account f or the thermal-hydraulic conditions defined in the new design basis, the surge line f atigue analysis includes thermal stratification, pressure ranges between the thermal stratification conditions, thermal striping, fluid flow and temperature changes leading to through-wall temperature gradients, and the additional localized stress due to the non-linearity of the top-to-bottom temperature profile.

I in the NB-3600 simplified elasto-plastic fatigue analysis, all applicable surge line locations were analyzed,

,n ading the drain line nozzle and the snubber stanchion which were considered as branch connections, lhe total cumulative usage factor is less than 1.0 at all surge 'ine locations.

I In addition to the st ruc t ural analysis of the surge line described above, detailed stress analyses of the pressurizer and hot leg nozzles were pe: f ormed to demonstrate compliance with the ASML Code, Section 111.

I f inite element models were made of both nozzles and the thermal and pressure stresses were calculated using the revised design basis transient descriptions as input.

Piping loads acting on the nozzles were taken from the structural analysis of the surge line and were combined with the pressure and thermal I

Stress and f atigue analyses were performed in accordance with the stresses.

requirements of the 1986 Edition of the ASME Code, Section 111, NB-3200 and NB-3600.

The analyses demonstrate that the cumulative usage f actor for each nozzle is less than 1.0.

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3.

Pl. ANT-SPECIFIC APPROACH FOR DAVIS-BESSE UNIT 1 f

Section 3 of the main report (Reference 1) describes the decision to evaluate the lowered-loop plants generically and the need for o plant-specific approach for the Davis-Besse Unit 1 plant.

The assessment addressed two different types of factors: (1) those that are inherent in the equipment design and (2) the plant-specific operating and surveillance procedures that may influence the surge line conditions.

L.1 Compr_11ED_ elf 0nfi9REALL0ni As descrioed in Section 3.1 of the main report, the lowered-loop plants are sufficiently similar to be evaluated generically.

However, the differences between Davis-Besse Unit I and the lowered-loop plants led to the decision to install special instrumentation at Davis-Besse to gather data during the heatup from the 6th refueling outage in the summer of 1990. Consideratio,is leading to this decision were as follows:

a The Davis-Besse surge line configuration differs significantly from the lowered-loop plants as shown in figures 3.1 and 3.2 of the main report.

Tho iower horizontal run is somewhat shorter, ar.d there is an upper horizontal run in excess of 20 feet compareo to 2.5 feet in the lowered-loop plants.

n The surge line at Davis-Besse Unit 1 incorporates eight fixed pipe whip restraint structures, with impact collars clamped to the surge line.

in addition to placing a limit on pipe displacement, these impact collars interrupt the insulation, permitting gaps on either side.

The increased heat loss affects the magnitude of the stratification temperature differences.

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a At Davis-Besse the power-operated relief valve (PORV) inlet condensate drain is connected to the surge line drain upstream of the drain isolation valve.

Condensate reflux into the surge line depends upon heat losses I

from the line, and could have some influence on the surcje line strstification response.

3.2 P1 ant Oprittigni Section 3.2 of the main report considers the plant operational aspects of a generic evaluation of the B&WOG plants.

1he magnitude and number of thermal l

cycles. applied to the pressurizer surge line were evaluated to formulatt the I

design basis cycles, lhe evaluation included review of applicable plant operating procedures and plant data as well as interviews of the plant operators.

Plant data from the instrumented Davis-Besse Unit I surge lint

'd historical operating data for plant heatup and cooldown events for Davis-Besse Unit 1 I

provide most of the bases for describing the design transients for Davis-Besse Unit 1.

All of the B&W plants operate in a similar f ashion as described in the main report; however, certain differences between the sets of design transients for Davis-Besse Unit 1 and those for the lowered-loop plants have resulted from this evaluatien, lhese differences are discussed in detail in Section 4.4.

3.3 Conchtsion While the lowered-loop plant configuration and plant opeiations are quite similar and a generic development of design basis transients is justified, Davis-Besse Unit 1, which is a raised-loop B&W plant, requires a plant-specific analysis due to the differences discussed in Section 3.1.

The analysis for Davis-Besse Unit 1 is addressed in this supplemental report.

The methodology described in the main report is generally applicable to the Davis-Besse Unit 1 analysis. This includes the correlation of stratification and striping, the synthesis of design transients, the structural modeling techniques, the structural loading analysis, and the fatigue analyses of the surge line and its associated nozzles.

Differences from the material contained in the main repot due to plant-specific structural and operating conditions are identified and,n

fied in this supplement.

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DEVELOPMENT OF NEW DESIGN BASIS FOR SURGE LINE 4.1 Instrumentation of Davis-Bene Unit 1 Surag_line Plant-specific thermal and displ cement data for the Davis-Besse Unit I surge line were collected for the following reasons:

The original Davis-Besse Unit 1 instrumentation does not supply sufficient a

data for an understanding of the thermal conditions throughout the line, s

Differences inherent in the design of Oconee Unit I and Davis-Besse Unit I result in different thermal conditions for the two surge line l

configurations, a

The surge lines have differences in geometric

layout, draining arrangements, piping supports and restraints, and insulation.

The objectives and the technical approach for data collection have been identical I-for the Oconee and Davis-Besse instrumentation programs. The objectives of the Davis-Besse surge li a instrumentation program have been to determine:

a The magnitude of the thermal stratification including the maximum top-to-bottom piping temperature differential, a

Variations in the thermal stratification with axial position along the surge line, a

The changes in surge line displacement that result from thermal stratification, The plant operations that cause thermal stratification cycles, and a

a The temperature response of the surge line to changes in surge line conditions.

To meet the objectives, a comprehensive instrumentation package was installed that has included 46 thermocouples mounted on the outside circumference of the surge line and 14 displacement instruments affixed to various parts of the line.

The thermocouples and displacement instruments were connected to a data 4-1

I acquisition system allowing continuous monitoring of all instruments.

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addition, numerous permanent plant computer signals were recorded with this data acquisition system and the plant computer system. Details of the instrumentation and the data acquisition system are included in Appendix A.

1 The instrumentation package and data acquisition system were installed in April and May of 1990 during the Sixth Davis-Besse Unit 1 Refueling Outage (6Rf0).

Data were recorded as the plant prepared for and went into its normal heatup in g

early June.

There was no interference with normal plant operations and no I

changes to procedures were made to accommodate the data acquisition or to reduce the effects of potentiai thermal stratification. Data were recorded throughout the heatup, power escelation, and for several days near full pc..er.

l L.2__lprrelAligLgLhroe line Temancthtn The correlation of the invis-Besse Unit I surge line temperatures versus plant l

conditions paralleled that of the lowered-loop plants, as described in Section 4.2 of the main report.

Measurements and nithods specific to the Davis Besse analysis are described below.

Elant Tempera ure He wiramelti t

The Davis-Besse Unit I surge line temperatures were measured at eight cross sections distributed throughout the horizontal portions of the line, as shown in I

figure A-2, At six of these instrumented cross sections, seven thermocouples were distributed over the pipe circumference to provide uniform and complete coverage of the temperature profile across the height of the pipe; top-and bottom-of-pipe thermocouples were used at the remaining two locations.

I Surge line temperatures were recorded at twenty-second intervals through much of June and July 1990.

These measurements spanned two heatups, a cooldown, power escalation, and operation near full power.

For ease of handling end analysis, the interval of almost continuous data from 10 June through 12 July were divided I

into the following three data sets:

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O June 10-18, 1990 June 23-30, 1990 July 1-12, 1990 e

b lhe major plant conditions during data collection are outlined below.

The indicated events were identified by reviewing logs, and by examining time-based traces of spray valve position, spray line temperature, pressurizer level, core power, primary flow rate, and reactor coolant pump power.

411m._10dL_1919_DALA Th' Nne 13-18,1990 data encompassed a heatup.

Initially, the plant was cold tsurized with nitrogen.

Considering time zero as 0000 hours0 days 0 hours 0 weeks 0 months , on June a pressurizer steam bubble was established at about 20 hours2.314815e-4 days 0.00556 hours 3.306878e-5 weeks 7.61e-6 months , pumps were gerated briefly and independently for venting, between 83.7 and 85.3 hours3.472222e-5 days 8.333333e-4 hours 4.960317e-6 weeks 1.1415e-6 months .

I Reactor coolant heatup was begun at 122 hours0.00141 days 0.0339 hours 2.017196e-4 weeks 4.6421e-5 months using 2 pumps; a third pump was added at 146.5 hours5.787037e-5 days 0.00139 hours 8.267196e-6 weeks 1.9025e-6 months and the fourth at 202 hours0.00234 days 0.0561 hours 3.339947e-4 weeks 7.6861e-5 months , d1UIL2h39.19.9LQ31a The measurements of June 23-30, 1990 began with a plant cooldown in progress, Two reactor coolant pumps were operating until 6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months .

Auxiliary spray was actuated intermittently between 6,6 and 33.3 hours3.472222e-5 days 8.333333e-4 hours 4.960317e-6 weeks 1.1415e-6 months .

The pressurizer was blanketed with nitrogen beyond 20 hours2.314815e-4 days 0.00556 hours 3.306878e-5 weeks 7.61e-6 months . Conditions remained largely quiescent through 90 hours0.00104 days 0.025 hours 1.488095e-4 weeks 3.4245e-5 months . Then the pressurizer level was reduced, and a steam bubble was drawn beyond 95 hours0.0011 days 0.0264 hours 1.570767e-4 weeks 3.61475e-5 months . The reactor coolant pumps were operated individually and briefly for venting, between 110,4 and 111.1 hours1.157407e-5 days 2.777778e-4 hours 1.653439e-6 weeks 3.805e-7 months .

The heatup was started at I

123.3 hours3.472222e-5 days 8.333333e-4 hours 4.960317e-6 weeks 1.1415e-6 months using 2 pumps; a third pump was setivated at 140.5 hours5.787037e-5 days 0.00139 hours 8.267196e-6 weeks 1.9025e-6 months , and the fourth at 157 hours0.00182 days 0.0436 hours 2.595899e-4 weeks 5.97385e-5 months .

Spray was activated briefly at 165.3 and 175.2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months .

Minimum continuous spray flow was interrupted for approximately I hour starting at 169.2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months .

'All references to times for various changes in plant conditions are referenced to 0000 hours0 days 0 hours 0 weeks 0 months of the first day of the data collection period.

4-3

J_u1 y 1-121_1129_D.aLa L

The data taken during July 1-12, 1991 involved power escalation.

Four reactor coolant pumps were operated throughout the measurement period, except for 3-pump operation from 38.4 to 38.7 hours8.101852e-5 days 0.00194 hours 1.157407e-5 weeks 2.6635e-6 months .

Spray was used intermittently, and was maintained for 8 hours9.259259e-5 days 0.00222 hours 1.322751e-5 weeks 3.044e-6 months starting at approximately 210 hours0.00243 days 0.0583 hours 3.472222e-4 weeks 7.9905e-5 months .

Power operation began at (22 hours2.546296e-4 days 0.00611 hours 3.637566e-5 weeks 8.371e-6 months ; power was increased beyond approximately 40% of full power at l

105 hours0.00122 days 0.0292 hours 1.736111e-4 weeks 3.99525e-5 months , and beyond approximately (>01 at 172 hours0.00199 days 0.0478 hours 2.843915e-4 weeks 6.5446e-5 months , lhe acquired plant data were extensively cross-plotted and compared. The local temperature distributions, the sequential response of temperature versus location in the surge line, and their responses to insurges and outsurges generally I

confirmed the sensitivity and self-consistency of the surge line temperature measurements.

The Davis-Besse Unit I surge line temperature measurement s were processed in the same fashion as the Oconee measurements, as described in Section 4.2.2 of the main report.

interface elevations and both local and extreme temperature differences were extracted from the data.

Maximum top-to-bottom temperature dif ferences were determined separately for the lower-and upper-elevation piping runs.

I EDItalAUDB lhe Davis-Besse Unit I surge line correlations were developed in much the same i

fashien as those of the lowered-loop plants, as described in Section 4.2.3 of the main report.

The major plant conditions affecting the surge line temperatures were:

Surge line flow rate (or pressurizer level versus t ime) a System pressure (or saturation temperature) s a

flot leg temperature The supplementary plant conditions included: reactor coolant pump status, spray status, magnitude of pressuriter level oscillations, core power level, and the status of the decay heat removal system during a cooldown.

Correlations were 4-4

I l

developed for the elevation of the thermal interface in the lower-and upper-elevation piping runs, as well as for the following temperatures:

a Top-and bottom-of-pipe, lower-elevation run a

Top-and bottom-of-pipe, upper-elevation run a

Pressurizer nozzle (fluid)

I Each of these temperature correlations pertained to a surge line piping outside metal temperature.

The exception is the pressurizer nozzle fluid temperature, as dese hed below. Additionally, an estimate of the riser average temperature was fv'.sd from the bracketing pipe temperatures.

The temperature at the hot l

leg-to-surge line nozzle was taken to be that of the top of pipe in the upper-elevation run. This temperature was used to determine the temporal extremes of the hot leg-to-surge line nozzle temperature.

The top-to-bottom temperature difference in the upper-elevation piping was used to characterize the stratification temperature difference at the nozzle.

I The pressurizer nozzle correlation provided a direct estimate of the temperature of the nozzle fluid, rather than metal.

This correlation basically varied the nozzle fluid temperature toward the temperature of the source fluid in proportion to the volume of fluid displaced during a flow event.

The source fluid temperature for an outsurge was the saturation temperature.

However, the temperature at the nozzle required treatment of the 200 ft' vo'ume of the lower pressurizer which is located below the pressurizer heaters and may contain coolant below the saturation temperature.

The current outsurge fluid I

displacement was obtained by integrating the preceding surge line volumetric flow rates. The predicted nozzle fluid temperature thus increased toward the current saturation temperature as the current outsurge displacement approached 200 ft'.

Insurge predictions were handled in two phases.

The first phase involved the temperatures predicted for the lower-elevation piping; the source temperature for the second phase was the hot-leg temperature. The associated fluid volumes were 5 f t' (approximately one-half of the volume of the lower-elevation piping run),

and the total surge line volume, 22.3 f t'.

The outsurge and second-stage insurge 4-5 I

I

N L

source temperatures were modifited usir.g rudimentary heat balances to estimate the L

heat losses to ambient.

This type of correlation was necessitated by the absence of a temperature measurement near the nozzle. The performance of this correlation was checktd by comparing its predictions to the nozzle metal temperatures observed in Oconee, as wtll as by examining its response to test cases, The correlations of pressurizer nozzle fluid temperature provided estimates of g

I temperature versus time. These temperature were processed to obtain the extreme temperatures (peaks and vall3ys, or PP'

'he same menner as for the other l

surge line temperatures. The incremen' the nozzle fluid temperature were compared to the corresponding tir

.o obtain rates of change.

A change-weighting method was used to ot,'

'ange which were more n

om incremental time appropriate for stress analysis than thost steps.

I The result ing weighted rates of change tended to reflect those rates of change which were large, persistent, and significant for the stress analysis.

The surge line outside-wall temperature measurements, and therefore the correlations based on these measurements, reflected the thermal time constant of the wall.

Temperature changes of relatively short duration were separately identified for comparison with those predicted by the correlations.

The temperature changes during selected short duration events were combined with the general results of the temperature predictions for further consideration.

The temperature change associated with a short duration event was the difference between the pre-event temperature and the fluid temperature based ca the fluid volume displaced during the event.

The selection criteria for short duration events were as follows:

1.

Surge line mass flow rate greater than 10,000 lbm/h (insurge or out; urge).

2.

Duration (of flow rates greater than one-half of the maximum flow rate) less than 12 minutes.

3.

Temperature changes greater than 50f.

4-6

These criteria were selected such that all events of significance for stress or l

fatigue were considered, s

iJ IbttnLStr_iplna u

Thermal striping in the Davis-Besse Unit I surge line was evaluated using the same correlations and techniques as have been used for the lowered-loop plants, and as described in Section 4,3 of the main report, In the Davis-Besse calculations, h9 wever, the surge line fluid velocity was modified to account for the makeup system cycling experienced at Davis-Besse, lhls modified velocity has I

been used whenever the makeup system controls were placed in automatic, lhen the fluid velocity was required to be at least as large as the surge line velocities that had been observed (based on rates of change of pressurizer level) during the more extreme cyclic variations of pressurizer level.

This velocity was used to evaluate the Richardson number and hence the maximum striping amplitude; an increased velocity obtained a smaller Richardson number and a larger maximum striping amplitude.

4.4 Reylet_.0LODeralll90ALilhl9EJ A review of the operating history of the Davis-Besse Unit I plant was performed in a manner similar to that performed for the lowered-loop plants.

llistorical I

operating data were collected for plant heatup and cooldown events for Davis-Besse Unit 1.

These data were retrieved for 31 of 40 heatups and 32 of 39 cooldowns, ending with the heatup from the 6th refueling outage in June 1990.

The recorded parameters included hourly data for the pressurizer, RCS cold leg, and surge line (original thermocouple) temperature and pressurizer level.

The data were sut ficiently complete to characterize the limiting temperatures for RCS and pressurizer heatup nd cooldown for past operations with a high level of confidence.

This historical review provided substantial support for the selection of limiting temperatures for use in stress and fatigue analysis.

In combination with Davis-Besse data, data collected earlier for the analysis of the lowered-loop plant surge line were employed in describing the Davis-Besse heatup and cooldown transient.

Specifically, the lowered-loop plant data were 4-7

I used in definition of the typical flow rates in the surge line as derived from pressurizer level variations with time, In conjunction with the review of plcnt data, operating and surveillance procedures were reviewed for Davis-Besse Unit I to identify those events that might cause thermal stratification cycles.

Because of this review, certain differences between the sets of design transients for Davis-Besse Unit I and I

those for the lowered-loop plants resulted.

Transient events were added to include the conditions of (1) inservice makeup pump testing, and (2) complete interruption of pressurizer spray bypass flow.

Certain design events were deleted including HPl injection tests and miscellaneous pressurizer spray actuations. The HP! injection tests were eliminated as a design transient for the surge line for Davis-Besse Unit 1 because this test is conducted with the pressurizer at or near ambient temperatures.

It was determined that the miscellaneous spray actuations occur infrequently at the plant and that spray actuations are predominately ef the variety associated with operations to change I

the boron concentration in the pressurizer; as a result, the miscellaneous spray flow events are included in the separately described transient for pressurizer boron equilibration (Transient 2002).

The operating procedures for Davis-Besse Unit I were used as the basis for determining the major flow events accounted for in the heatup and cooldown design transients.

Random flow events, the particular causes of which have not been identified, were also included to ensure that typical pressurizer level changes are reflected in the design transients.

The data for pressurizer level versus I

time collected for the analysis of the lowered-loop surge line, in addition to data specific to Davis-Besse Unit 1, were used to characterize the random flow events included in the heatup and cooldown design transients for Davis-Besse Unit 1.

The design heatup and cooldown transient descriptions were generated based on plant data as well as plant limits for the relationship between pressurizer I

temperature and RCS temperature.

To conservatively generate design transient conditions for the surge line for past heatup events, two sets of temperature differential curves were used to ensure a conu rvative representation overall.

4-8 I

4 I

L One curve represents the maximum differential likely to be seen in any operation r

L experienced in the past by Davis-Besse Unit 1.

This maximum temperature difference (MID) curve is configured to be strongly conservative, and to cover

~

those few heatups not covered in the recorded hourly data. The second curve is configured to prcduce a more realistic, but still conservative representation of the data.

This curve bounds 80 percent of all the plant data.

I A certain fraction of the past events were specified to have occurred with conditions defined by the strongly cer

.tive MID curve. The remaining events were evaluated on the basis of the conservative representation of plant data provided by the second curve.

I lhe maximum temperature difference (MID) curve bounds all the plant data with the exception of three short operating periods (a few hours) which occurred during the actual heatups, lhe overall maximum temperature dif f erentials f or the heatup covered by the MlU curve were not exceeded, however, the dif ferences during these I

short periods exceeded a small portion of the bounding curve by 20T or less.

These short periods are negligible in the analysis, and the MID curve used as a basis for the development of design transients is a strongly conservative representation of past operations.

for generating descriptions of future design heatup artd cooldown events, the recommended operating limits for the pressurizer and RCS temperatures given in I

Section 8 of BAW-2127 were used with the added restriction that the maximum pressurizer temperature be limited to 41SF when the RCS temperature is below 185f.

The RC temperature of 185F was selected suf ficiently low to provide adequate flexibility for pressurizer operation; a more restrictive temperature of as high as 230f could be justified based on a review of plant data but would not provide the same degree of operating flexibility.

As a result of the review of plant data and operating procedures, the following major differences were identified between operations at Davis-Besse Unit I and those for the lowered-loop plants:

4-9

1.

At Davis-Besse Unit 1, the maximum pressurizer pressure and temperatures i

at which the plant is operated during low temperature conditions in the RCS are lower than for the other operating plants; this is because the Decay lieat Removal System relief valves are used for low temperature overpressure protection below about 250f in the RCS, which limits the I

maximum allowable pressurizer pressure.

This is consistent with the restriction that the maximum pressurizer temperature is limited to 415f when the reactor coolant temperature is below 185f as reflected in the analysis.

2.

The control of pressurizer level with the makeup valve operating under automatic control at Davis-Besse Unit I results in cycling of the makeup valve and small amplitude variations in pressurizer level on the order of

+/- 1 inch of level. This cycling is pronounced during plant operation at I

low pressure and substantially subsides during operation at high temperature and pressure (Modes 1, 2, and 3).

3.

The duration of heatup operations at Davis-Besse Unit I appears to be significantly longer than is typical for the other plants. Also, a large amount of time has been spent between consecutive cooldown and heatup events with the RCS at low temperature and the pressurizer maintained at an elevated temperature (up to 415'F).

The above differences in operations were factored into the descriptions of the design heatup and cooldown transients for Davis-Besse Unit 1.

4.5 Ogy_qlopment of Jeoisedj)1dgn Basisjnnliggi The revised surge line design basis transients are listed in lable 4-1.

These redefined transients comprise the bases for the reevaluation of the structural integrity of the surge line piping and nozzles. Table 4-2 presents the design number of events for each type of transient. The bases for the number of evants for design purposes are provided in the following subsections.

4-10

a The design basis plant heatup and cooldown transients were completely redefined r

L in this program, Other transients included in the design basis were generally retained in terms of the existing surge line boundary conditions of pressurizer

{

and RCS temperatures and surge line flow rates, but thermal stratification and striping were included in the surge line transient descriptions, in addition to the changes made to the design heatup and cooldown transients, a number of transients were added and other modifications made to the set of design basis I

events as a result of the review of the operating history and the operating procedures for Davis-Besse Unit 1.

In general, the development of the revis(d design basis transients for Davis-Besse Unit 1 followed the same process used for the lowered-loop plants.

The process used to generate the boundary conditions of terperature and the surge line flow rates for plant heatup and cooldown transients are discussed below, Plant data were used to characterize the variations in RCS and pressurizer g

B temperatures with time.

Using the plant data, tabulations of the durations of various modes of operation, i.e., time spent below 200f temperature, time spent in the actual heatup process, at temperature plateaus, and at hot zero power at the end of the heatup, were used to arrive at a description of the typical variation of RCS temperature with time for the design heatup and cooldown transients.

i Plant data were used to plot the paths taken during actual t.eatup and cooldown operations in terms of pressurizer-to-RCS temperature differential as a function of RCS temperature.

These traces established both bounding and typical variations of pressurizer-to-RCS temperature dif ferentials with RCS temperature.

For one of the design heatup t ransients, the bounding variation of thc pressurizer-to-RCS temperature ditferential was in sed upon the maximum temperature difference (MID) curve which generally bounds the plant data and is strongly conservative.

Given the RCS temperature at any point in time for the design transient, generated as described above, the variation of pressurizer-to-RCS temperature differential was used to establish the pressurizer temperature for tr.e heatup and cooldown transient based on either the generally bounding MID curve or the plant data.

4-11

1he pressurizer surge line flow rates were generated based on known, quantifiable L

plant operations, or other flow rates based on typical plant data for pressuriter level change versus time (for which the exact cause of the flow could not be identified).

A simulation of the makeup system control of the pressurizer level was used to generate the flow response in the surge line as a result of changes to the net I

make up volumetric flow rate (dif ference between in-flow to and out-flow from the RCS).

With the makeup controls in automatic, a change to the not ef fective makeup volumetric flow rate will perturb the pressurizer level, causing the makeup valve to be repositioned to restore pressurizer level to the setpoint.

The calculation of the surge rate includes the effects of the following:

Changes, or upsets, in the net makeup volumetric flow to the RCS; a

a Effects of RC volume change caused by the t i.ne rate of 1ange of temperature of the reactor coolant; a

Ef fects of the general trend in pressurizer level, e.g., ef f ects of change in pressurizer level setpoint by the operator; I

e Ef fects of spray rate, either automatic or manual spray operations (e.g.,

for cooling the pressurizer to depressurize the RCS or for adjustments of pressurizer baron concentration);

a Ef fects of operator adjustments of letdown to maintain both the level and the desired makeup flow rate.

I Based on a review of plant procedures, a list of the operations in the plants that potentially affect the surge line flow rate was generated.

To the extent I

possible, the likely numbers of flow changes anu the magnitudes of the changes of flow rate into (or out of) the RCS were estimated. Using these estimates for upsets of the net volumetric makeup into the RCS, and including additional

" random" events, the pressurizer surge line flow rates and pressurizer level response were calculated. A number of random events were added to the simulatinn to ensure that the total number of level changes agrees with the number of level change events derived from the plant data for heatup and cooldown operations.

The random flow rate event s were characterized by a statistical analysis of the plant data for pressur w level changes.

4-12

The surge line boundary conditions of pressurizer temperature and hot leg temperature, along with the surge line flow rates described for each design transient were used simultaneously to generate consistent sets of both

{

stratification and striping temperature differences.

The striping calculations utilize a correlation for the cumulative frequency of occurrence of striping temperature change as a function of striping amplitude.

The maximum striping amplitude was correlated to the Richardson number.

For a given design transient, the thermal response of the surge line was calculated as l

a function of time, and at each time interval (time cut) as the calculation advances, the distribution of the frequency of striping cycles was calculated for each degree f increment of striping temperature difference. The number of cycles for each increment of striping temperature differences was calculated by multiplying the f requency by the length of the time interval.

T ht. number of cycles in each increment of striping temperature difference was accumulated as the calculation progresses to arrive at a total for the entire transient, for each piping location and type of component (pipe or elbow), the thermal striping was evaluated considering a range of from 1/2 to 4 seconds for the period.

The critical period giving the maximum stress caused by the thermal striping was used for each particular location.

4.5.1 Heatuo Transilrttji 4.5.1.1 Heittup Transient DescriolhLrd and Number of Osmrfntn A number of different categories of design heatup transients have been described for Davis-Besse Unit 1.

The various categories arise from considerations of (1) time in plant life (past er future), (2) whether the pressurizer temperature represents a bounding uoper limit or a variation typical of plant operations, (3) the various time intervals corresponding to operations with different measured I

whip restraint clearances, and (4) accommodation of the period of limited I

clearance between the snubber stanchion and the west wall of the compartment.

The various time categories for the design heatup transients are defined as follows:

4-13

I I

Description Restraint Gaps Ca ry 11 For time period 1/77 through 5/80 Minimum original gaps T2 Time period 6/80 to 5/82 As measured gaps 6/80 Time period 6/82 to time of modification of snubber stanchion 13 As measured gaps 6/82 clearance to west wall, 12/84, and I

through 11/88 Time period 12/88 to time of modification of gap clearances As measured gaps 12~88 I

T4 made during the month of 4/90 and and 4/90 through the 7" fuel cycle, anticipated to end 9/91 Future events, beginning after Restraints gappec :o 15 Refueling Outace,No.7 allow free motion I

A set of two types of heatup events are specified for each time catep ry, one odd-numbered and one even-numbered design event type.

Generally, odc-numbered design heatup transients represent the generally bounding var W ion of pressurizer pressure with RCS temperature, e.g.,

lAl, lA3, lA5, etc.

Even-I numbered design transients generally represent variations of pressurizer temperature that are more typical of the available plant data.

Considered together, these transients provide a conservative representation for both historical and future operations.

For purposes of sitcplification, each of the design heatup transients is specified with the same heatup duration and sequence of major events.

Particular transients may differ in terms of the following parameters: (1) initial variation of RCS temperature prior to starting of RC pumps for plant heatup, (2)

I pressurizer temperature versus time, and (3) the number and timing of random flow events caused by unidentified operations in the plant.

Variations of RCS temperature at the beginning of the heatup are based on typical operating data for the plant. The pressurizer temperatures and pressures versus time that have been specified for each of the design heatup transients are based on either the maximum allowable RCS pressure limit or the available plant data.

4-14 I

Refer to figure 4-1.

The number of random flow events included in the design

(

heatup transient varies with the average lengths of time spent in the various phases of heatup, as derived from the available plant data for the particular

~

time category, Each heatup transient includes a phase at the beginning of the plant heatup, prior to starting the RC pumps, where the RCS temperature is maintained at approximately 100F and the pressurizer temperature is somewhat greater than 400F, The duration of this time period for the design heatups is set at about 40 hours4.62963e-4 days 0.0111 hours 6.613757e-5 weeks 1.522e-5 months .

To account for time spent in this operating condition in excess of 40 hours4.62963e-4 days 0.0111 hours 6.613757e-5 weeks 1.522e-5 months on a per heatup basis, a separate transient is described, Transient 101.

I The diration of the design heatup transient was selected to bound the majority of historical data for the heatup times. In some cases, this bounding value for the duration of heatup is significantly greater than the. average of the plant data for the set of heatup events corresponding to the particular time category.

g R

However, the numbers of random flow upsets are based on the appropriate average duration obtained from the plant data, not the bounding value of time specified for the design heatup.

A process was used in generating plant parameters for the design heatup transients similar to that used for the lowered-loop plants, described in the main body of this report.

Descriptions of the various specified design heatup I

types are presented below.

Transients lAl, lA2 Time category T1 for Transient 1Al, the RCS temperature at the beginning of the heatup is specified as 70F, heating to 100F in 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months prior to starting RC pumps for plant heatup; the pressurizer temperature is bar,ed on (1) the upper bound of plant data for RCS temperatures from ambient to 280F and (2) a generally bounding RC pressure curve corresponding to the maximum temperature differential (MID) curve for RCS temperatures from 280F to hot, zero power conditions.

4-15

for Transicat lA2, the RCS temperature is specified to L

start and be maintained at iOOr prior to running RC pumps. Ihn pressurizer temperature is based on (1) the I

upper bound of plant data for RCS temperatures from ambient to 280F and (2) the typical variation with RC temperature as obtained from plant data for RCS temperatures from 280F to hot, zero power conditions.

I for Transient

1A3, the RCS Transients lA3, lA4 Time category 12 temperature is specified to start and be maintained at 100f prior to running RC pumps.

The pressurizer temperature is based on (1) the upper bound of plant l

data for RCS temperatures from ambient to 280F and (2) a generally bounding RC pressure curve corresponding to the maximum temperature differential (Hip) curve for RCS temperatures from 280f to hot, zero power conditions, for Transient lA4. the RCS temperature is specified to start and be maintained at 100f prior to running RC pumps. The pressurizer temperature is based on (1) the upper bound of plant data for RCS temperatures from ambient to 280F and (2) the typical variation with RC temperature as obtained from plant data for RCS temperatures from 280f to hot, zero power conditions.

Transients lAS, lA6 Time category T3 -- for Transients lA5 and 1A6, same as Transients lA3 and 1A4 except the pressurizer conditions are based or, plant data for the time spanning category T3.

For RCS temperatures above 280r, the pressurizer conditions for Transient lAS are based upon the maximum temperature differential (MID) curve.

Transients lA7, lA8 Time category T4 -- plant data are available describing the variatP ' of pressurizer pressure conditions for all heatup transients in this time category; consequently, 4-16

i I i

the design transients are limited to one representative type of transient, LAB.

(There was no need to specify a

bounding curve for conditions of pressurizer temperature at RCS temperatures above

280F, corresponding to the odd-numbered, i.e., lA7, transient type.) The pressurizer temperature is based on (1) the upper bound of plant data for RCS temperatures f rom I

ambient to 280F and (2) the typical variation with RC temperature as obtained from pl ant data for RCS temperatures from 280F to hot, zero power conditions.

Transients lA9, ;A10 Future heatup events category -- for Transient lA9, the RCS temperature at the beginning of the heatup is specified as 70F, heating to 100F in 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months prior to running RC pumps for plant heatup.

The pressurizer temperature is based on the upper bound of the plant I

data for RCS temperatures from ambient to approximately 185F and the recommended operating pressure-temperature limits specified in Section 8 of BAW-2127 for RCS temperatures above 185F (refer to item 1, page 4-9).

l The pressurizer temperature has been specified at a high value (sufficient to operate RC pumps) early in the heatup; this reflects the conditions experienced frequently in the plant where the RC pressure is raised to supply the required NPSli for extended periods prior I

to actually starting the pumps.

For Transient lA10, the initial RCS temperature of 100F is maintained until running RC pumps for plant heatup.

The pressurizer temperature is based on the recommended operating conditions for heatup.

The pressurizer temperature is specified to increase to the value required to operate RC pumps just prior to actual running of the pumps to begin RCS venting operations and plant heatup.

4-17 I

I

s The recoaxnended operating conditions for allowable pressurizer temperature L

variation with RC temperature for future heatup operations, Transients IA9 and 1 A10, are equivalent to the recommendations made for plant heatup of the lowered-

[^

loop plants given in Section 8 of ilAW-2127 except for the added restriction that L

the maximum pressurizer temperature is limited to 415F when the RCS temperature

~

is below 185f, I

the total numbers of occurrences of heatup events f or a given time category are based on a tabulation of historical heatup events supplied by Toledo fdison. In each of the time categories for the heatup transients, the numbers of transient events are distributed with 15 percent of the events for the type based on the strongly conservative MID curve (odd-numbered design events) and the remaining 85 percent for the type with the more typical variation of pressurizer temperature (even-numbered design events).

The total number of events for time category il from the tabulation of heatup events supplied by Toledo Edison has been increased to account for the two hot f unctional tests conducted during the I

initial operations for plant startup.

Based on the durations of these hot functional test operations, an equivalent of seven additional heatup events has been included in the design numbers of events specified for this time category.

Lithl_MMimuPLlrC55gr]Itr:10-RC$_.l0PDer31tu rLDiff tEt0Ct A maximum stratification temperature dif ferential was specified for three destyn transient events to account for operating conditions during which the system pressure and the pressurlier temperatu e may have reached the maximum llowed by either (1) the RCS pressure corresponding to operation at the relief setting of the Decay lleat Removal System reitef valves at RCS temperatures below 280f, or (2) the strongly conservative MID curve at temperatures above 21101 these I

operating pressures were specified for Transient lAl and the resulting maximum thermal stratification in the surge line for this transient is 358f.

LhbL.J9t! nkfl.10nd Ll i o n Ll RE._ lemile rA t tiru lyt riy 3_LLme lhe variation of RCS temperature with time for the revised design heatup transients is based on plant data.

The available plant data for heatup events was used to arrive at average time durations for the various phases of heatup me I,

I

I such as (1) operations with the RCS temperature below 200f, (2) operations when RCS temperatures are increasing, (3) intermediate temperature plateaus, and (4) at hot, zero power conditions prior to power escalation.

For purposes of simplification, the durations of each of the design heatup transients are

'g identical even though the average duration of the historical heatup events differ for the individual time categories. The total duration and the durations of each I

of the different phases of the design heatup were selected to bound the average values obtained from plant data for the heatup events in each time category.

The l

total duration specified for each of the design heatup transients is seven days, or 168 hours0.00194 days 0.0467 hours 2.777778e-4 weeks 6.3924e-5 months .

The transient is described for the operations ranging from cold conditions to 8 percent power, consistent with the range of conditions specified in the original design heatup transient for the plant.

The traces of pressurizer temperature versus time shown in figure 4-1 for the boundary condition on the surge line at the pressurizer are based on one of the g

following: (1) the MID curve for the RCS (for lAl, lA3, and 1AS), (2) plant data 2

(remainder of historical heatup transients), or (3) recommended limits for heatup operations (future transients).

The pressurizer tempe rature versus RC temperature relationships for each of the various design transients for plant heatup are described in Section 4,5.1.1.

4,5,1.4 Sutgt Line fIow Rain _19LD_e119n_lleal_up_Irimsjin11 Changes of flow rate in the pressurizer surge line piping can lead to thermal stratification transients. It is not possit le to describe every plant event that influences the surge line flow rate and affects the thermal transients for the surge line piping and nozzles. Ilowever, by quantifying the major influences on the flow rate and supplementing these with random flow events, design heatup transients can be generated which are conservative representations of the actual E

plant transients in terms of the number and magnitude of surge line flow events.

Each heatup transient type (Transient lAl through 1A10) is specified with the same basic set of quantifiable flow rate events; however, the number and timing of the added random flcw events may vary for some of the transients.

The number and timing of the added random flow events varies for the different time 4-19 rW'"

~

!I categories of transients based on historical differences in the average times required to complete the plant heatup.

I Typical plant heatup operations that may af fect the net makeup flow to the RCS are listed in Table 4-3.

Those operations judged to be significant and quantifiable are the major events taken into account in the descriptions of the design transients. These events include RCS temperature changes, RC pump starts, I

certain surveillance tests, and RCS venting operations.

The response of the makeup flow rate controls to each of these events is accounted for in the development of the transients.

The random flow events incorporated into the design transients are based on measured pressurizer level data for both the lowered-loop plants and the Davis-Besse plant.

The available plant data was statistically analyzed to characterize the random flow events in terms of the magnitudes of flow rates and pressurizer level changes. Descriptions of the random flow events are based on I

the following:

1.

Plant heatup data were evaluated to determine the mean and standard I

deviation parameters for flow rates and pressurizer level changes during plant operations over various ranges of RCS temperatures.

The average numbers of flow events were determined for heatup operations for the various ranges of RCS temperatures.

l 2.

The average numbers of flow events described for the design heatup transients were determined so that the numbers of events per unit of operating time specified for the design events are consistent with the plant data.

I 3.

The numbers of random flow events included in the design events were set so that the sum of the defined events and the random events are equivalent to the average of the total number of events per heatup transient as obtained from the plant data.

I 4-20 I

4.

The flow change data were treated as a normally distributed, random variable and divided into three representative ranges of magnitude of equal probability based on the normal distribution curse.

Flow rates bounding these three ranges of flows are used to describe the random flow events and the three bounding flow rate events are specified to occur in a recurring sequence during the heatup transient, 5.

The random events were spaced uniformly over the appropriate times corresponding to the spe'efied operating RCS temperature ranges.

[

The automatic operation of the makeup valve controls at the Davis-Besse plant has

_3 resulted in cyclic stroking of the control valve and small amplitude variations in pressuri2er level.

The cycling is most prominent when reactor coolant pressure is low and the makeup control valve must control flow at low rates

.=

-E across a large differential pressure. Under these cenditions, the makeup valve operations are characterized by opening of the valve and adding flow to the RCS qj at a high rate for a period of about I minute followed by closing of the valve with a minimum flow for a period of about three minutes.

During the time the

- valve is open, a surge flow into the pressurizer exists and upon closing, an outuurge takes place.

These oscillations are not directly described as a

.;'nponcnt of the surge line flow rate since the effect on the global stratification in the surge line is small.

However, the surge line flow rate into and out of the pressurizer does affect thermal striping in tha line. The effects of the makeup cycling on thermal striping are taken into account in the thermal stratification and stripir.g calculations.

The design transients are conservatively specified to inc'lude this makeup cycling for all past historical heatup events, future heatup events are also specified to include makeup cycling and the associated thermal striping, for the design heatup transients, the pressurizer spray, when active, is considered to be actuated in one of three modes, (1) manual actuation for purpotas of pressurizer boron equilibration, (2) automatic spray operation, and g

J (3) minimum continuous bypass spray. To describe the effects of large spray flow I

rates, i.e., ca".es of boron equilibration in the pressurizer or automatic actuation of spray, the spray flow rate component of the surge line flow rate is 4

4-21 1

specified.

When the main spray valve is closed the minimum continuous bypass spray flow rate is used. This flow rate may range from 1.5 to 5 gpm with all RC pumps operating.

Since there is a significant uncertainty in determining the magnitude of the minimum continuous spray flow rate, the st'atification correlation model instead uses the number of operating RC pumps to determine the thermal response of the surge line; the bypass spray rate is not explicitly specified.

The Davis-hsse pressurizer spr y valve is adjusted to pass a maximum flow rate of approximately 190 gpm fcr control of pressure transients which potentially might occur during plant power operation.

The valve is occasionally throttled open during normal operations to adjust the pressurizer boron concentration, for

~

the design heatup transients, pressurizer spray actuations are sper fled for the

~

]

manual operations to adjust the boron concentration at cold conditions and at hot J

conditions.

A number of automatic actuations are specified to account for potential actuations of the spray during power escalation to 8 percent power at the end of the heatup transient.

]

4 1 2 _ 1qqldgyn Tran11cn h 4.5.2.1 C1qldgwLlraps!ent_ Ducrlplions _ and thtattqr_qLQrgurrences Similar to the design heatup transients, the cooldown events were described for five different categories of operating timec.

Categories Il through T4 are specified for historical events, and category 15 for future events.

Refer to Section 4.5.1 for a description of each of the time categories, The design cooldown transient describes the plant operations and the thermal response of the surge line during the power reduction from 8% power to hot, zero power and then plant couldown to refueling temperature, approximately 140F hot leg temperature.

=

The RCS temperatures and pressurizer temperatures for the design cooldown transients are shown in figure 4-2.

The descriptions of the design cooldown transients are essentially identical except 'or the pressurizer temperature versus time and the associated spray flow ra*.es required for spraydown of the 3

pressurizer. For purposes of simplification,.;ther parameters such as duration 4-22 1

i of cooldown, RCS temperature versus time, and sequence of events are essentially s

the same for the various cooldown transients.

Each time category has associated with it a set of two cooldown events, with these cooldown events dif fering coly la the temperature versus time trace for the pressurizar temperature.

The set of r tant transients for time category 11 consists of design cooldown transients 181 and 182, The two types of transients

{

in each tine category, i.e.,

odd-numbered and even-numbered transients, are f

spet.lfied to describe (1) a strcagly conservative envelope based on the generally bounding MID curve and (2) a temperature trace that is typical of the plant data.

Overall, approximately 15 percent of the transients specified for a time category are assigned to the transient type oith the bounding trace of pressurizer temperature (odd-numbered events) and the remaining 85 percent assigned to the transient described with the typical trace for pressurizer temperature (even-numbered events).

The design cooldown transients for Davis-Besse Unit I are described below.

Transients 181, 102 Time category T1 -- for Iransient 181, the pressurizer j

temperature is based on (1) a generally bounding RC pressure curve corresponding to the maximum temperature dif ferential (HTO) curve for RCS temperatures from hot, zero power conditions down to about 280f and (2) upper bound of plant data for RCS temperatures frov <80F and below with the Decay Heat Removal System operating.

for Transient 182, as discussed in Section 4.5.1.3, the pressurizer temperature conservatively represents the 1

available plant data for RCS temperatures above 280F and bounds all of the plant data for conditions with the Decay Heat Removal System in operation below about 280f Transients 103, 184 Time category 12 -- for Transient IB3, the pretsurizer temperature conservatively represents the plant data at RCS temperatures above 280f and boands all thc available 4-23

I plant data at temperatures below 2B0f with the Decay Heat Removal System opertting.

I for Transient 184, the pressurizer temperature is based on a

conservative representation of pressurizrir temperature from the plant data for RCS temperatures above 280f and an upper bound of all the plant data I

below 280f.

Transients 1B5, 106 Time category T3 -- for Transients 185 and 106, the bases for the pressurizer temperature variations with RCS temperature are defined to be identical to that for Transients 183 and 184, respectively.

The bounding envelopes selected for Transients 183 and 184 also bound the plant data for time category T3.

Transient 1R8 ilme category 14 - for Transient 108, the pressurizer

'emperature bases are defined to be identical to that for Transient 104, which are bounding for the plant data of time category T4. (No transient is described for 107 for the same reasons as discussed previously in Section 4.5.1 for the design hea+up events for time category T4.)

Transients 109, 1810 Time category.) -- for Transient 189, the pressurizer tenperature is based on the recommended guidelines given

~

in Section 8 of the main report below about 280f with the additional restriction that the pressurizer temperature is less than 415f when the RCS temperature l

is below 185f.

At RCS temperatures above 280F, the specified pressurizer temperature bounds the available plant data for the entire operating history of the I

plant.

4-24 I

k P

L for Transient IB10, the pressurizer temperature is based

[

on the recommended guidelines given in Section 8 of the main report with the additional restriction that the

(

pressurizer temperature is less than 4 MF when the RCS temperature is below 185F.

LLL2 Bqundary Ttuttal_ur1LtLL[ unction of Tie Based on the available plant data, the time specified in the design cooldown transients for cooling the plant and filling and depressurizing the pressurizer is 72 hours8.333333e-4 days 0.02 hours 1.190476e-4 weeks 2.7396e-5 months .

The time durations for each portion of the design cooldown transient have been estimated using the available plant data for the following phases of the cooldown operations:

power decrease from 8 percent power to hot, zero power, Cooldown operations (with average RCS temperature decreasing),

RCS temperature plateaus during cooldown, RCS temperature below 200F with pressurtier hot, and Pressurizer fill and spraydown at end of plant cooldown.

I Similar to the original design transient for cooldown of the plant, the RCS temperatures versus time for the design cooldown transients are defined over the iange of operations from an initial power level of 8 percent to the point where the hot leg temperature reaches the ref ueling temperature of 140f. The duration I

of the typical design transient wt.s lengthened to be more representative of the actual plant operations.

Temperature plateaus were added and cooldown rates adjusted to give reasonable agreement between the design transients and the available data. The pressurizer temperature versus time plots for the design cooldown transients were developea based on the relationship of nressurizer temperature to RCS temperature based on the plant data.

I Based on the 4.vailable plant data, the plant cooldown frequently is t<:rminated without completely depressurizing and cooling down the pressurizer. for purposes j

J of determining typical values for the total duration of cooldoun operations from the historical data, the cooldown was considered to end at a time corresponding to about 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months after the RCS temperature decreases below 200F, The excess 4-25 m

L operating time not included in the plant cooldown with the pressurizer hot and

(

the RCS at a low temperature is it.cluded in a separate design event, Iransient 101.

For those leistorical cooldown transients where the pressurizer is cooled

{

to near ambient, the plant cooldown was considered to end when the temperature difference between the pressuriter and the RCS hot leg decreases below about 50f.

(

$12. 3_...htge L ine flpx EtttLfE_De1.Lq[LfDDldOMLhAR11CD11 The flow events for the design cooldown transients were developed in a manner similar to the methods used to describe the flow events for the design heatup transients as discussed in Section 4.5.1.4.

Operations that were judged to be significant and quantifiable are considered the

(

major events to be accounted for in the cooldoc transients.

The makeup flow response to each flow event is accounted for in the development of the transients.

Surge line flow rates include the effects of operations to spray

{

down the pressurizer for either boron concentration adjustments or cooling and Jepressurizing the system.

Random type of flow events were added to ensure that the total number of flow events f or the design transients properly represent the historical operating experience.

Available plant data for both Davis-Besse Unit I and the lowered-loop plants were statistically analyzed to describe the flow events included in the design transients.

The method used to characterize the random flow events for the design heatup transients, outlined in Section 4.5.1.4, was used to describe the flow events for the der.ign cooldown.

411 Other_Dn ign Tranniten1; Plant parameters for the original design transients previously described for Davis-Besse Unit I for operations at hot conditions were generally retained, however, the surge 11oe conditions for these transients were revised to include the effects of thermal stratification and thermal striping.

Existing descriptions of RCS parameters of temperatures, pressures, and spray flow rates remain unchanged for most of these original design basis transients. The set of these transients originally described for operations at hot conditions was 4-26

revised and expanded somewhat to reflect the results of the review of the plant operating history and procedures. Also, in some cases, changes were made in the numbers of events f or the design transients to more appropriately reflect the

(

types end frequencies of certain operations in the plant.

Modifications and additions to the original set of design transients for operations at hot conditions are discussed below.

Unlike the set of design transients for the lowered-loop plants, the set of design transients for Davis-Besse Unit I does not include the effects of testing HPl safety injection or HPl suction check valve tests.

These tests do not produce any significant thermal

(

transients on the surge line piping and nozzles because the tests are conducted at very low pressures (i.e., pressurizer temperature near ambient), with the

~

reactor vessel head removed.

I 4AL 1 OngnlignLAt Co1A. premtdreL(pndilian$ - IBulichLK1 The historical records for Davis-Besse indicate a substantial amount nf operating time has accrued under conditions with the RCS average temperature at cold conditions, i.e.,

approximately 100f, and the pressurizer temperature at I

approximately 400F. As discussed previously, this type of operation has occurred under conditions where the plant was maintained with an elevated pressurizer temperature either between successive cooldown and heatup events or between initial pressurization and the time of actual RCS heating in the early phases of plant heatup operations.

10 properly describe these operating conditions in ter.ns of the thermal effects on the surge line piping and nozzles, actual plant data for pressurizer level versus time was used to establish a representative history for the flow rate variation with time. Typical data for these operations over a period of 10 days were used to characterize the surge line conditions.

The particular data used for describing the design trar.sient were taken f rom measurements for the month of November, 1988 (12" through the 22'"').

4_d,3,2 SieaiSigttlemnenLun Jariatio_ps - 1ransienLil The design transient for describing steady state operations at power was redefined.

In order to properly reflect the operations under these conditions which involve surge line flow and temperature oscillations caused by normal 4-27

control variations in averane RC tempeteteras and makeup valve cycling, a representative set of plant measurements of pressurizer level versus time was 5

used to characterize the surge line variations of flow rate. The plant data were recorded during the time that the plant variables were being monitored to determine the surge line temperature stratification parameters.

A representative period of eight hours of operation of the plant near fell power

(

was selected f or this design transient description.

The operat ion includes ef fects of the cycling of the makeup valve.

The number of design events specified for stress and f atigue evaluation of the surge line piping and nozzles for this design transient corresponds to the total possible number of eight-hour

(

operating intervals at power over the 40-year service life of the plant based on a plant capacity factor of 0.8.

LLLl__Prnut!rker4SJJor91LLquilibration - 1miskt1L20D2 Transient 200 was added to the set of design transients for tht surge line to describe the ef fects of spray and heater operations to equalize the pressurizer and RCS boron concentrations. The operation involves use of spray ilow through the pressurizer to cause the boron concentration to approach that in the RCS.

I A modulated spray flow rate of about 50 gpm was used in the description of the spray transient.

This heater and spray operation is normally performed approximately twice a week f or a period of about eight hours for each operation.

The number of baron equilibration events specified allows for these operations biweekly over the 40 year service life of the plant, with a'i additional number of design events included to allow f or other miscellaneous, undefined, spray actuations occurring in the plant.

1911_lrt(c rr u nli g uLS p r aL Ll pt_1_l tnish nll0[

1 4AL4 A minimum spray line flow rate is normally maintained in the plant when the g

pressurizer to RCS temperature dif f erence is greater than about 250f, to minimite P

the ef fects of thermal transients on the pressuriter spray nozzle.

Although inf requent, this flow rate is reduced to zero occasionally in the plant if the spray isolation valve is closed, lhe surge line temperatures are af fected by the change in the flow rate and the surge line temperatures are normally decreased somewhat under conditions of zero bypass flow until the flow is restored, causing 4-28

E the surge line to be subjected to one thermal cycle for each interruptian in the minimum flow. The number of events for design purposes is specified as 20, based on past experience indicating a frequehty of one interruption every two years in F

the plant.

L 1.5.3.5 Inservitt.Mdeyp Pumplest - Transient 22[

1he Quarterly inservice Makeup pump Test involves the starting and running of a standby makeup pump for a short period of tiine.

Upon starting the pump, the additional makeup flow rate causes an insurge into the pressurizer.

This I

insurge, and the following action to restore the pressurizer level to the desired setpoint produces a surge line thermal transient. The number of events of this type specified for design purposes is 160.

l iJ Duion ltansients Summan for the purposes of design analysis of the Davis-Besse Unit I surge line, the design transients were redefined to incorporate the effects of thermal stratification and striping. The design plant heatup and cooldown transients for the surge line were completely redefined. Certain other design transients were added or revised to more accurately reflect the actual operations in the plant.

Table 4-2 lists the design transients and the number of events of each type of I

transient for analysis purposes.

Calculations of the stratification and striping thermal cycles were performed for each type of design transient.

These numbers of thermal cycles for each event and the number of design transients of each type as given in lable 4-2 determine the total numbers of thermal cycles considered in the revised design analysis of the surge line.

Table 4-4 provides a brief summary of some of the important results of the stratification and striping calculations including the maximum stratification temperature reversal, the distribution of cumulative numbers of stratification temperature reversals, and the maximum striping amplitude.

Results shown are for the lower horizontal section of piping which generally esperiences the maximum thermal stratification magnitude in the piping.

4-29

Two columns are shown for the maximum significant striping delta-T.

The pst

(

valtes pertain to the existing surge line configuration.

Toledo Edison has committed to making modifications to the surge line supports configuration and thermal insulation.

These changes are expected to result in stratification temperature differences in the surge line comparable to the Oconee measurements.

On this basis, for future transients, the maximum significent striping delta-T results calculated for the lowered-loop surge line design transients for I

operations at high temperature (i.e., power operations) have been assumed for the Davis-Basse plant.

Monitoring of the Davis-Besse Unit I surge line will be

(

necessary following the changes to confirm the existence of stratification delta-T magnitudes similar to those measured at Oconee.

I l

l l

1 I

I I

I 4-30 l

lable 4-1.

Surge t.ine Design Basis Transient List L

Transient Transient Description Hodification from Original 10 Transients (0DB - Original

_ _ _.._ Des _1gnjlasis L ___ _ _ _,

lAl Time category 11 - RC Defined to realistically Temperature of 70F to 8%

represent the most severe full power (FP), a general-heatup from the available plant l

ly bounding pressure curve data for time il specifies P/T relationship 1A2 Time il - Trc of 100f to Defined to realistically I

8XfP, RC pressures specify represent heatups cccurring P/T relationship - bounding with RC pressures higher than below 280f RCS and typical typical values over the range l

of plant data above 280F of RC temperatures lA3 Time category 12 - 1rc of Defined to realistically 100f to 8xf P, RC presst.res represent the most severe 1

specify P/T relationship -

heatup from the available plant bounding colow 280F RCS and data for time Tl conservatively representa~

tive of plant data above 280f lA4 Time 12 - Trc of 100f to Defined to represent the 8XFP, RC pressures specify typical heatup events for time P/T relationship - bounding 12 below 280f RCS and typical of plant data above 280F 1A5 Time category 13, otherwise same as lA3 1A6 Time 13, otherwise same as lA4 1A7 (No transient defined) 1A8 Time T4, otherwise same as lA4 1A9 Time category 15 - Trc of defined to conservatively 70F to 8%fP, RC pressures represent future heatups specify P/T relationship -

bounds plant data below an RCS temperature of about 185F and recommended limits above 185F 4-31 l

Table 4-1, surge Line Design Basis Transient List (cont.)

!L Transient Transient Description Modification from Original ID 1ransients (000 - Original DeSi9nBasis),

lA10 Time 15 - 1rc of 100F to Defined to represent typical I

8tiP. ItC pressures specify future heatups P/T relationship - bounds plant data below an 110S temperature of about 185f I

and recommended limits above 185f i

1B1 lime category 11 - 8%f P to Defined to represent a bounding refueling temperature, P/I cooldown based on least relationship based on a limiting Appendix G limits g

generally bounding pressure (highest pressures) and plant R

curve above 280f and upper data for time 11 bound of plant data below 280f 182 Time 11 8%fP to refueling Defined to conservatively bound temperature, P/T most cooldown events for time relationship based on a 11 I

conservative representation of plant data above ?801 and bounds all plant data l

below 280f 183

~ lime category T2 - 81FP to Defined to conservatively bound refueling temperature, P/T cooldown events for time 12 I

relationship based on a conservative representatton of plant Jata above 280f I

and bounds all plant data below 280f IB4 Time T2 - 8xfP to refueling Defined to represent typical temperature, P/T cooldown events for time 12 relationship based on typical plant data above 280f and bound of all plant data below 280f 185 Time category 13, otherwise same as 1B3 IBS Time 13, otherwise same as 1B4 4-32

Table 4-1.

Surge Line Design Basis iransient list (cont.)

lu Transient Transient Description Modification from Original 10 Transients (ODB - Original Design Basis)

IB7 (No transient specified) l 188 Time 14, otherwise same as 184 189 Time category 15 - 8xf P to Defined to conservatively bound refueling temperature, P/1 most cooldown events for time relationship based on bound 15 I

of all plant data above 280f and recommended operating limits below 280F l

1810 Time 15 - 8xfP to refueling Dcfined to represent typical temperature, P/l cooldown events for time 15 relationship based on I

recommended operating 1imits 1C1 Cold RCS, Pressurized Defined to represent operatlons i

Operations with the RCS at 100f and the pressurizer at 415f 2A Power Change from 01 to 151 Surge line temperatures based I

FP on ODB boundary conditions.

2B Power Change from 154 to 0X Surge line temperatures based FP on ODB boundary conditior.s.

3 Power loading 8X to 100X, FP Surge Line temperatures based on 000 boundary conditions.

Power unloading 100 to 8 Surge Line temperetures based l4 percent on 008 boundary conditions.

5 Ten percent step load power Surge Line temperatures based increase on ODB bounday conditions.

6 Ten percent step load power Surge Line temperatures based decrease on 000 boundary conditions.

7 Step Load decrease 100 to Surge line temperatures based 8X FP on ODB boundary conditions.

4-33

Table 4-1.

Surge Line Design Basis Transient List (cont.)

i h

Transient Transient Description Modification from Original i

10 Trans'ents (ODB - Original

__._...De_s.ign Basis L _____ _ _ _,_ _.

8 Reactor Trip All trips now included under I

the category of type 8A, 88, or 8C transients.

Previously, certain other trips were specified separately.

9 Rapid Depressurization Surge line temps based on ODB boundary conditions.

10 Change of RC Flow Rate Surge Line temps based on 00B boundary conditions.

13 Steady State Temperature Surge Line temps based on Variations typical plant data for RCS temperature variations and makeup valve cycling.

14 Control Rod Drop Surge line temps based on ODB boundary conditions.

19 feed and Bleed Operations Surge Line temps based on 000 boundary conditions.

l 20 Miscellaneous Transients A new transient was incorporated to describe pressurizer spray and heater operations used to equilibrate I

pressurizer & RCS boron l

concentrations.

Transient 208, previously described as a al miscellaneous spray actuation event, was deleted. An additional transient was I

included to describe the complete interruption of spray flow.

22 Test Transients A transient was added for the Inservice Makeup, Pump lest.

4-34

Table 4-2.

Design Transients - Numbers of Events b

Description Event Number lotal Type Events Heat from cold conditions to 8X full power lAl (TI) 3 1A2 (T1) 17 1A3 (T2) 1 1A4 (12) 5 1A5 (T3) 2 1A6 (T3) 12 lA7 (T4) lA8 (14) 5 1A9 (TS) 32 1A10 (15) 163 240 Cooldown from 8% full power 181 (T1) 3 IB2 (TI) 17 1B3 (12) 1 s

184 (12) 5 185 (13) 2 186 (T3) 12 1B7 (14) 188 (T4) 5 IB9 (TS) 32 1810 (TS) 163 240 RCS Cold, Pressurized Operations 101 (TI) 9 101 (12) 1 1C1 (T3) 4 101 (14) 1 101 (TS) 35 50 Power change 0% to 15%

2A 1440 Power change 15% to Ox 2B 1440 Power gpding 83 to 100%

3 1800 4-35

L Table 4-2.

Design Transients - Nun.ber of Events (cont.)

l L

Description Event Number Total Type Events Power unloading 100% to 8%

4 1800 Step load increase of 10%

5 8000 Step load decrease _of 10%

6 8000 Step load reduction 100% to 8%

7 310 Reactor trip 8A 80 8B 232 BC 88 400 Rapid depressurization 9

40 Change of RC flow rate 10 20 Steady-state temperature variation 13 35000 _

Control rod drop 14 40 Hakeup and letdown feed and bleed operations 19 40000 Miscellaneous makeup flow change 20A 30000 208 20C 4.0E6 Misc. - Pressurizer heater / spray operation 2002 6000 Complete interruption of spray flow 20E 20 Quarter _ly_inservicemakeul pump test 22E 160 l

I I

4-36

Table 4-3.

Events Affecting Surge Line flow I

for Plant Hettup and Couldown L

Plant Heatyp High pressure injection check valve testing (at reactor coolant pressure <50 psig) forming steam pocket in pressurizer at 50 psig

~

Pressurizer pressurization (af fects letdown rate and reactor coolant system volume and mass)

Purging of pressurizer nitrogen through vent lines Adjusting pressurizer level setpoint Controlling pressurizer level in auto (w/ valve and controller deadbands)

Drawing of pressurizer chemistry sample Testing of PORV (at 100 psig or 200 psig)

Adjusting pressurizer level with LD flow Placing makeup and purification in service (start makeup pump at <l50 psig)

Pressurizing to 150 < reactor coolant pretsure < 175 psig I

Placing reactor coolant pump seal return in service (makeup and purification system)

Adjusting of reactor coolant pump seal injection / return flows Venting reactor coolant system RCS heating without reactor coolant pump operating Venting reactor coolant pumps for initial operation Running each pump for 5 minutes (initial run after filling of reactor coolant system)

Re-venting reactor coolant system (control rod drive mechanism and high point vents), re-venting reactor coolant pump Drawing steam generator vacuum by opening turbine bypass valves Starting 2 reactor coolant pumps (in same loop) to commence heatup w/ pump power Pressurizing steam lines by closing turbine bypass valves (affects HV rate) 4-37

T able 4-3.

Events Affecting Surge line flow for Plant Heatup and Cooldown (cont.)

Holding reactor coolant system temperature (e.g., at 250f for reactor coolant system chemistry in spec)

Isolating low pressure injection at E80f (affects heatup rate)

Opening spray line block valve (when reactor coolant system temperature reaches 200f)

Closing letdown orifice manual bypass Adjusting makeup flow rate with increasing reactor coolant system pressure Starting third reactor coolant pump Performing steam generator fill, soak, drain operations (300 - 400f)

Pressurizing reactor coolant system af ter steam generator fill, soak, drain operations Adjusting makeup bypass flow rate (at reactor coolant pretsure >$00 psig, 1000 psig, and 1500 psig)

Holding reactor coolant system temperature for reactor trip and reactor protection system reset (at 1700 to 1725 psig)

Pressurizer spray controlling in auto at 2155 psig, heaters on for boron equilibration Starting 4th reactor coolant pump at 480f Cycling of turbine bypass valves (in manual) every 20 minutes - temperature holds I

at >500F Turbine bypass valves controlling steam pressure at 870 psig in auto Adjusting boron concentration in reactor coolant system (changing makeup /LD)

Physics testing at hot zero power Surveillance testing ElitnLLoslstown Degassing pressurizer and reactor coolant system (vent pressurizer to waste gas header) 4-38

Table 4-3.

Events Affecting Surge Line flow

[

for Plant lleatup and Cooldown (cont.)

}

Decreasing pressurizer level, 28% to 0% full power (operator adjusts setpoint)

Decreasing reactor power demand in manual at <0.5%/ min and reducing average temperature Reducing turbine load demand (manually) w/ auto opening of turbine bypass valves to hold pressure reducing turtine load to <20 MW and tripping turbine I

Adjusting pressurizer level to 85 inches at 532f average temp Controlling pressurizer level in auto (w/ valve and controller deadbands)

Sampling boron concentration in reactor coolant system Adjusting boron in reactor coolant system (changing makeup /LD flow rates)

Tripping reactor coolant pump to go to 1/2 operating status Raising steam generator levels for hot soak (at 532f and at 400f to 300f)

Adjusting turbine bypass valve positions for desired cooling rate (turbine bypass valves in manual)

Spraying down pressurizer I

ilolding reactor coolant system temperature for placing reactor protection system in shutdown bypass Tripping reactor coolant pump to go to 0/2 operating pump status a

Performing core flood tank valve tests (at reactor coolant pressure 750 to 700 g

psig)

Performing power-operated relief valve cycle tests (at reactor coolant pressure 725 to 675 psig)

Decreasing pressurizer level setpoint to 60 inches (prior to tripping reactor coolant pumps)

Opening letdown orifice manual bypass Holding for steam generator chemistry, continuing fill, soak, and drain operations Holding reactor coolant system temperature at 280F (adjusting turbine bypass valve positions) 4-39

Table 4-3.

Events Affecting Surge Line Flow for Plant Heatup and Cooldown (cont.)

Valving in the low-pressure injection system

)

Adjusting low-pressure injection cooler outlet temperature to cold leg temperature Raising steam generator level for natural circulation cooldown Tripping last two reactor coolant pumps (average temperature increases) t Establishing 25"/hr increase in pressurizer level Raising and lowering pressurizer level when on decay heat removal 3

Open decay heat removal Aux spray valve and set to ensure surge line net flow l3 from reactor coolant system to pressurizer l

Spraying down to final pressurizer pressure filling steam generators to wet layup level (at less than 200F)

Securicq reactor coolant pump seal injection / return flows Shutting down makeup and purification system (at less than 75 psig)

I I

4-40 I

Table 4-4 Summary of Results for Thermal Transient Parameters (for lower horizontal section of surge line piping)

Hax Cumulative Number of Max Signif.

I delta T Significant Reversals at > AT

$_tricina del t al Transient Reversal 50/100/_L50/200/250/300/350/400 EAll fMlura 260 NA HUlAl 298 96/ 68/ 26/ 15/ 8/ -

197 NA l

Hula 2 295 94/ 63/ 25/ 15/ 8/ -

203 NA HULA 3 295 92/ 63/ 26/ 13/ 8/

197 NA Hula 4 295 90/ 57/ 23/ 16/ 8/ -

201 NA HULA 5 295 88/ 57/ 27/ 14/ 8/

198 NA HULA 6 295 84/ 52/ 22/ 15/ 8/ -

197 NA HULA 8 295 90/ 57/ 23/ 13/ 8/ -

260 260 Hula 9 298 96/ 65/ 29/ 14/ 8/ -

Hula 10 262 96/ 61/ 27/ 13/ 3/ -

180 180 196 NA 00181 218 56/ 39/ 18/ 2/

C0102 218 52/ 36/ 17/ 2/ -

196 NA 196 NA CDlB3 218 54/ 38/ 17/ 2/

I CDlB4 218 48/ 34/ 15/ 2/ -

196 NA C01B5

<------------------ (same as CDlB3) ------------------->

CDlB6

<------------------ (same as CDlB4) -------------------->

CDIB8

<------------------ (sama as C0184) ------------------->

199 199 CDIB9 223 54/ 38/ 19/ 2/ -

199 199 C01810 223 54/ 39/ 19/ 2/

154 154 HUICI 310 28/ 20/ 19/ 12/ 9/ 1/

TRAN2A 142 2/ 2/ -

137 TRAN2B 165 2/ 2/ 2/ -

142 TRAN3 115 2/ 2/ -

154 98 I

149 1RAN4 137 2/ 2/

TRAN5 72 2/ -

119 127 TRAN6 TRAN7 138 2/ 2/ -

139 TRAN8A 135 4/ 2/ -

108 TRAN88 158 2/ 2/ 2/ -

113 I

TRAN8C 129 2/ 2/ -

118 TRAN9 64 2/

TRANIO 72 2/ -

119 TRAN13 TRAN14 81 2/ -

125 4-41

i d

)l Table 4-4 Summary of Results for Thermal Transient Parameters (cont )

Max Cumulative Number of Max Signif.

4 delta f Significant Reversals at > AT S.tr_ipjAQAtl.tLI i

Ir_4nsient flgygr.nl

_hDR00/150/2qQR5DL1091150/400 East f_Rtura i

1 TRAN19 62 4/ -

i TRAN20A 51 2/ -

TRAN20C 99 TRAN20D2 82 2/

TRAN20E S9 2/

101 101 TRAN22E 141 2/ 2/ -

!I

'I il 1

1 i

l 1

l il 4-42 lI d

-_,..,.,._r.

--....,s_

. ~. -,

Figure 4-1.

Design Heatup Transient Temperatures

[

Pest Transients 700-l.b

$00 f

1 A 1. 3. 5 p

IA2 l

500 c

eressuriter 5 400 f

d -~ ~ ~ ~

I L1A4.6,8 e

300 Res

~

200

'k 1A1 0

O 25 50 75 100 125 150 175 Tirne, hours Future Transients 700 600

-~

Pressuri:er f.._

500 1A9 400

_ y-t 1110 g300 g

e 200 100 1A9 0

0 25 50 75 100 125 150 175 Time, hours 4-43

I figure 4-2 Design Cooldown Transient Tenperatures Past Transients I

600 I

\\ '.

181 t

183 g'

/-182 500 g

y,... r....l....

Pre ssurizer 2...

3 I-1B4 thru 8

~

I

@ 300 P.C S 6-I 200 100 O

i I

0 m

e g

g Time, hours I

Puture Transients T

600

~

180 s

1B 10's 500 t

J Pressurizer I

B 400

-~

C 300 IB9 and 10' 200 100 0

s 0

20 40 60 80 Time, hours 4.o lI I

5.

PIPING ANALYSIS 5.1 Structural loadina Analysis The structural loading analysis which generates the internal forces and moments in the surge line for the thermal stratification conditions defined in the design basis transients is essentially the same for Davis-Besse Unit 1 as that performed for the lowered-loop plants, with the exception of the gapped whip restraints for Davis-Besse Unit 1.

The structural loading analysis of the surge line for Davis-Besse Unit I was performed using the computer program ANSYS (Reference 5).

j i

.h. l.1 Mathematical Model An " extended" mathematical model was built consisting of the pressurizer, surge line, hot leg, reactor vessel, and steam generator.

The mathematical model of the Davis-Desse Unit I surge line is shown in figure 5-1.

Gap elements were used to model the whip restraints at the following locations:

I joint 48 SL#1 (upper horizontal) joint 44 St#2 (upper horizontal)

I

~

joint 41 SL#3 (upper horizontal) joint 37 SL#4 (upper horizontal) joint 30 SLA5 (riser) joint 22 SL#6 (lower horizontal) joint 18 St#7 (lower horizontal) joint 15 SL#8 (lower horizontal)

The veadweight support, PSU-ill at joint 36, a Grinnell Type f spring support w;;

modeled as a gap element and a spring element.

If conditions cause the spring support travel to be exceeded, the spring support will bottom out and become rigid, thereby, supporting the pipe in a compressive mode from below, in addition, the past experiences of snubber PSU-R1 interference at joint 36 were 5-1 l

?

L modeled as a gap element.

Before 1984, there were two snubbers located where

[~

pSU-R1 is currently positioned, one on each side of the surge line, in 1984, the L,

snubber closest to the wall was found to be broken and both snubbers were re aced with one snubber, the current pSU-R1, located on the side of the surge li..e farthest from the wall.

The stanchion (3/8" x 4" x 4" structural tubing) closest to the wall was also removed.

The interference prior to removal was considered in the analysis.

h12 RQn-linear lemarllyre prefily As for the lowered-loop plants, the temperature profile on the pipe cross-section for the Davis-Besse Unit 1 plant is non-linear in the horizontal portion of the l

surge line..

A study of the non-linear temperature profile was repeated for Davis-Besse Unit 1 in the same manner as performed for the lowered-loop plants in order to find an " equivalent linear temperature profile",

four non-linear measured te mperature profiles at the outside surface were selected for the lower horizon'.al and another four for the upper horizontal. These selections exhibited I

th; most non-linear temperature profiles as described in subsection 5.1.3 of the main report.

A non-linearity ccefficient was calculated for each of these

_l profiles, using a piece-wise integration of the actual temperature profile on the pipe cross-section.

The non-linearity coefficient is used to obtain an l

equivalent linear top-to-bottom temperature profile which produces the same rotation as the non-linear temperature profile.

Therefore, the non-linearity coef ficient is actually the ratio of the rotation produced using the actual top-to-bottom temperature profile and a linear temperature profile with 'he actual top and bottom temperatures.

I Finite element conduction runs were used in an iterative process to match the calculated outside temperature profiles to the outside measured profiles and subsequently to obtain an average temperature profile, lhus, an equivalent linear temperature profile with an average modulus of elasticity and an average coef ficient of thermal expansion was obtained for each non-linear measured temperature profile with the modulus of elasticity and the coef ficient of thermal expansion varying on the pipe cross-section.

Using the non-linearity coef ficients from the piece-wise integration for each profile, the mathematical 5-2

I l

formula developed for the non-linearity coefficient for the lowered-loop plants was modified to obtain a worst case profile for the lower horizontal and a worst g

case profile for the upper horizontal of the Davis-Besse Unit I surge line.

Toledo Edison has cormnitted to making tndifications to the surge line insulation in order to eliminate excessive heat losses in the surge line.

These modifications are expected to result in stratification temperature differences in the surge line comparable to the Oconee measurements.

Therefore, the more conservative formula for the non-linearity coefficient which was developed for l

the lowered-loop plants was used for future transients.

51L. Verification Run for Dispigtnattij During the Davis-Besse "ntt 1 June 1990 heatup, the temperatures and the displacements of the surge lina were recorded. 1he instrumentation locations are shown in figure A-2 in Appendix A.

This data was used to verify the Davis-Besse Unit I nathematical model using the methodology from the lowered-loop analysis.

I Table 5.1 gives the temperature measurement locations with the measured top and bottom temperatures, as well as the top and bottom temperatures adjusted by the non-linearity coefficient to obtain the equivi..

linear temperature profile.

The non-linearity coefficient for these profiles were calculated using the piece-l wise integration described in Section 5.1.2.

The adjusted temperatures were given as input into the surge line ANSYS mathematical model shown in Figure 5-1.

The displacements obtained from the ANSYS computer run were compared to the measured displacements and are in good agreement as shown in figures 5-2 through 5-4.

I 5.1.4 Structural Loading Analysis for the ThersllLrat i ficat ion Condit ions The thermal stratification conditions were defined in the design basis transients documented in Section 4.5.

The structural loading analysis for the Davis-Besse Unit I surge line included the pressurizer, the hot leg, the steam generator, the reactor vessel, the surge line, and the surge line supports and restraints. This analysis was performed using the non-linear effects of the gap whip restraints, the restricted hanger travel, and the snubber / stanchion interference with the I

5-3 I

L wall (corrected in 1984).

The surge itne and other reactor coolant system

[

compcnents were represented linearly. Running the ANSYS model with the gaps for all gap and temperature conditions would have been prohltiitive; theref ore, base cases were executed.

These base cases were utilized to develop an iterative interpolation scheme to obtain the forces anu moments in the surge line.

1his g

a method was verified by comparing ANSYS results to the results of the iterative scheme for the same gaps and temperature condit tons.

Thus, using a nominal n eber of ANSYS non linear base cases, the resulting loads were determined for all the te v rr.ture and gap conditions.

I id Gfnttillhf1Lindkt.s for tht_httqtlint[lts Puoly clastic and elasto-plastic finite element analyses were performed on the four 90 degree elbows and the 45 degree elbow of the Davis-Besse Linit I surge line. The method was the same as that used for the lowered-loop analysis. The resulting generic stress indices for the surge line elbows were found to be the same as for the lowered-loop plants:

C, - 1.58 and K, 1.48, f or a total peak I

stress index K,

C, - 2.34 The stress index was shown to be conservative for the 45 deg elbow. These stress indices were used in the Iatigue Analysis of the surge line.

The use of these lower stress indices is permitted by Subparagraph NH-3f;81(c) of the ASME Code with justification by appropriately referenced specific data. Nfb 3213.11 defines peak stress as that increment of stress which is additive to the primary plus secondary stresses by reason of local discontinuities or incal thermal stress including the effect of stre:.s concentrations.

The basic I

characteristic of a peak stress is that it does not cause any noticeable distortior ind is objectionable only as a possible source of a fatigue crack.

The "clasto-plastic" finite element analysis of the ten inch Schedule 140 elbow, subjected to the thermal stratification piping loads showed that the highest stress intensity is highly localized as in the case of a stress concentration.

The analyses also showed that the elbow still behaves in a very linear fashion af ter the highest stressed locations enter the elasto-plastic domain.

These 5-4

L highest bending stress intensity values have a negligible impact on the i

distortion of the elbow.

L Based on these findings, it was concluded that a distinction could be made between the secondary and the peak stress intensity associated with the specific g

I geon.etry and loading.

The default indices of NB-3680 make no such distinction for an elbow subjected to thermal stratification piping loads.

Using the code definitions for stress categorization and the results of the "clasto-plastic" finite element analyses, that portion of the elbow stress associated with a l

stress concentration and not associated with the deformation of the elbow, was effectively reclassified from the secondary category to the peak category.

The stress-strain curve used to derive the stress indices or to calculate the j

highest secondary bending stress in the surge line elbow was consistent with the I

ASME code.

in the elastic range, the E value was taken directly from the Code (27000 ksi for an average temperature of 300T).

In the plastic range, the E l

value was conservatively low when compared with tests of similar materials at much higher temperatures (6007). The 35m discontinuity point, where the purely l

elastic and plastic slopes meet, was per the ASME Code. The stress allowable for the thermal stress range (EON.12) is 3Sm.

If the stress range exceeded 35m at some locations on the pipe cross-section, this was taken into account. For this reason, a bi-linear stress-strain curve was given as input for the elbow material, with a stress limit of 35m for the change in material behavior.

If g

I local stresses exceeded 35m, the stress redistribution on the pipe thickness lead to a higher rate of increase in the equivalent stress for that portion of the cross-section which was still in the purely elastic domain. This higher rate of increate in the equivalent stress was therefore greater than the one which occurs in a purely elastic analysis.

5.3 Verification of NB-3600 Equations (Equations 12 I

and 13. and Thermal Stress Ratchetin_q)

The Primary Plus Secondary Stress Intensity Range, Equation 10 of NB-3653 g

p (Reference 6),

exceeded the 3*5m limit for the most critical thermal stratification cycles.

The excessive stress intensity range occurred most typically for the thermal stratification cycles associated with very high top-to-5-5 l

l

___________m.______

- - - * = - - - -

e s

bottom temperature differences in the surge line linked with high temperature

(

flushing events.

For the load sets which did not satisfy Equation 10, it was necessary to verify Equations 12 and 13 of f4B-3653.6 and the Thermal Stress r

Ratcheting Equation of NB-3653.7. lhese verifications were perf ormed using the L

method applied to the lowered-loop plants as described in Subsection 5.3 of the nain report.

Equation 12 requires the calculation of the secondary stress range due to thermal expansion and a comparison of the secondary stress range to the 3*5m allowable.

l The secondary stress range of the Ocvis-Besse Unit I surge line was calculated from the thermally adjusted internal forces and moments associated with the most severe range of thermal stratification conditions.

The thermally adjusted internal forces and moments were the internal forces and moments from the thermal stratification structural loading analysis multiplied by the ratio E,,,,/E,...

This was perfor...ed in accordance with NB-3672.5. T he ra t t u E,,,,/ E,,,,, whe re E,,3, was taken at the ambient temperature of 70*f, was always greater than 1.0.

Equation 12 secondary stress was verified at every surge line location, using the simplified ASME-Code equations and the conservative generic elbow stress indices, l

1.58 (derived as de>cribed in Section 5.2), except for elbows A, D, and E (see figure 5-1).

1 The three elbows which f ailed the simplified Equation 12 were addressed by means of an elasto-plastic t inite element analysis.

For each of these elbows, the thermally adjusted internal forces and moments from the most severe range of thermal stratification conditions were applied on the finite element elbow modci documented in Section 5.2 and figure 5.5 of the main report.

Under these conditions, the three elbows are subjected to intarnal forces and moments in all three directions:

torsion, '.n-plane bending and vut-of-plane bending.

An elasto-plastic finite element analysis was performed for each one of these three elbowc, applying the " actual" thermally adjusted internal forces and moments from the most severe range of thermal stratification conditions.

For out-of-plane g

bending, the C, stress index is 1.58; for in-plane bending, the C, stress index II is 1.3, and fc r torsion, the C, stress index is 1.0.

Therefore, depending on the orientation of the applied moments, the "attual" equivalent strest index can be between 1.0 and 1.50.

The method used to calculate the resulting maximum 5-6

s secondary stress was the same as that used for the generation of the generic I

stress index C, in Section 5.2.4 of the main report.

For these elasto-plastic finite element analyses, the 3*Sm allowable was taken at the average temperature r

in each elbow.

The maximum resulting secondary stress was calculated to be l

5 smaller than the 3*Sm allowable for each of the three elbows.

Equation 13 involves the calculation of the primary plus secondary membrane plus bending stress intensity, excluding thermal expansion, and comparing the total resulting stress with the 3*Sm limit. [quation 13 stress is due to dead weight, l

operating pressure and Operating Basis Earthquake (OBE). The surge line does not contain any material or thickness discontinuity.

Therefore, the third term of Equation 13 stress was equal to zero (no variation of modulus of elasticity and no abrupt variation of average temperature in the axial direction of the piping).

Equation 13 stress was shown to 'e acceptable at every surge line location for I.

Davis-Besse Unit 1.

lhe maximum Equation 13 stress was equal to 26.4 ksi, which is 53% of the 3*Sm limit of 50.1 ksi.

ihe maximum Equation 13 stress occurred at the drain line of toe lower horizontal section of the surge line.

The verification of Thermal Stress Ratcheting consisted of comparing the highest occurring Delta 1, range with an allowable value to be calculated in accordance with NB-3653.7, where Dt.'ta 1, range is the range of the linear through-wall temperature gradients.

The verification of Thermal Stress Ratcheting was performed in the fatigue analysis of the surge line described in Section 5.5 and I-was found to be acceptable (by at least 15%).

5.4 Developmeni of peak Stresses The development of the peak stresses due to fluid flow, thermal striping, and the non-linearity of the temperature profile was discussed in detail in Section 5.4 of the main report.

The development of the peak stresses for the Davis-Besse Un#t I surge line was accomplished with the same method utilized for the lowered-loop plants.

I 5-7

I 5.4.1 Peak Stresses Due to Fluid Floy The peak stresses due t,a fluid flow were calculated as described in Section 5.4.1 of the main repert. These stresses are due only to the through-wall temperature gradient.;

if the peak stresses due to fluid flow were greater than the maximum thermal striping peak stresses, they were added to the peak stresses due to thermal stratification indeced bending moments in the fatigue analysis of the surge line as explained in Section 5.5 of the main report. The water temperature I

ramp rates and fluid P.w rates were used internal to the code which calculated the through wall gradienu in order to obtain an appropriate film coefficient per B&W standard film coeffic.ont correlations.

5.4.2 Peak Stresses Due to Thermal Strinina from the lowered-loop surge line analysis, a striping period of 4.0 seconds for as-welded locations and 0.5 seconds for locations remote from welds was found to result in the maximum possible peak stress intensity range due to thermai striping.

These critical striping period! were used to determine the maxirum possible peak stress intensity ranges for the Davis-Besse Unit I strge line. Ti.e temperature variations on the pipe thickness as a result of thermal striping were I

calculated through ANSYS iinite element analysis.

Given the inside metal temperature variations on the pipe thickness, the linear and non-linear temperature gradients were calculated using the equaticns given in NB-3653.2, leading to the maximum peak stress values due to the* mal striping.

Film l

coefficients were built into the striping correlations.

The " cut-sawtooth" pattern used for the calculation of the stress intensity ranges,as compared with the wave-forms from the Battelle-Karlsruhe erper ients.

(See Section 2 of Appendix C of the main report.) The comparisen showed that the ani,1yzed " cut-sa a.,th" patterns are representative of the measured wave forms.

I Tne modifications described in Section 5.1.2 are encettu to result in stratification temperature differences in th; surge line comparable to the Oconee measurements which were used in the lowered-loop analysis.

Therefore, the ma::imum significant striping delta-T results calculated for the lowered-loop surge line design transients for operations at high temperature (i.e., power operations) were assumed for future transients for Davis-Besse Unit 1.

5-8 I

I

W wi s

Monitoring of the Davis-Besse Unit I surge line will be necessary following the

(

changes to confirm the existence of stratification delta-1 magnitudes similar to those measured at Oconee.

5.4.3 ErALStrfanLQur_10 the tiprtLitturity of ttte_.JImpfratute_frofile

{

The peak stresses due to the non-linearity of the temperature profile were calculated as described in Section 5.4.3 of the main report.

This peak stress reflects the non-linearity of the top-to-bottom temperature profile, usually eferred to as the delta 1, peak stress.

The ABAQUS finite element model used for this analysis was the same as the model used in the lowered-loop analysis.

The model consisted of 27 rows of elements in the axial direction with each row containing 24 elements going from the bottom to the top of the pipe, covering an angle of 180 degrees. Boundary conditions were applied so that only half of the pipe required analysis. The peak stress intensity was calculated using the top-to-bottom delta-1, the elevation of the fluid interface centerline and the maximum difference between the actual ttmperature profile and the cauivalent linear top-to-bottom temperature profile, delta T.,

I 5.5 f atiage Artgly1LLof the Synge line A Code reconciliation was performed for the Davis-Besse Unit I surge line g

I analysis and t e results are identical to those listed in Section 5.5 of the main report.

The CoJ of record for the Davis-Besse Unit 1 plant is B-31.7 USA Standards (1968).

The fatigue usage factors for the Davis-Besse Unit I surge line were calculated l

using the methods described in Section 5.5 of the main report.

The thermal stratificat ton ceaks and valleys were combined into pairs for each joint of the surge line model, in calculating the main fatigue, the assumption was made that all of the cycles in the transients that produce thermal stratification can occur at any tirm.

Using that methodology, the different stress states were ranged using the following pattern:

s the highest possible state of stress (peak) with the lowest possible state of stress (Valley),

5-9

M e

the second highest state of stress (Peak) with the second lowest state of E

stress (Valley), etc...

L This procedure was used for the total number of occurrences of each peak and p

valley.

The main fatigue usage is the usage factor due to all the thermal stratification conditions which are characterized l'y a top-to-bottor temperature difference

~

(thermal stratification pV's).

These top-to-bottom temperature differences induce bending moments in the surge line, for the calculation of the main fatigue usage, the absolute values of the peak stress ranges f rom the following contril,utions were conservatively added:

1.

moment loading range due to thermal stratification, 2.

moment loading range due to OBE (for 30 future cycles),

3.

internal pressure range in tha surge line, 4.

non-linearity of the top-to-bottom temperature profile, 5.

maximum of the peak stress due to thermal striping or the peak stress due I

to fluid flow.

The addition of the contributions listed above is very conservative, because the assumption was made that the different peak stress ranges occur at the same location on the pipe cross-section.

The peak stress due to fluid flow and the additional peak stress due to the non-linearity of the top-to-bottom temperature profile were both directly considered in the calculation of the main fatigue usage.

These two peak stress contributions are not highly cyclic in nature and generally occur with moment peaks or valleys.

No additional fatigue usage was considered for the delta T.

stress.

However, there are fluid flow conditions, involving a complete or partial flushing of the surge line, which are not in concert with moment peaks or valleys.

These fluid flow conditions were considered in a separate fatigue analysis and the fatigue usage was added to the main fatigue.

5-10

I Thermal striping is a highly cyclic phenomcnon which also induces an additional f atigue usage. This additional f atigue usage was calculated and was simply added to the main fatigue usage. Again, these two contributions to the total fatigue usage probably do not occur at the same location in the pipe cross-section.

The moment loading due to OBE was considered in the calculation of the main fatigue usage, The OBE moments at each surge line location wer: conservatively added to the thermal stratification moments for the 30 most critical future thermal stratification ranges, assuming that one OBE cycle will occur exactly at the time of that most critical thermal stratification range. This procedure was done for the future only, as it is known that an OBE has not occurred in the past at Davis-Besse Unit 1, and 30 occurrences of an OBE have to be assumed for the 40-year plant life.

In addition, a total number of 650 OBE cycles must be assumed for the 40-year plant life. As 30 of these cycles were considered in the main fatigue usage, an additional fatigue for 620 independently occurring OBE cycles was added to the sum of the main fatigue usage and the fatigue usage due to the thermal striping cycles.

5.6 F_3tique Analysis Results for the Surce t.ing The total fatigue usage factors for a 40-year plant life (including past and future fatigue) are listed in Table 5-2, All total fatigue usage factors were less than the allowable of 1.0.

The highest cuenlative f atigue usage f actor was 0.62 and occurred in the lower horizontal run just past the first elbow (A) from I

the pressurizer.

The highest usage factor for an elbow was 0.56 and occurs in the first elbow (A) from the pressurizer.

The f atigue usage factor for the stanchion was 0.08.

I I

I 5-11 I

I

Table 5-1, Measured and Equivalent Linear Temperature Profiles L

(($l[E"j ttEr((iN[ <n l,Q YtNIYANab (sttISt A-2) ltop Tbottom

( 5" " W"I 5 0 Itop 1 bottom 110T LEG 131 131 1101 LEG 131 131 LOCATION 2 146 126 50-53 136 136 LOCATION 3 149 129 45-50 149 126 LOCATION 5 149 129 42-45 150 128 LOCATION 6 149 130 36-42 150 131 32-36 139 139 LOCA110N 9 195 126 30-32 172 172 24-30 194 194 21-24 185 115 LOCATION 10 197 124 18-21 187 114 LOCATION 12 199 127 13-18 187 116 LOCA110N 14 201 126 11-13 189 117 9-11 189 117 4-9 201 201 I,380 PRESSURIZER

_ 380 380 PRESSURilER 380 I

I 5-12 i,

I Table 5-2.

Total Fatigue Usage Factors for the Davis-Beste Unit 1 Surge Line I

Surge Line Maximum Locations Usage factor Most Critical 0.62 Straight-I Most critical 0.56 Elbow I

Drain Nozzle 0.08 Branch Stanchion 0.08 I

I I

I 5-13 I

I

sum ums uma sus amm aus ams nem nas mas sus amm aus sua mas sum

&++-

Figure 5-1 Surge Line Mathematical Model NSSS PLANT l t

,4 Stanchion 14 DB1 l

3#

33 35 g

l 36 32l

s PRESSURIZER 37 31 33

'N 39 l

x 40

(

'N[' 42 i

x.

43

'x l

30 45 l

s EWOWB

'~ *8 HOT LEG T

13 I

29 10 47 f4 5

,/

'N9 53 48 g

N 51 15 /

,/ 5 49 50 p

ms, 6

l 16 8 7 2al tr/_

EWOW E t9 EWOWA l

/

27

/.

20

,/22 2s -

Y 25 a

,_Dra,in L,ne l

i EWOWC Z

x

'N[

t

~

FIGURE 5-2 Comparison of Surge Line Displacements For 6/13/90 at 03:29:59.9 hrs.

0.1 7 005 l

u..

n w n 0,

3 n

f -0.05

^

8

~

w b

$ #l L~

9 U -0.15 c

.9 M

-0.2 6

54

$ -0.25

-0.3 O

10 20 30 40 50 60 Horizontal Distance from Pzt (feet)

MEASURED ANSYS CALCULATION WHIP RESTRAINTS HANGER

^

O O

FIGURE 5-3 Comparison of Surge Line Displacements For 6/13/90 at 03:29:59.9 hrs.

l l

0.1 7

0

^

E

._E

~

r

" 0.1

?.

5-0.2 T

8 s

-a g-0.3 o

x 8

B -0.4 (

3 2

g 4-G.5 3

-0.6 O

10 20 30 40 SO 60 Horizontal Distance from Pzr (feet)

MEASURED ANSYS CALCUl.ATION WHIP RESTRAINTS HANGER O

- ^

O

~

FIGURE 5-4 Comparison of Surge Line Displacements For 6/13/90 at 03:29:59.9 brs.

0.1 l

7 an a

,s j

04 r**

g

~

E-0.1 sa e

E T

8 s

.9 -0.2 9

~

is g

E -03 4)

O l

'8

~

.0 7-0.4 f3 O

N L

i t

i I

I

.5 0

10 20 30 40 50 60 Horizontal Distance from Pzt (feet)

MEASURED ANSYS CALCULATION WHIP RESTRAINTS HANGER O

O

^

l

l 6.

N0ZZLE ANALYSES in addition to the structural analysis of the surge line described in Section 5, detailed stress analyses of the pressurizer and hot leg nozzles have been performed to demonstrate compliance with the requirements of the ASME Code, Section 111. The thermal and pressure parameters for each nozzle are described in the design basis transients of Section 4.5.

In addition, each nozzle is subjected to piping loads from the surge line itself.

These loads have been taken from the piping analysis described in Section 5 and the resulting stresses have been combined with those from pressure and thermal loadings.

Detailed descriptions of the analyses of the pressurizer surge nozzle and of the l

surge line to hot leg nozzle are contained in the following sections.

The analysis of the surge drain nozzle is part of the surge line structural analysis described in Section 5 (Table 5-2 gives the total fatigue usage for the surge drain nozzle of Davis-Besse Unit 1).

I 6.1 Pressuri'er surae Nozzle 6.1.1 Methodolooy Modifications from lowered-loop Analvslji The purpose of this supplemental report is to describe the evaluation cf the raised-loop pressurizer surge nozzle of loledo Edison's Davis-Besse Unit 1 (08-1). The method used for the analysis of the Davis-Besse Unit I nozzle was basically the same as that outlined in Section 61 (main report) for the lowered-loop surge nozzle evaluation.

However, becau:,e of differences in transient thermal data, resulting external loads, and the experience gained from the lowered loop evaluation, some modifications to the analytical method were made.

This supplemental report describes those Todificatinni, and summarizes the results of the Davis-Besse Unit I nozzle evaluation.

6-1

s 1.

The geometry of the Davis-Besse Unit I nozzle is identical to that of the lowered-loop plants shown in Figure 6-1 (main report).

Therefore, the finite element model used for the lowered loop nozzle evaluation (shown in figures 6-1, 6-2, 6-4, 6-5, 6-6, and 6-7, main report) is applicable to the Davis-Besse Unit I nozzle.

-2.

New transient thermal and pressure conditions were defined for the Davis-Besse nozzle; the Davis-Bessa data are discussed in Section 4.5.

In addition to containing a larger number of peaks and valleys, the Davis-l Besse date include ramp rates much greater than those specified for the lowered-loop plants. Because of these higher ramp rates, additional base case runs were made to insure that all transient ramps would be bounded by a base case.

3.

In the lowered-loop evaluation, the thermal stresses for the chosen base case were used directly. For the Davis-Besse Unit 1 evaluation, the base case thermal stresses were adjusted by the ratio of the transient AT-to-base case AT.

l 4.

For the lowered-loop evaluation, the moments used were based on the moment set occurring at the time closest to that of the transient pv.

For the Davis-Besse Unit 1 evaluation, the external loads at the exact time of the transient pv are defined and used.

This allows for a more accurate time I

phasing of 1004 than was used (available) in the lowered-loop evaluation.

~5.

The ramp rates usad to define the pv base case in the lowered-loop evaluation were based on the maximum rate at any time throughout the l

transient pv.

The ramp rate used for the Davis-Besse Unit 1 evaluation was based on the change-weighting technique discussed in Section 4.2 under

" Correlation."

6.

In the calculation of tiie fatigue usage factor for the lowered-loop g

p plants, two separate usage factors were determined and then added together. One value was the transient-to-transient usage factor which was determined from the absolute maximum and minimum stresses from each 6-2

k transient. The second value was the transient internal usage factor which

(

was determined for each transient based on all other transient stresses.

~

for the calculation of the Davis-Besse Unit I usage f actors, all pvs are conservatively used in the transient-to-transient approach to determine J

one complete usage factor.

7.

The evaluations of the safe end and safe end-to-elbow weld for the lowered-loop evaluation were conservatively combined into one cvaluation.

l The analysis conservatively used the stress indices from NB-3600 (piping, radial gradient stress is peak) for the as-welded condition and applied them to the stresses as classified in NB-3200 (components, radial gradient is secondary).

I for the Davis-Besse Unit 1 evaluation, the safe end and safe end-to-elbow weld were evaluated separately; this is in accordance with ASME Section 111, NB-ll31, which states that the connecting weld shall be considered part of it s niping. The safe end was evaluated using the requirements of NB-3200 i gradient is secondary) without the stress indices of the as-welaec aition.

The safe end-to-elbow weld was evaluated using the requirements of NB-3600 (radial gradirnt is peak) with the stress indices of the as-welded condition.

I The items listed above provide a summary of the differences between the evaluations of the Davis-Besse Unit 1 pressurizer surge nozzle and the lowered-loop surge nozzle of Section 6.1 (main report) and retain a conservative basis for the stress and fatigue evaluations, it should be noted that other assumptions as listed in Section 6.1.4 (main report) remain valid for the Davis-Besse analysis. These original assumptions, when combined with the oifferences noted above, result in a conservative estimate of the fatigue usage.

6.1.2 Summan_sf Results and Conclnim The following table provides a summary of the results for the Davis-Besse Unit 1 pressurizer surge nozzle.

6-3

SUMMARY

OF RESULTS LOCATION fATICUE USAGE FACTOR _

ACTUAL ALLOWABLE SAFE END-10-ELBOW WELD 0.51 1.0 I

(STAINLESS STEEL) max SAFE END 0.29 1.0 (STAINLESS STEEL) max N0ZZLE-TO-HEAD CORNER 0 93 1.0 (CARBON STELL) max l

In suneary, the pressurizer surge nozzle, safe end, and safe end-to-elbow weld meet the requirements for Class 1 components of the ASME code, Section !!I,1986 Edition with no Addenda for the loading conditions identified in the Transient Specification for the surge line.

6.2 Hot Le.q Surae Nozzle This section completely replaces Section 6.2 in the main report. The purpose of i

this section is to describe the evaluation of the hot leg surge nozzle for Davis-Besse Unit 1.

Due to the differences (geometry and loadings) between the l

lowered-loop and raised-loop hot leg surge nozzle, an independent analysis for the raised-loop nozzle was performed.

The stress analysis of the nozzle and nozzle-to-surge line weld has been performed using the finite element method as implemented by the "ANSYS" computer code, Reference 5.

The loads used fcr evaluat n were the ther.ual and pressurt loads identified in the design basis transients for the surge line stratification (see Section 4.5).

The acceptance criteria for the evaluation were the requirements for Class I components of the ASME B&PV code, Section 111, 1986 edition with no Addenda, Reference 6.

The nozzle and nozzle-to-surge line weld were evaluated using the detailed requirements of Subsection NB-3200 as permitted by NB-3600, 6.2.1 Geometry An axisymmetrical representation of a small segment of the surge line, hot leg surge nozzle, and hot leg is shown in Figure 6-8 with an effective radius for the 6-4 I

L sphere (hot leg) equal to 3.2 times the hot leg pipe radius.

The 3.2 : 1 f

equivalent spherical vessel is a modeling technique recommended by Reference 7.

Using the 3.2 factor instead of the more common 2.0, assures that the maximum r

pressure stress at the critical location in the nozzle is adequately predicted L

by the axisymmetric model.

This modeling technique is conservative for predicting the membrane stress, but is accurate for predicting the maximum stress

(

in the critical locations for use in a fatigue analysis.

This piping junction consists of a carbon steel nozzle welded to the carbon steel hot leg. Both the

(

nozzle and hot leg are clad with stainless steel to prevent reactor coolant fluid from contacting the carbon steel base metal.

512 Descriplign of todjngs

~

1he loadings on the hot leg surge nozzle consist of thermal gradients, internal pressure, and external piping loads.

The thermal gradients are caused by the various fluid temperature swings (peaks and valleys) associated with the in-and out-surges of fluid between the e

pressurizer (hot) and the hot leg pipe (cold).

The surge line fluid thermal stratification can extend into the hot leg nozzle producing circumferential temperature gradients and thermal striping which are in addition to the l

axisymmetric (radial and longitudinal) temperature gradients produced by the transient. Also, the temperature differential between the surge line fluid and the hot leg fluid contributes to these temperature gradients.

The temperature swings for the surge nozzle and hot leg fluids are defined in the design basis for the surge line. The thermal gradients and stresses due to these temperature g

B swings are determined using the ANSYS finite element code.

The RCS pressure is applied to the internal surfaces of the nozzle and hot leg pipe. The pressures are defined in the design basis and the resulting stresses are determined using the AN!.YS finite element code.

There are significant external loads developed due to the heating and cooling of the surge line as well as the stratification in the surge line.

The external

loads, forces and moments, for each peak and valley are given in the 6-5

documentation of the surge line analysis. The stresses due to these moments and l

forces are calculated using the ANSYS finite element code.

L 6.2.3 Discussion of Analy11s in a typical fluid temperature spike, the top fluid in the nozzle will have a larger temperature change than the bottom fluid. Thus, for the determination of the radial and longitudinal temperature gradients and the associated thermal I

stress, it is conservative to use an axisymmetric analysis with the top fluid as the fluid boundary. An axisymmetric thermal and thermal stress analysis has been performed using the ANSYS finite element computer code. The transient thermal analysis consists of imposing time dependent boundary conditions (bulk fluid temperatures and heat transfer coefficients) on the finite element model. Nodal temperatures from the thermal analysis were stored on magnetic tape for each iteration (time step) of the transient. The ANSYS postprocessor " POST 26" uses the nodal temperatures to calculate Delta-T's between various locations in the structure. Tables of the Deita-T's versus time for each transient were used tc determine when the maximum and minimum stresses are likely to occur. The nodal temperatures for each critical time step were input to the ANSYS stress routine for the determination of stresses. The ANSYS postprocessor, " POST 11", was used to linearize the stresses at critical sections.of the structure.

Stresses due to pressure and resultant external force (along the nozzle axis) were also determined at the critical sections using ANSYS and POSTil.

I Due to the large number of temperature swings (peaks and valleys) associated with the thermal stratification transients, it was not practical to evaluate each peak and valley as an individual case.

Instead, a few " base cases" were created to envelop the large number of identified peaks and valleys.

The base cases were chosen using the following parameters; the hot leg fluid temperature, the maximum instantaneous ramp rate for the top fluid temperature excursion in the stratified nozzle, the top fluid starting temperature, the top fluid temperature change (Delta-T) betwaen a peak and valley in the nozzle and whether or not the RC pumps are on or off. The temperature distribution and resulting thermal stresses for each of the base cases were determined using the procedure described above.

Parameters describing the actual transients identified in the design basis were 6-6

l L

used to determine a representative base case for each peak and valley which will

[

now approximate the actual transient.

The linearized and maximum thermal stresses from the chosen " base case" were used directly for combining with stresses due to pressure, resultant external force, and non-axisymmetric load stresses.

Pressure stresses for a base case with an internal pressure of 1000 psi were

[

determined using ANSYS.

The pressure stresses for each peak and valley were determined by multiplying the stresses from the base case by the ratio of the I

actual pressure for the peak or valley from the design bases to the pressure used i

in the pressure base case.

The stresses due to a axial load of 10' lbs were determined using ANSYS.

The axial load stresses for each peak and valley were determined by multiplying the i

stresses from the base case by the ratio of the resultant force for the peak or valley from the surge line evaluation to the force used in the base case.

As described above, the thermal analysis performed is axisymmetric in that it is assumes that the top fluid concpletely fills the nozzle and creates the two-dimensional axisymmetrical temperature fields in the nozzle for the various thermal transients. The next task was to determine the additional stresses due to circumferential temperature gradients produced by the fluid stratification in the nozzle.

The stresses due to this fluid stratification were conservatively assumed to occur at the time of maximum thermal stresses due to the radial and longitudinal temperature gradients.

The stresses due to thermal stratification were determined for six base cases by use of the ANSYS harmonic element STif 25.

This element is used for two-dimensional modeling of an axisymmetric structure with non-axisymmetric loading.

In the case being considered, the non-axisymmetric loading is the temperature field in the nozzle which varies in the circumferential direction as well as in the radial and axial directions.

The stresses due to a circumferential temperature gradient are independent of the radial and axial gradients. These stresses are primarily a function of the temperature dif ference between the top f

and bottom fluid and the transition zone between the two fluid temperatures.

6-7 I

l l

~

The circumferential temperature gradient was approximated by assuming the top and I

bottom of the nozzle is at a steady state condition for a thermal peak and L

valley, respectively.

The transition between these two temperature fields was r

assumed to be linear over the same 1" height of nozzle that contains the fluid interface zone between the hot and cold fluid.

From the design basis, the centerline elevation of this interf ace zone in the hot leg nozzle, relative to the centerline of the nozzle, varies from 42" to -3" in steps of -1" during the various PV temperature excursions.

Thus, six thermal stratification load base I

cases were required.

I The six stratification base cases used a 200T temperature differential between the hot and cold fluid.

Therefore, the stress due to the base case circumferential temperature gradient could be determined by subtracting the steady state stress due to the radial and axial temperature gradients from the I

combined stress due to radial, axial, and circumferential temperature gradients (from harmonic element results which included the same sicady-state temperatures as were used in the axisymmetric load).

The thermal stratification st. esses dua to the circumferential temperature gradient for each peak and valley were determined by multiplying the stresses from the appropriate circumferential temperature gradient base case by the ratio I

of the actual stratification Delta-T for the peak or valley from the design basis to the Delta-T used in the stratification base case (200T).

Stresses for two base case nozzle bending moments of 10' in-lbs (MY and MZ) were also determined by using the ANSYS harmonic element, STif 25. The bending moment stresses for each peak and valley were determined by multiplying the stresses from the base case by the ratio of the moment for the peak or valley from the surge line evaluation to the moment used in the base case.

Shear stresses for a base case nozzle torsion load (MX) were calculated by hand for a unit load of 10' in-lbs. The shear stresses for each peak and valley were determined by multiplying the stresses from the base case by the ratio of the torque for the peak or valley from the surgo line evaluation to the torque used in the base case.

6-8

The thermal, stratification, pressure, and external load stresses were multiplied I

by the appropriate stress indices (Table NB-3681(a)-1) or stress concentration factors and then combined for determination of naximum stress and f atigue usage.

1he results are given in Section 6.2.9.

6 2.4 List of Assumoljanillnp_ujj_Used in Ana.lyJh 2

1.

The surge line nozzle to hot leg junction is a 3D structure and the thermal stratification in the nozzle produces non-axisymmetric loads. Two significant assumptions are necessary in order to analyze this nozzle using a 2D finite element model.

a.

The 3D structure can be approximated as a nozzle attached to a I

sphere whose radius is 3.2 times the radius of the hot leg.

This assures that the pressure stress in the model will be equivalent to the maximum pressure stress in the actual structure at the critical I

location.

The thermal stress from the model due to temperature gradients is approximately equal to those in the actual structure since thermal stress is not a strong function of the radius of the sphere.

b.

The circumferential temperature gradient is approximated by assuming the top and bottom of the nozzle is at a steady state condition for I

a thermal peak and valley, respectively.

The transition between these two temperature fields is assumed to occur over the same l' height of nozzle that contains the fluid interface zone between the I

hot and cold fluid.

This is a conservative assumption since heat conduction in the circumferential direction of the nozzle will increase the height of the transition zone in the metal which would tend to reduce the thermal stresses.

2.

The outside of surge line nozzle and hot leg are assumed to be fully l

insulated (no heat loss).

3.

The surge line nozzle and hot leg are assumed to be at a steady state condition at the beginning of each up ramp (peak) or down ramp (valley).

This is a conservative assumption since it maximizes the radial and axial gradients in the structure for each peak or valley, 4.

The fluid temperature ramp rate ("F/Hr) used in the analysis in determining the applicable base case is the maximum ramp rate at any time throughout the temperature change (PV) as defined in the design basis.

i l

6-9 l

1

5.

The transition between the nozzle and hot leg fluid temperatures is assumed to be a step change occurring at the intersection of the nozzle and hot leg pipe.

This is a conservative assumption as a more gradual transition will actually occur which would reduce the thermal stresses in this region of th2 nozzle.

6.

Fluid stratification is assumed to occur over the entire nozzle length l

(i.e., there is no mixing at the entrance to the hot leg).

7.

The outside nozzle surface temperature and maximum rate of change for each pV, as predicted in the design basis, is also assumed to be l

representative of the nozzle fluid temperature and ramp rate. This l

is an appropriate assumption for the majority of the PVs, however, it is slightly unconservative for some short duration transients (i.e., the outside metal temperature would not see the total fluid temperature change).

8.

The transient stresses are conservatively assumed to apply to all i

l angles around the nozzle and not just to the regions in contact with the top (hot) fluid.

l 9.

The temperature excursions predicted for the surge line upper horizontal run are assumed to be applicable for the hot leg nozzle.

10.

The OBE seismic events are assumed to occur at steady state conditions and not at the point of maximum or minimum transient stress. Even if an event wert to occur during a time of maximum transient strest, the effect on fatigue usage for only one occurrence of OBE would be minimal.

El 5 Thermal _ Anal 335_ij_pf Axisvmmetric Loyh An axisymmetric heat transfer analysis using the finite element code ANSYS was performed to obtain the temperature distributions in the surge line nozzle and hot leg.

The thermal transients evaluated are those specified in the design 6-10 1

I basis and discussed in Section 6.2.5.1.

The resulting nr,dal temperatures from the thermal analysis will be used as input to the stress analysis.

I The finite element model of a small segment of the surge line, surge nozzle, and hot leg is shown in figure 6-9.

These components are represented by I

isoparametric quadrilateral thermal elements, STlf 55. The required inputs for this element are four nodal points and material properties: thermal conductivity, density, and specific heat.

6.2.5.1 Selectji_on of Transients The operating transients for the surge line (and nozzles) are identified in the design basis for the surge line and discussed in Section 4.5.

A review of the transients revealed a significant number of temperature fluctuations during each transient.

The temperature fluctuations involved 50 to 60 different peaks or valleys per heatup or cooldown transient. The fluctuations include temperr.ture g

changes (Delta-T) with magnitudes ranging from approximh ely 40 to 400 degrees E

F.

As previously stated, an evaluation of each peak and valley was not practical, therefore only a few cases were considered. These cases are referred to as " base cases" and were selected to insure all peaks and valleys are enveloped.

I 6.2.5.2 Thermal Boundary Conditions The thermal boundary conditions consist of convective heat transfer at the inside surfaces of the model. Depending on the flow velocity in the nozzle and hot leg, either free or forced convection may be the predominant mode of heat transfer between the fluid and metal surf aces. For natural convection, the heat transfer is caused, primarily, by the difference in temperature between the metal surface 1

and the reactor coolant fluid. The film coefficient versus Delta-T is input in tabular form to the Af4SYS thermal runs. Af45YS uses the actual surf ace-to-fluid Delta-T at each time step (iteration) to determine the appropriate film coefficient. For forced convection, the film coefficient is constant for a given fluid velocity, temperature, and geometry.

The film coefficient used in the analysis is the maximum of the coefficients for free or forced convection. Free convection film coefficients for the nozzle are not used in any of the base cases 6-11 I

I

since the calculated force convection coefficient is larger.

A sample of the

{

film coef ficients for the two regions of the model is given below.

REGION FILM COEFFICIENT (BTU /HR-fT'-f)

FREE FORCED N0ZZLE 140 400 HOT _ LEG 115 3000 I

The outside surfaces of the model are assumed to be fully insulated.

in addition, for symmetry, the ends of the model are assumed to be adiabatic surfaces.

The transition between the nozzle and hot leg fluid temperatures is conservatively assumed to be a step change at the intersection of the nozzle and hot leg pipe.

6.2.5.3 Results of Ther. pal Analysis The results of the thermai analysis are in the form of nodal temperatures. These nodal temperatures were read into the thermal stress analysis and provided the I

model with the axial and radial thermal gradients that produce the thermal stress. The times at which the maximum gradients occur were used for the thermal stress analysis since they were likely to produce the maximum stresses.

To determine when the maximum gradients occur the ANSYS postprocessor, POST 26, was used. POST 26 provides a time history of the gradients at defined locations for the duration of the transient. For the :. urge nozzle evaluation fifteen pairs of nodes were used to examine the thermal response (Delta-T) of the structure. The locations of these 15 nede pairs at e shown in figure C-11. Figure 6-12 shows the temperature contours at an extreme Delta-T time point of a typical fluid temperature spike.

6-12 I

6.2.6 Stress Analysis of Axisvpmetric LQadi L

An axisymmetric stress analysis using the finite element code ANSYS has been performed to obtain the stress distribution in the model for the base case

(

axisymmetric loadings. The loadings for the analysis are the nodal temperatures from the thermal analysis (Section 6.2.5), a unit pressure load (1000 psi), and a unit axial force (10' lbs).

L L6.1 Description of finite Element _tigdel The finite element model used for the thermal analysis was also used for the stress analysis. The only difference between the two models is the element type designation.

The STif 55 thermal element was replaced with a isoparametric I

quadrilateral stress elements, STif 42. The required inputs for this element are four nodal points and material properties: coef ficient of thermal expansion, modulus of elasticity, and Poisson's ratio.

l 6.2.6.2 Structural Boundary Condilion 1

structural boundary conditions applied were required to simulate those portions of the structure that were not modeled.

The end of the model representing the hot leg was restrained from motion in the meridional direction (UY displacement = 0.0). This restraint simulates the restraint of the adjacent hot leg pipe material. The end of the surge line segment was assumed to be free.

The location of this free boundary condition is sufficiently remote from the I

nozzle-to-surge line weld such that any stress induced by the assumed boundary cadition will have attenuated to a negligible value at this critical section.

6.2.6 3 Selection of Transient Times As stated in Section 6.2.5.3, the selection of transient times for use in the stress analysis was dependent upon the thermal gradients through the stracture.

The thermal gradients cause differential growth between adjacent mater!al which results in thermal stresses.

The times at which the maximum radial and axial gradients (Delta-T) occur were evaluated for stress.

6-13 i

en de 62usd_ilRillfltmenLllrauJimilh r

L 1he results from the ANSYS stress runs are not in a format which can be directly compared to ASME code allowables, in order to get stresses compatible with the

[

ASME code requirements it was necessary to use the ANSYS postprocessor POST 11.

POST 11 performs stress linearization by converting the non-linear through-wall stress distributions into the stress components required for an ASME code evaluation: membrane stress, bending stress, and peak stress.

lhe pertinent I

information about the linearization methods and detailed input is given in Section f,.31 of the ANSYS users manual. Twelve stress classification lines (SCE) were selected to evaluate stresses in the various regions of the model. The line locations are shown in figure 6-13.

The sum of the linearized stresses for a given ioad set was used to compare to the ASME code limit for the range of primary-plus-secondary stress intensities (35m limit).

I in addition to the linearized stresses each " base case" also contains maximum stresses.

lhe maximum stresses represent the stresses at the surface of the I

component and are given in the ANSYS element su ass printout as the element surface stress, lhe sum of the maximum stresses for a given load set was used in the evaluation for fatigue usage.

S M S.iren_Atglyftjj of Non-AxisvmyLgjrlti ndJ A non-axisymmetric stress analysis using the finite element code ANSYS was performed to obtain the stress distribution in the model for the base case non-axisymmetric loadings. The loadings for the analysis were the circumferential nodal temperature gradients for six stratification cases and two nozzle bending moments as described in Section 6.2.3.

1 5,2. 7.1 De s rlplinn o f f i n i t e_f l eme n t Mod el r

The finite element stress model used for the axisymmetric loads was also used for i

the stress model for the non-axisymmetric loads. The only dif ference between the two models is the element type designation.

The STIf 42 element was replaced with a harmonic element, Silf 2S. The required inputs for this element are four nodal points and constant material properties: coef ficient of thermal expansion, modulus of elasticity, and Poisson's ratio.

6-14 t

1 1

...a

i I

6.2.7.2 St.ructural BoRDAstry Cq. ditions n

The structural boundary conditions are the same as was used for the axisymmetric loads in Section 6.2.6.2.

6.2.7.3 ANSYS load Sten Data A harmonic element model requires the load to be input as a series of harmonic functions (fourier series). The ANSYS preprocessor PREP 6 was used to generate the Fourier series for the stratification temperature fields described in Section 6.2.3.

Stresses were obtained for all the modes up through mode number 20.

I These stress modes were then combined using the ANSYS postprocessor POST 29. The unit (10' in-lb) nozzle bending moment was applied as described in Section 2.25, Case C of the ANSYS user's manual.

This bending moment was represented by applying peak axial (nozzle) force values at the end of the nozzle.

The load varies as a first harmonic wave (MODE 1) with a cosine symmetry condition (ISYM

= 1).

I 6.2.7.4 Finite Element Stress Resultji The stresses output from POST 29 were linearized at critical sections of the model. The results were then combined with linearized stresses from other loads and compared to the ASME code allowables as described in Section 6.2.8.

6.2.8 ASME Code Calculations The linearized thermal, stratification, pressure, and external load stresses for each peak and valley were combined to obtaio the total linearized stress.

The linearized stresses for all peaks and valleys were tabulated and the difference between the maximum and minimum linearized stresses was used for comparison to the ASME code limit of 35m.

When the 3Sm limit was exceeded, NB-3228.5

" Simplified Elastic-Plastic Analysis" was used to justify the stress conditions.

The maximum thermal, stratification, pressure, and external load stresses for each peak and valley were multiplied by the appropriate stress indices or stress concentration factor and combined to obtain the total maximum stress.

The I

maximum stresses for all peaks and valleys were tabulated for evaluation of fatigue.

A sample of the linearized and maximum stresses at the end of nozzle 6-15 I

I

E L

taper with the associated fatigue usage for a typical PV is given in Table 6-S.

[

The fatigue evaluation took into account the number of cycles for each peak and valley, the maxiroum stress ranges, the linearized stress range associated with the maximum stress range, and the resulting Ke factor for the maximum stress range when the linearized stress range exceeds the 35m allowable. fatigue usage due to thermal striping on the stainless steel regions of the nozzle was conservatively issumed to be equal to that calculated for the surge line.

The maximum linearized stress and f atigue usage f actor for both the stainless steel and carbon steel portions of the surge nozzle are given in Section 6.2.9.

All requirements of the ASME code are met.

5.2.9 5qrsty_0LBenth3_and C_0AclV5100 A summary of results for the hot leg surge nwzie evaluation is given in the following table.

Although the 3Sm limit is exceeded for both the carbon steel and stainless steel, the requirements of the ASME code were satisfied by performing a " Simplified Elastic-Plastic Analysis" as defined in Subsection, NB-3228.S of the code.

SUMMARY

Of RESULIS, H01 LEG SURGE N0ZlLE LOCA110N FAllGUE USAGE FACTORS ACTUAL Att0WABLE N0ZZLE-TO-SURGE LINE WELD 0.52 1.0 (STAINLESS S1 EEL)

N0ZZLC-TO-SURGE LINE WEl.D 0.20 1.0 (INCONEL)

END OF N0ZZLE TAPER 0.72 1.0 (CARBON STfEL)

N0ZZLE-TO-HOT LEG CORNER 0.76 1.0 (CARBON _S, TEEL ( _ _ _

in conclusion, the hot leg surge nozzle and nozzle-to-surge line weld meet the requirements for Class I components of the ASME Code, Section 111, 1986 Edition with no Addenda for the revised design basis transients discussed in Section 4.S.

6-16

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": ::3 E 15 l" g :a H

- 5:- 4 m

E:s:sr I m: m# I a 33 g

c 3.3 tI:23 a

8 aa

-r a c:

a va 6-17

~

s k

. Figure 6-8.

Geonetry of Hot leg Surge 140z:10 r

j i

I llo r. Lee i

I 38.6" IR 61.475" OR I

Cladding I

Surge Line-To-Nozzle '.leid

[

Surge Line I

i l

JJ u

5.375" 5.8125" 9.1875" 4.375" p

Center Line Of Nozzle 6-18

g4 m.

4 m

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!I 6-19 I

I I

I figure 6-4 4.

Location of Lkilta-T Values (Hot Leg Surge !;or:le) j l

fjf g

y I

/o 1

I O

s

~'

I g

e..

2 g - ;

- }Os i

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h-DE'.4'A-T LOCAT"NS 1 to 10 are radial Delta-T 11 to 15 are axial Delt+.-T I

6-20 I

I m.

n w

w

-9

k figure 6-12.

Hot I.eg burge Nortle Temperature Coritours i

(f) for a lypical PV L

/

(

/

t

/

(

1 4

9 F

t 350 325 [ \\

l o'{- -

%~N 400 4-rM& -

423 9

6-21

figuru 6-13.

Location of Stress Classification Lines

< am

~ s e n.,,a

'y j,

i f

x i

/

5 mm o

1 y),

q-ff y~

m f /!/

Detail A 1

O Ww

!O s.. no.n i

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. '3 w.-s e, ak= w%

t O

6-22

7,

$UMMARY Of RE$UI.15 1he Bl.W Owners Group developed a program to comprehensively address the requiremeet5 of NRC Dulletin 88-11 "pressuriter Surge tine lhermal Stratification.*

lhe Owners collected the necessary information required to evalitate the surge line, in addition to operational records and plant design inf ormation,,lant thermal stratification data and thermal striping test data were obtained.

lt was determined that the lowered-loop plant configuration and plant operations I

are sufficiently similar for a generic development of the design basis transients. lhe subsequent analysis and results for the lowered loop plants was documented :n BAW-21?7 of December 1990.

Ilowever, Davis-Desse Unit 1 (DD-1), a raised-leop plant required its own instrumentation and a separate, plant-spttific, set of design basis transients because of inherent dif f erences in it s dasign. These differences are discussed in Section 3 of both the main report (llAW-2127) and this supplement, lhe Davis-liesse Unit I analysis has been addressed in this supplement to the enain report.

Revised surge line design basis transients accounting f or plant evolutions af f ecting the surge line for the 40 year drsign lif e of Davis-llesse Unit I were dueloped.

The plant heatup and cooldown transients were thn most significant contributor to the fatigue usage factor for surge line components.

A structural loading analysis of the surge line was performed to take into account the global elfects due to thermal stratification. 1he resulting internal forces and moments were applied f or the f at igue stress analysis of the surge line and the associated nortles.

7-1 I

E

s The f atigue stress analysis considered the stress ranges for the global ef f ects

[

due to thermal s t r at i f ic at ion.

the locallied effects due to thermal otratification, the pressure ranges, the Operating liasis t arthquake, the thermal striping and the fluid flow conditions.

All resulting stress intensities were shown to be within their allowable limits. As a result of the f atigue analyses, the (umulative usage f actor is less than 1.0 at all locations of the surge line and its nozzles, in sununary, the f ollowing is a tabulat ion of the highest usage f actor f or the most important surge line components of Davis-liesse Unit 1.

Surge line Component Usage factor (40 year Lif e)

Surge line Ilbow 0.56 Straight Pipe Section

_0.6?

Drain liottle tiranch Connection 0.08 Stanchion 0.08 Pressuriier flozzle 0.93 llot t og floitic 0.76 In view of the conservatism accumulated ir, the synthuis of the design transients and in the analysis of resultant stresses, the*.e f atigue usage values provide assurance that the 40 year licensed lif e of Davis-Besse Unit I will be met with acceptable margin to accommodate normal variations in operations.

7-?

!I I

I 8.

DASES FOR Tile DAVI5-BESSE 1 ANALYSIS 1he generation of the revised Design flasis transients and the thermal stratification f atigue stress analysis of the surge line for Davis-!! esse Unit I were tased on conditions stated in this section.

I (he thermal stratification fatigue stress analysis was based on the following:

I for past operations, limitations to motion from supports, restraints, and snubbers were identified and taken into account. The following were included in I

the analyses:

[11tTAAl l00djtnak e

Restraint clearances with the restraint structures and the secondary concrete wall were taken into account throughout the bistory of operation.

I These restraint clearances were evaluated and extended unchanged from the last measurements until the 7th Refueling Outage (7Rf 0). After 7Rf0, the analysis assumes that no interference with other structures will occur, l

s An interference between snubber pSV-R1 stanchion and the 0-ring wall resulting in failure of the snubber occurred in 1984.

The interference dimensions were taken into account in limiting surge line displacement for all heatups and cooldowns until redesign of the snubber and elimination of this potential interference in 1984.

s Apart f rom the interference resulting in f ailure of snubber PSU-R1 in 1984, both past and future transients are based upon the amplitudes of surge line displacements within the free travel range of each snubber.

LI g

limitations on surge line downward displacement by reaching the hard stop a

g limit of spring support PSU-ill were taken into account for all transients 8-1 I

._ M

L, to the 8RIO.

Af ter BRl0, modifications to the spring support will eliminate this limit to displacement.

Branch moments at the surge line drain nortle connection within their respective maximur.. allowable (for deadweight, Operating Batis Earthquake, and thermal stratification) were assumed for past and future transients.

lhtrN1 hip 0MC Surge line thermal response on Davis-Desse Unit I dif fers f rom the lowered-loop plants and the measurements in Oconee I in two respects:

1.

The surge line insulation and restraint impact collars result in increased heat loss, particularly noticeable in the lower horizontal run of the surge line as the plant temperature increases.

?.

The makeup pumps during the heatup following the 6Rl0 were placed under automat ic control immediately prior to initiating reactor coolant pump operat ions.

The resultant makeup flow was cyclic, varying around the pressuriier water level setpoint and resulting in periodic reversals of I

flow in the surge line during most of the heatup.

lhe results of these observations were tasen into account in the analysis in the following manner:

a ror past transients and extending through the 8Rf 0, the thermal responses due to makeup valve cycling neasured following the 6Rf0 were l

conservatively taken into account although it is known tha* in some past heatups the valve was maintained in manual and the significant cycling during the heatup did not occur.

m Similarly, the measured temperatures representing the heat losses through the thennal insulation in addition to potential cooling ef fects from the PORV inlet drain were taken into account extending through the 7RIO.

e following the BRf0, the fatigue stress analysis was based upon the assumptions that cycling of the makeup control valve will be reduced and 8-?

A______-_________--_-_-_________________-__-_________--________-_____-

that valve cycling will not contribute significantly to thermal striping I

in the line for any of the design transients at power.

The striping L,

rec.ains the same f or heatups and couldowns, a

Also, for transients following the 7Rf0, the stratif; cation temperature dif f erences in the piping during steady reactor operation will be reduced b

to values approximating those measured in Oconee 1 and used in the analyses for the lowered-loop plants.

1he improvement will be

(

accomplished through upgrading in the thermal insulation at the time of modifications of the whip restraints to ensure freedom f rom interference.

lhe top-to-bottom temperature profiles on th t Davis-Desse surga line have s

been more nearly linear than those observed during the instrumented heatups at Oconee 1.

With the improvements in insulation, the conditions i.1 the Davis-liesse surge line will more nearly resemble those at Oconee 1.

Ilence, the non-linear profile developed for Oconee I was conservatively applied to the Davis-Besse analytical model for the operations following the 7RFO.

I Toledo (dison has committed to making the plant modifications necessary to ensure the validity of the bases stated above for future transients in terms of structural and support interferences, makeup flow cycling, and insulation, loledo Idison will monitor the surge line temperature following the restraint and insulation modification to confirm that surge line thermal response i.dequately l

supports the conclusion that striping f atigue does not challenge the full 40-year licensed lifetime of Davis-!3 esse Unit 1.

I Onenling_Linila The generation of the revised Design Basis Transients (for future events) was based on operating guidelines that limit the pressuriier to RCS temperature difference during plant heatups and couldowns (imposed with RCS pressure / temperature limits).

s h

8-3

+-

4

lhe heatup and couldown Design flasis transients defined for f uture operation will remain conservative if the RCS pressure is limited in accordance with figure 8-1.

1he curve shown in F igure 8-1 is a composite of various subcooling limit curves that vary over the range of RCS temperatures. The operating procedures at Davis-Desse will maintain pressure and tem..erature during heatup and cooldown operation to the right of the se:ected maximum allowed subcooling limits. (The curve shown

(

is similar to that developed for the lowered-loop piants except that a somewhat more restrictive limit of 280 psig has been placed on the pressuriier pressure when the RCS temperature is below 185f. This limit is consistent with the normal plant operations since the pressure is normally limited in this range of temperatures to avoid lif ting the Decay Heat Removal relief valves, which provide protection a'jainst overpressute at low RCS temperatures.)

lo meet the pressure limit specific f or heatup in the temperature range 70f to 150f, preheating the RCS has been recommended.

This may be accomplished by throttling back on the decay heat system cooling water (i.e., component cooling water) and/or bypassing reactor coolant flow around the decay heat removal heat exchanger.

The availability of decay heat and the requirements of the heatup schedule will dictate the capability of maintaining the recommended p/1 profile prior to achieving the condit ions necessary for starting an RC pump. The fatigue evaluation was performed on the basis that 85% of the heatups for the remt.inder i

of the plant life can meet the recommended limit shown by path CDEN in figure 8-1.

for those heatups involving pressurization at an RC temperature of 70f to 120i, a less restrictive limit is included in order to permit RC pump operation at lower RCS temperatures (path ABlN in figure 8-1) when core decay heat is nct adequate f or raising RC temperature. The f atigue evaluation was pertormed on the basis that 15% of the heatups for the remainder of the plant life will follow this heatup path, in summary, future heatups were divided into path COIN (85%)

and path ABLN (15%).

l I

I Figure 81. Surgo Lino Operational Limit I

2 4 00 -- --- ---- -- - - - - - - - = - - - - - - - - -

I 2200 F 001 10 trPy N

NOT LimR 1800 --

M i

Manimum Pressure Uma i

I 1600 (--

for the surge une 4

I

? 1400 h 6

CL

~

(_.t Press i Temp SC d

IA ' 232 ! % 0, 3*Al h

1200F' iDl 260 05 0 i 3*A' I

'Ci 150< 66 0 ! 300 Q.

0! 232 i 100 0 1 300 4

k 1000 g

!Ei 200 i 115 0 1 300 I

!F t 200 i 185 0 ' 230 l

IG I 3451 185 0, 250 l l 600 -

. - 5 E

E I

i j

i 477 1 265 0 I 200 '

,.El..72yJ_3gO__J_200_

600 -

LK]OLO i 346 5 i '?00 L,1.400j 389 01 200 I

, N 2.200 1 450 5 200l

..../

M g B00 4422O T E 400 p O,E

-- F B-mg 200l-

{

C' Surge hne lima appucable to l both heatup and cookkwn. !

I I

O l--

0 200 400 600 RC Hotleg Temporaturo, F Operating Umrt

-- Append'x G Umit I

m I

A

[

I L

9.

REFERENCES 1.

BAW-?l27, Ligl_Sybal111LigtJ[uchttjlegulgigry_SpagnhijgajlyllnLinJD-jl "PrtWtLiitt_luntLLint.lhttaLSitattif_itatlon," December 1990.

l 2.

NRC Bulletin 88-11 "Pressuriter Surge Line lhermal Stratification,"

December 20, 1988.

I 3.

BAW-2085, " Submittal in Response to NRC Bulletin 88-11 Pressurizer Serge line lhermal Stratification," May 1989.

4.

NRC Letter dated May 18,1990, J.1. Larkins to M. A. llaghi, " Evaluation of Babcock and Wilcox Owners Group Bounding Analysis Regarding NRC Bulletin 88-11."

I 5.

"#15J$" Computer Code. Versions

.lc and 4.3.

Ingineering Analysis System, User's Manual Volumes I and 11, Swanson Analysis Systems, Inc.

6.

"ASML lioller and Pressure vessel Code," Section 111, 1986 Edition with no I

Addenda.

7.

J.

B.

Truitt and P.

P.

Raju, "lbree-Olmensional Versus Axisyneet ric iinite-flement Analysi:. of a Cylindrical Vessel Inlet Nottle Subject to Internal Pressure -- A Comparative Study," ASML Pre;sure Vessel and Piping Division, Paper No. 78-PVP-6.

l l

l 9-1

W 10.

DOCUMENT SIGNATURES Septembe r 11, 1991 F

1his document prepared by:

b fi(LQp rl C. \\ l. D A e 4~ ~3 u u ---~

C K7 Ihaiid1R

[

_ g._g,,.

GTE7 eat Fr1F This document raviewed for technical content and accuracy by:

l QwaAwkw JT RT G16iidemans, Anilysis Wriices 7.E $6 *ti l

S T75hepard Material & Struct tral Analysis d ddk Material & Structural Analysis E STTowgW'=CiedicitiWTeams vs n

ian I

Verification of independent review:

C ETilly, aiager a

Performence lysis

[

,j g u

II Y$

L_. E.

re, Manager _

Materials & Structural Analysis This document approved for release:

W*

E.

. D5 aTsikT Engineering & field Services 10-1

L I

L 1

1 I

I I

APPENDIX l.

A.

Surge Line Data Acouisition at Davis-Desse Unit 1 CDlilullS Page l

1.

Thermocouple fabrication and Qualification A-- 2 B

2.

Displacement Transducer Description.

A-3 3.

Data Acquisition System Description and Operation.........

A-3 4.

General Description of Data....................

A-4 1

A-1

I Davis-Besse Unit I was instrumented extending the B&WOG program data as discussed in Section 3.1.

This appendix supplements Section 3.1 with additional detail on the data acquisition, L

ltwrrreteele f abricities antonalifitnion g'

The thermocouple assemblies and associated extension wire assemblies required for instrumenting Davis-Besse Unit I were fabricated by B&W Nu:lcar Service Company, thermocouple assemblies were f abricated from ANSI Type K, 20 gage solid thromel-5 Alumel c mm rcial grade assembly wire having p rallel c nduct rs individually B

insulated with ceramic fiber braid, an overall Jacket of ceramic fiber braid and stainless steel protective overbraid. The thermocouple junction was mechanically f ormed and then spot welded to a band which was later strapped around the surge

(

line.

Standard 2-pole connector plugs with integral cable clamp were attached to the ends of the thermocouple wires opposite the hot junction.

Thermocouple j

extension cable assemblies were f abricated from ANSI lype K, 20-gage solid Chromel-Alumel commercial grade extension wire having twisted connectors individually insulated with Tefzel, Mylar-backed aluminum foil shielding with drain wire, and an overall extruded leftel insulation jacket, l

Qualifying the commercial grade thermocouples f abricated for the safety-related pressurizer surge line temperature measurement application was accomplished by the standard practice of " t y p e testing, in this approach, duplicate thermocouple assemblies prepared from the same materials and following the same l

procedures that applied to fabricating the thermocouples installed at Davis-Besse l

were placed in a furnace (Jofra Temperature Calibrator) with certified Platinum Resist ance Thermomete. s (RIDS) and heated.

Comparison between the temperature registered by these qualification test thermocouples and the reference g

B temperature monitored by the R1Ds provided a means of qualifying the surge line t he rmocouples.

The furnace was set to ramp to a new level at a rate of 10 f or 20 f per minute to temperatures levels between 75F and 70bi. At each test level, the temperature was allowed to stabilize within lf for 3 minutes. The test setuo was then allowed to equilibrate for at least 5 minutes before the reference and test specimen temperatures were recorded.

il m

I

w Comparison of thermocouple data and the R1D ref erente indications dernonstrated

[

that thermocouple readings were consistently within 1.5I of the reference temperature. This agreement was well within the estabitsbed acceptance criteria p

at all test conditions for qualifying surge line thermocouples.

L L._DiiPlEMf!tllnuid titer.DhCLlRLiDB lhe displacement transducers manuf actured by Celesco Transducer products. Inc.,

I (Model p1101) provide an electrical signal proportionsi to the linear extension of a stainless steel cable. Displacement was measured by attaching the cable to the surge line and the body of the t ansducer to a fixed surf ace. Retraction is ef fected by means of a constant tension spring motor which maintains uniform l

tension on the cable. The manuf acturer specified the accuracy to be within 0.25%

for the expected range of surge line dispiacements.

I h._DALLAtmLilition.SM 11ELOOM EM10nELOPSD110B 1he schematic in figure A-1 depicts the general interface of components which make up tne data acquisition system.

The system included a fluxe llellos mainframe cor. trolled by a host Compaq computer utilizing Lablech Noteboak I

sof tware which was configured to receive the desired data.

The computer and Helios mainf rame, which were located in the Davis-liesse Unit I control room, interfaced with a remote Helios extender chassis housed in an instrumentation cabinet located in the reactor containment building near the pressurizer surge l

line.

This inst rument cabinet also contained the power supply, signal conditioning and other interfacing equipment required for the surge line therm 0 couples and displacenent transducers.

With the system installation complete, the integrity of each instrument and acquisition component was checked.

proper electrical loop resistances for the thermocouples, lead wires, and extension cables were verified. A polarity test and a complete checkout were then performed for all instrument channels.

Data collection was started before the pressurizer heaters were turned on and continued until the plant had reached power and rcmained there f or several days.

There were only short time periods in which data collection was interrupted in order to download t he daia f rom the host computer.

A-3

DWl45 and loledo (dison personnel monitored the data acquisition system to ensure

~

proper operation.

in addition, plant operations were monitored closely and the events which affected the surge line were noted. lhe operator and unit logs have

~

been obtained to assist in associating plant operations to surge line transients.

L L_0enenL0eitrMicrteLDitta g

1he temperature and displacement of the surge line were monitored, as well as the B

reactor coolant system (11C5) conditions, A liat of recorded parameters is contained in lable A-1.

lo record surge line temperature, thermocouples were placed at eight dif ferent axial locations (shown in f igure A-?).

1here are two or seven thermocouples at each locatton as given in figure A-?.

All locations contain a thermocouple at the top (11) and bottom (11) of the surge line. Where seven thermocouples are used, they are spaced with nual elevation dif ferences between thermocouples as shown in figure A-3.

The thermocouples are identified by axial location and relative position as follows: "1017" means t he thermocouple at location 10 and thermocouple position 7 as given in ligures A-? and A-3.

1he I

displacement of the surge line was monitored with position transducers (string potentiometers) at locations noted in f igure A-2.

position transducer, are identified by axial location and the measurement direction as f ollows: OllY denotes the potentiometer which Indicates the Y direction movement at location 1.

I plant data have been recorded for two heatups, one couldown and during periods of power operation.

The dat a were stored at sampling laterval time of 20 seconds. 1he large quantity of data (more than POO Megabytes of disk space) was stored in ASCll format and transferred to the 11WidS lip 9000 59rtes 800 computer I

where the data was processed into functional information (plots, calculations, etc.).

A4

s Table A.I.

Signal Identification

[

10!ilf.!UVH115 DISIlllEUM 0211 I

Location 2 Position 1 0212 I 2

2

[

0213 f 2

3 0214 f 2

4 02T5 f 2

5 0216 f 2

6 0217 I 2

7 1

3 1

031' t

03" r

3 2

0:

f etc.

0314 f 1

0315 f 0316 f l

0317 f 0611 f

06T2 f l

0613 f g

0614 f 0615 i 0616 f l

06T7 f 0911 f

location 9 Position 1 0912 f 9

2 j

0913 f 9

3 m

0914 f 9

4 0915 f 9

5 0916 f 9

6 0917 f 9

7 1211 F

12 1

1212 f 12 2

1 1213 f etc.

1214 f 1215 f i

12T6 f 1217 f 1411 f

1412 f i

1413 f 1414 f 1415 f l

1416 f P

1417 f 0511 F

tocation 5 Position 1 05T7 f 5

7 10T1 f

10 1

10T7 f 10 7

Mll4 f Mirror Insulation Temperature location 1, lhermocouple Pos.4 M2T4 f Mirror Insulation lemperature Location 2, Thermotouple Pos.4 A-5

lable A,1.

Signal identification (cont.)

L JDtnBZW1115 QMluP110ti 0111 1 Decay lleat Drop Line location 1, Thermocouple Pos, 1

[

0112 f Decay lleat Drop Line location 1, Thermocouple Pos.

2 D113 i Decay lleat Drop line Location 1, Thermocouple Pos.

3 D211 f

Decay lleat Drop Line location 2, Thermocouple Pos.

1 D311 f

Decay lleat Drop Line location 3, lhermocouple Pos.

1 D312 I Decay lleat Drop Cine Location 3, lhermocouple Pos.

2 I

D313 i Decay lleat Drop Line Location 3 Thermocouple Pos.

3 0411 f

Decay lleat Drop Line Location 4, lhermocouple Pos.

1 Ol2Y Inches location 1 Direction Y OlZZ inches 1

Z I

04ZY Inches 4

Y 04ZZ inches etc.

07ZX Inches I

072Y Inches 07ZZ inches 087X Inches 1

08ZY Inches 08ZZ inches ll2X Inches ll2Y Inches I

132X Inches 132Y Inches 157Y Inches I

RS46 Percent Auctioneered Average Power QS85 Tripped Turbine Trip Signal Master 1358 i Dil CLR 1 Out Terr,p 1361 f

Dil CLR 2 Dut lemp I

F4619 GPM llP INJ 1-1 flow I4649 CPM llP INJ 1-2 flew I4679 GPM llP INJ 2-1 flow F4709 GPM llP INJ 2-2 flow F592 GPM LP INJ 2 flow F593 CPM LP INJ l flow I

[S57 inches Pressucizer Compensated level 1776 f RC PkZR lemp, RC15-1 l

Z576 Position RC PRZR Spray line VlV, RC2 PS57 Psig RC Loop i llLG WR Press P721 Psig RC Loop 1 IlLG NR Press, RPS Cll!

T753 f RC Loop 1 IILG WR lemp, Clll IS63 f 1 cold Wide Range, Loop 1 3788 MW RCP l-1 MIR PWR J808 MW RCP l-2 MIR PWR J828 MW RCP 2-1 MIR PWR J848 MW RCP 2-2 MTR PWR I722 MPPil RC Loop 1 IlLG flow, RPS Clll f728 MPPil RC toop 2 HlG flow, RPS Cil2 f719 inches RC l.etdown flow F738 GPM RC MU flow, low Range A-6

I lable A.I.

Signal Identification (cont.)

JD11HRWillS ELSG1P110ti f740 GPM RC MU llow, liigh Range LB81 Percent SG 1 Operate Level, 9BI 1883 Inches SG 1 SU Range, 9B3 I

L891 Percent SG 2 Operate Level 9Al t893 Irahes SG 2 SU Range, 9A3 167) i M!i IW 1emp to ICS,111-1 I

P931 Psig SG 1 Out S1M Press, P112B1 1503 l

SG 1 Out SIM lemp P936 Psig SG 2 Out SIM Press, PT12Al 1504 i

SG 2 Out SIM 1emp I

ilX11 I

fluke llelios Extension Chassis inside I/O cabinet External lleat Exchanger inlet Temperature llX10 f

fluke llellos Extension Chassis inside I/O cabinet I

External lleat Exchanger Outlet Temperature lilli I

iluke llelios (xtension Chassis inside I/O cabinet internal lleat Exchanger Inlet lemperature 11110 f

fluke !!elios lxtension Chassis inside I/O cabinet I

External lleat [xchanger Outlet Temperature IP01 i

Ambient Temperature near Displacement Location 1 1P13 I

Ambient Temperature near Displacement locat.on 13 I

1R01 f

Ref lemperature in P7R room - 595' elv 1R02 f

Ref Temperature in hallway outside D-ring - $95' elv 1R03 f

Ref Temperature above P7R - 640' elv I

1R04 f

Ref lemperature at bottom elevation - 575' elv PSE1 Volts DC 20 Volt Power Supply PSL2 Volts DC Backup 20 Volt Power Supply CITI I

internal 1/0 Cabinet Temperature I

Cllll

'XRil Internal 1/0 Cabinet Relative llumidity R[7Y Inches Referente String Potentiometer CX11 i

External 1/0 Cabinet temperature CXill Ril External 1/0 Cabinet Rel?tive ilumidity C913 Percent Core Power f210 MPPil Core RC flow I

f460 GPM llP If4J 1-1 flow f488 GPM llP INJ ?-1 I1ow I489 GPM

!!P INJ 2-2 flow f490 GPM llP INJ 1-2 Ilow I

f712 MPPil RC Loop 1 lilG Ilow F713 MPPil RC toop 2 ilLG flow F717 GPM RC Letdown Ilow I

f782 GPM RCP Seai In flow f783 GPM RCP Seal Ret Ilow L769 inches RC PR7R AVG lVL P732 Psig RC LOOP 2 tilG WR Press, SI AS Cil 2 I

0613 MfPT 1 (Main feed Pump #1 Trip)

OG34 MIPT 2 (Main feed Pump #2 trip)

Q764 RC PRlk IllR SOURCL I

0810 RPS Cil 1 1 RIP A-7 I

I

huusJ 1able A.I.

Signal Identification (cont.)

[

1DN1fR/UNilS

[lLSGifilQB 0818 RPS CH 2 TRIP E

Q826 RPS CH 3 1 RIP L,

Q834 RPS CH 4 1 RIP T713 f RC AVG 1LMP (120/920) 1724 I RC Loop 1 CLG 1emp (0/700)

T733 f RC Loop 2 CLG Temp (0/700)

,1 T753 f RC Loop i HLG WR Temp, CH 1 1773 f RC PRZR PWR RLF Out Temp, RC12-1 1774 f RC PRZR Spray line Temp 1775 r RC PRZR Surge Line Temp I

1777 i RC PRZR Temp, RC15-2 1783 f RC Loop 2 HLG WR lemp, CH 2 1885 f SG 1 Out SIM Temp I-1901 F

SG 2 Out SIM lemp XO38 1-G Master lurb Trip Z768 Position RC PRZR PWR RLF VLV Z769 Position RC PRZR PWR Rif Shutoff VLV I

Z771 Position RC PRZR Spray line VLV, RC-10 L768 Inches Pressurizer level f738 GPH RC MU ILOW, low Range I

f740 GPM RC MU FLOW, High Range P725 Psig RC Loop 1 HLG WR Press, SFAS CH3 R790 Mwt Auctioneered N1 Linear Power

.I T781 f

RCP l-1 Oschrg CLG WR iemp, RC482 Z772 Position RC PRZR Spray Line VLV, RC2 2750 % Open RC MV CTRL VLV L768 Inches Pressurizer Level I

I 5

A-8

sus a

sus mes m

m m

M M

m m

M M

Figure A-1. Davis-Besse Data Acquisition Hardware Configuration INSIDE CONTAINMENT

_ OUTS!DE CONTAINMENT s.-

SURGE LINE p

~

CLAMP-ON BANDS REFERENCE g

M TCOMP.

[

\\

(

t/O CABINETS r

1 1

~

^

UPTO 8

()

/

O o

60 SIGNALS IN 6 CABLES STRING L/C HARDWARE 14 o.

[] --

(DATA ACO.)

f~'

CONTROL NEMA 12 ROOM i '. NNET

/

DATA CONTAINMENT

urum, m-<e ges_s v^

FIGURE A-2. Instrumentation Locations at Davis-Besse Unit 1 g

iL, t5 27TC 97TC r,0(

3 7TC 10 2 TC PZR u i i Tc i2 7 TC

}

$ 2TC M2~TC

[

l 6 7 TC 14 7 T/C h

C 14 i

3 ro l

ii~ ~ 10 mi t-L' HL i

6 6

3 s./

s i

_t I O

-- j ius,

f

= = s e -.) + p s*,

7

\\,

-,,a.~

w

-- ri r u wn4

- w Thermocouple Locations X: NORTH Y;UP VERTICAL I

1 PT Y.2 6 PT. X.2 N w PZR

. ri Y.2

n. ei. x Y

7. PT. K.Y,Z 13 Fi. X,Y

-'~

n

.__13 15 iS FT. Y g

t E'

- - 11

$$ --e HL 3

1 9_.}

__. -}?

-f,,,.

,., _-+ s

. L.E N GT H pg e

*4 Displacement Transducer Locations A 'O

[

Figure A-3 Thermocouple Positions L

l T1 gm _..s.

.g.

1 79"

~~ N T2 l

T3 1

8.75" w_

g-

]

T4 l

)

TS l

TG

/

l

'N.tf l

10.75" i

Note: Distance between adjacent thermocouples (T1-T2, T2-T3, etc.) is 1.79".

A-12

....

ML20085D147

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