ML20078J583

Nonproprietary Braidwood Unit 1 & Byron Unit 1 Model D4 SG Limited Tube Support Plate Displacement Analysis in Support of Interim Plugging Criteria

ML20078J583
Person / Time
Site:	Byron, Braidwood
Issue date:	11/30/1994
From:	WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19311B498	List: ML20078J552 ML20078J558 ML20078J577 ML20078J583
References
SG-94-11-005, SG-94-11-5, WCAP-14223, NUDOCS 9411210361
Download: ML20078J583 (200)

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Attahcment B WCAP-14223 (Non-Proprietary Version)  !

"Braidwood Unit 1 and Byron Unit 1 Model D4 Steam Generator Limited Tube Support Plate Displacement Analysis in Support of Interim Plugging Criteria" dated November 1994

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WESTINGHOUSE NON-PROPRIETARY CLASS 3.  ;

I WCAP-14223 SG-94-11-005 i

l BRAIDWOOD UNIT 1 'AND BYRON UNIT 1 [

MODEL D4 STEAM GENERATOR LIMITED i TUBE SUPPORT PLATE DISPLACEMENT ANALYSIS l IN SUPPORT OF INTERIM PLUGGING CRITERIA NOVEMBER 1994 f l

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WESTINGHOUSE ELECTRIC CORPORATION l NUCLEAR SERVICES DIVISION i P. O. BOX 158 MADISON, PENNSYLVANIA 15663-0158 .

C 1994 Westinghouse Electric Corporation i All Rights Reserved

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BRAIDWOOD UNIT 1 AND BYRON UNIT 1 MODEL D4 STEAM GENERATOR LIMITED TUBE SUPPORT PLATE DISPLACEhENT ANALYSIS IN SUPPORT OF INTERIM PLUGGING CRITERIA TABLE OF CONTENTS SECTION PAGE 1-1

1.0 INTRODUCTION

2-1 2.0

SUMMARY

AND CONCLUSIONS 2.1 Overall Conclusions 2-3 2.2 Summary MODEL D4 S/G DESIGN DESCRIPTION 3-1 3.0 3.1 Overall Design 3-1 3.2 Tube Support Plate Design 3-2 3.3 TSP Supports 3-3 3.4 Secondary System Considerations 3-3 4.0 THERMAL HYDRAULIC MODELING 4-1 4.1 TRANFLO Code Description 4-1 4.2 Model D4 TRANFLO Models 4-1  ;

4.3 Calculation of TSP Pressure Drop from Dynamic Analysis 4-2 4.4 Model D4 S/G Operating Conditions 4-3 4.5 TSP Pressure Drop Data 4-3 4.6 Acoustic Pressure Wave Considerations 4-4 4.7 Balance of Plant Modeling 4-5 5.0 5-1 QUALIFICATION OF TRANFLO CODE 5-1 5.1 Qualification Plan 5.2 Previously Reported Efforts 5-1 5.3 MULTIFLEX Code Description 5-4 5.4 MULTIFLEX Models 5-5 5.5 Comparison of MULTIFLEX and TRANFLO Results 5-6 5.6 Conclusions  : 5-7 5.7 References 5-7 6.0 THERMAL HYDRAULIC SLB LOADS ON TSPs 6-1 6.1 Analysis Plan 6-1 6.2 Reference Full Power and Hot Standby I. cads 6-3 6.3 SLB Load Dependence on Water Level 6-4 i

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TABLE OF CONTENTS (continued)

SECTION EADE 6.4 Best Estimate Loads 6-5 6.5 SLB Load Sensitivity Analyses 6-6 6.6 Acoustic Wave Considerations 6-8 6.7 Conclusions 6-10 7.0 STRUCTURAL MODELING FOR TSP DISPLACEMENTS 7-1 7.1 General Methodology 7-1 7.2 Component Materials 7-1 7.3 TSP Support System 7-1 7.4 Finite Element Model 7-3 7.5 Revised Material Properties 7-3 7.6 Dynamic Degrees of Freedom 7-5 7.7 Displacement Boundary Conditions 7-6 7.8 Integration Time Step / Structural Damping 7-6 7.9 Application of Pressure Loads 7-6 8.0 TSP DISPLACEMENT ANALYSIS RESULTS 8-1 8.1 Analysis Approach 8-1 8.2 Summary of Limiting Plate Displacements 8-2 8.3 SLB Displacements by Tube Locaticn 8-3 8.4 Summary of Stress Results 8-6 9.0

SUMMARY

OF BRAIDWOOD-1 AND BYRON-1 INSPECTION RESULTS 9-1 9.1 Braidwood Unit 1 1994 (EOC 4) Inspection Summary 9-1 9.2 Byron Unit 1 1994 (EOC 6) Inspection Summary 9-2 9.3 Indications at Regions of Significant TSP Displacement 9-3 9.4 Conclusions 9-3 10.0 ANALYSIS. METHODOLOGY FOR TUBE BURST WITH LIMITED 10-1 TSP DISPLACEMENTS 10.1 General Description of Analysis Methods 10-1 10.2 Burst Test Results for Cracks Extending Outside TSPs 10-1 10.3 Burst Probability as a Function ofThroughwall Crack Length 10-2 10.4 Durst Probability as a Function of Bobbin Voltage 10-4 10.5 Modeling for Burst Probability With Limited TSP Displacements 10-5 10.6 Conclusions 10-7 ii

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(continued)  :

SECTION PAGE 11.0 ' TUBE BURST PROBABILITIES FOR LIMITED TSP DISPLACEMENTS 11-1 l 11.1 General Approach to the Tube Burst Probability Assessment - 11-1 3 11.2 Deterministic Burst Margin Assessment Il-2 j 11.3 Bounding Tube Burst Probability with Limited TSP Displacements Il-3 '

11.4 Probability of Burst for Braidwood-l EOC4 TSP Indications 11-3 11.5 Probability of Burst for Byron-1 EOC6 TSP Indications Il-5 ,

11.6 Conclusions Il-6  !

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1.0 INTRODUCTION

This report describes a systematic series of analyses to demonstrate limited tube support plate (TSP) displacement in a steamline break (SLB) event, in support of alternate plugging criteria (APC) for the Model D4 steam generator (S/G). In particular, applicability to the Braidwood I and Byron 1 S/Gs has been confirmed through a review of the as-built drawings for these units. With limited SLB TSP displacement, constraint provided by the presence of the TSP substantially reduces the potential for a tube burst, even under the assumption of large outside diameter stress corrosion cracking (ODSCC). Tube burst probability c.nalyses are provided in this report to demonstrate that the limited displacements reduce the likelihood of a tube rupture to acceptably small values. The limited TSP displacemem analyses of this report replace the prior Braidwood I analyses given in WCAP-14046, Revision 1.

The evaluations of this report include hydraulic SLB analyses to obtain the time dependent pressure drop loads on the TSPs, structural analyses that apply the hydraulic loads to determine the tube location dependent TSP displacements, and analyses that apply the limited TSP displacements to conservative representations of the crack distributions to obtain the probability of a tube rupture.

The TRANFLO computer code (Section 4) is used to obtain the hydraulic pressure drop loads on the TSPs in a SLB event. To further support prior qualification efforts for the TRANFLO code, comparison analyses were performed with the MULTIFLEX computer code to assess the potential for acoustic wave effects to influence the hydraulic loads on the TSPs and to obtain independent analysis verification of the TRANFLO TSP loads. The MULTIFLEX code has been approved by the NRC for dynamic analyses of pipe breaks in reactor systems.

The MULTIFLEX results are described in Section 5 of this report.

An extensive series of TRANFLO analyses (Section 6) was performed to define reference loads for SLB events both inside and outside containment and for events at hot standby and full power conditions. In addition, a series of analyses was performed to assess the sensitivity of the SLB TSP loads due to the break size, S/G water level, TSP pressure drop loss coefficients, S/G downcomer loss coefficients and the Moody discharge coefficient.

To develop a conservative, bounding set of TSP loads including worst case uncertainties, the variables in the sensitivity analyses that increased the loads were combined in a TRANFLO analysis to develop a ratio of maximum to reference case loads. The upper limits of this ratio were then applied to the reference TSP loads to define the bounding set ofloads for TSP displacement analyses. The analyses then compare the reference TSP displacements and associated tube burst probabilities with that obtained from the bounding TSP loads.

The structural analyses for TSP displacements (Sections 7 and 8) utilize the WECAN code and analysis methods that have previously been reviewed and found acceptable by the NRC.

Displacement analyses are performed for the reference hot standby and full power events, as well as the bounding load conditions including worst case uncertainties. In addition, displacement analyses were performed using the MULTIFLEX loads for comparisons with displacements obtained from the TRANFLO loads.

T5P,01 WP5 1.I November 3,1994

Tube burst probability analyses (Section 10 describes methodology and Section 11 provides results) are performed for both hot standby and full power conditions, with TSP displacements obtained from both the reference and bounding TSP loads. A minimum tube burst probability is obtained by assuming that the TSP displacement at each tube location (all hot leg TSP intersections conservatively assumed to have a long throughwall crack) exposes a throughwall crack length equal to the displacement. The distribution of indications found in the Braidwood I and Byron 1 inspections (Section 9) are also applied to obtain a more likely estimate of the tube burst probability. The analysis results are used to demonstrate that acceptable burst probabilities will be obtained even iflarge ODSCC indications should occur, that APC structural and repair limits based on tube burst considerations are not required for Braidwood I and Byron 1, and that limited inspection intervals, such as mid-cycle inspections, are not required to obtain acceptable burst probabilities.

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TSP,01 WP5 1-2 November 8,1994 I

2.0

SUMMARY

AND CONCLUSIONS This report describes analyses performed to demonstrate limited TSP displacement analyses and associated low tube burst probabilities for a SLB event in support of alternate plugging cnteria applications at the Braidwood I and Byron 1 Model D4 S/Gs. The analyses performed include TRANFLO code analyses to obtain the pressure drop loads on the TSP in a SLB event, dynamic structural analyses to obtain TSP displacements as a function of tube location for the limiting TSPs, tube burst probability analyses (bounding estimate) assuming exposed crack lengths equal to the local tube TSP displacements and burst probability analyses for the actual TSP indications found at EOC 4 for Braidwood 1 and EOC 6 for Byron 1. TRANFLO sensitivity analyses were performed to develop a conservative uncertainty factor that is applied to the reference laods to obtain bounding loads on the TSPs.

For further verification of the TRANFLO code, analyses were performed with the MULTIFLEX code to assess the potential for acoustic wave effects, and for an independent analysis of the TSP loads.

2.1 Overall Conclusions Reference analyses were performed for hot standby and normal operating conditions using expected values for water level and pressure drop loss coefficients. Table 2-1 summarizes the TRANFLO analyses performed for this report. Cases 1 to 4 represent the reference analyses, which included SLBs for breaks at the S/G nozzle inside containment and for breaks outside of containment which could result in containment bypass radiation release. It is shown that breaks outside containment result in smaller or essentially equal loads to breaks inside containment such that containment bypass is less limiting than breaks inside containment. A best estimate analysis (Case 21) based on a limited break, rather than the guillotine break of the reference analyses, resulted in a significant reduction in the TSP loads to only 64% of the reference Case 1 results. The TRANFLO sensitivity results to assess the influence of key input parameters on the loads are also given in Table 2-1. The input parameters that increased the loads were then applied collectively in Cases 61 and 62 to determine very conservative uncertainty factors on the reference loads. Based on these results, an uncertainty adjustment factor on the TSP loads of 2.0 and 1.75 were defined to envelope the potential uncertainties in the reference analyses. These load factors were then applied to the time dependent loads of Cases 1 and 2 to develop Cases 11 and 12 as bounding loads for obtaining TSP displacements. Cases 11 and 12, with the reference loads adjusted by conservative uncertainty factors, result in upper bound TSP displacements and an associated bounding estimate for the tube burst probability with limited TSP displacements.

Analyses were also performed with the MULTIFLEX Code (Case 31) to assess the potential for acoustic waves to influence the TSP pressure drop loads. The results of these analyses show that acoustic waves do not effect the TSP loads and that the MULTIFLEX TSP loads provide an independent verification of the loads calculated with the TRANFLO Code. The TSP loads obtained with the MULTIFLEX code are slightly higher than obtained with the TRANFLO code but bounded by the uncertainty factor applied to the TRANFLO leads.

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Table 2-2 summarizes the SLB TSP displacements for the limiting TSPs 3 and 7 and the tube burst probabilities obtained for the reference analyses (Cases I and 2). TSP displacements for the reference analyses are s 0.35 inch and result in tube burst probabilities of 4.410" for a SLB at hot standby conditions and < 104 for a SLB at normal operating conditions. These results are based on the extremely conservative assumption that the TSP displacement at each hot leg intersection exposes a throughwall crack length equal to the TSP displacement. The bounding TSP loads of Cases 11 md 12 include the adjustment factor on the TSP loads that bounds uncertainties influencing the load analyses. The maximum TSP displacements, localized to the comers of TSPs 3 and 7, are 0.544 inch for the adjusted hot standby loads and 0.258 inch for the adjusted full power loads. The resulting tube burst probabilities for the 4

bounding estimates of Cases 11 and 12 were found to be 2.910 for hot standby conditions and < 10" for full power conditions. These results, like the reference analyses, very conservatively assume that the TSP displacements expose throughwall crack lengths equal to the TSP displacements. Even when the uncertainty adjustment factors are applied to the loads for a SLB at full power, the tube burst probability is extremely small due to the small TSP displacements (< 0.26 inch at worst tube location).

When distributions of indications found at Braidwood I at EOC 4 and Byron 1 at EOC 6 are applied, rather than assuming all hot leg intersections have throughwall indications, the 4

resulting tube burst probabilities for a SLB at hot standby conditions are 2.510 for Braidwood I and 1.1104for Byron 1 and < 10" for a SLB at normal operating conditions.

These results dramatically demonstrate the benefits oflimited TSP displacements as the free 4

span burst probabilities for these EOC distributions exceeded 10 . These results also reflect the expected range of tube burst probabilities with limited TSP displacements as they are based on actual indication distributions rather than the bounding assumption of a throughwall crack at every hot leg TSP intersection.

Hot standby conditions represent only about 4% of a typical fuel cycle (3.8% evaluated for Braidwood-1, average of 1.7% for the last three cycles at Byron-1) and operating conditions typical of full power represent the remaining 96% of the fuel cycle. As noted above, the tube burst probability is significant only for hot standby conditions. If the hot standby and full power tube burst probabilities of 2.9104and < 10"", respectively for Cases 11 and 12 with the uncertainty adjustment factor, are weighted by the operating condition frequencies, the resulting tube burst probability per operating cycle is only 1.210" even under the very conservative assumption of throughwall crack indications at all hot leg TSP intersections.

The TSP loads adjusted by the 2.0 and 1.75 uncertainty adjustment factors for hot sandby and full power operation represent loads that bound the analysis uncertainties in TRANFLO and the differences between codes such as the MULTIFLEX Code. These loads would be the appropriately conservative loads to apply for assessments of tube expansion at TSP intersections as a means of further reducing the SLB TSP displacements to negligible levels.

The results of this ieport demonstram that, even under worst case uncertainties, tube burst probabilities are xceptable based on t.se limited TSP displacements and that tube burst should not be limiting far the Braidwood I and Byron 1 S/Gs. Based on these low tube burst probabilities for the most conservative possible tube degradation assumption, it is concluded TSP,02 WPs 2-2 w ==6 e.i'94

that deterministic structural limits and low 1.0 volt tube repair limits to preclude tube burst are not required for the Braidwood I and Byron 1 APC applications. Even with the large voltage growth rates found for a few indications at the las' Braidwcod I and Byron 1 )

inspections, the S/Gs can operate to the planned, full cycle refueling outage and achieve low l EOC tube burst probabilities due to the limited SLB TSP displacements for these S/Gs.

2.2 Summary Hydraulic Loads o_ntthe TSPs in.a SLB Event The TRANFLO Code is used to obtain the hydraulic pressure drop loads on the TSPs in a SLB event. To further support prior qualification efforts for the TRANFLO Code, comparison analyses were performed with the MULTIFLEX Code to assess the potential for acoustic wave effects to influence the hydraulic loads on the TSPs. The MULTIFLEX Code has been approved by the NRC for dynamic analyses of pipe breaks in reactor systems.

MULTIFLEX solutions utilize the method of characteristics which is one of the best available solution methods for evaluating acoustic effects. The results of the MULTIFLEX analyses demonstrate that acoustic wave effects have no significant influence on the TSP loads from a guillotine SLB. In addition, the maximum pressure drops on the TSPs from MULTIFLEX are in acceptable agreement with the TRANFLO loads and the differences in TSP loads are bounded by the uncertainty factor applied to the reference TRANFLO loads.

Table 2-1 provides the matrix of TRANFLO analyses performed for the reference conditions and for the sensitivity ar.alyses used to define the bounding uncertainty adjustment. Cases I to 4 represent the reference analyses for hot standby and normal operating conditions. The TSP pressure drop ratios relative to Case 1 for hot standby conditions are also seen in the table. The hot standby conditions result in the largest loads on TSP 3 while the normal operating conditions result in the largest loads on TSP 7. Plates 3 and 7 are the limiting TSPs in that maximum TSP displacements are obtained for these plates. As shown in this report, only the TSP displacements at TSP 3 for hot standby conditions are significant enough to increase the tube burst probability to > 10-'

Cases 51 to 55 provide sensitivity analyses to break size, water level, TSP loss coefficients, downcomer loss coefficients and the Moody discharge coefficient. The changes for each of these cases represem the maximum range of uncertainty relative to the reference conditions.

It is seen that water level (21" decrease from the controlled setpoint) and TSP loss coefficients (change from expected or nominal to maximum loss coefficients) result in the largest increases in the TSP pressure drops relative to the reference conditions. A reduction in the downcomer loss coefficient was found to result in small (5%) increases in the pressure drops. These three variations that increase the TSP loads were then combined in Cases 61 and 62 to obtain worst case analyses. Based on these analyses,it was concluded that a facN of 2.00 bounds the uncertainties on the hot standby loads and a factor of 1.75 bounds uncertainties on the full power loads. The uncertainty adjustment factors were then applied to the loads in Cases 11 and 12 to provide bounding loads for TSP displacement analyses.

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Case 21 was performed to demonstrate the load reduction when best estimates are applied for 2

the analysis conditions including a limited SLB break size of 1.5 ft . It is seen that the best estimate analysis results in only about 64% or less of the reference TSP loads-Structural Analyses [gr SlJLTSP Disolacements The maximum SLB TSP displacements at any tube location are given in Table 2-2 together with the relative TSP pressure drop ratios from Table 2-1. For the reference analyses, the maximum TSP displacement of 0.350 inch (at the outermost tube position) occurs for TSP 3 at hot standby conditions. Conservatively assuming the 0.35 inch displacement exposes a ,

throughwall 0.35 inch crack, the associated tube burst probability for this postulated indication would be about 1.710-", which is negligibly small. For the reference hot standby conditions, only about 432 and 26 TSP 3 intersections have displacements > 0.1 and 0.30 inch, respectively, since only the comers of the TSP near the tube lane have significant displacements. For a SLB at full power conditions, the maximum TSP displacement occurs at the corner of TSP 3 and is limited ta a negligible 0.094 inch. When the MULTIFLEX loads are applied, the ma;cimum TSP displacement for a hot standby SLB is only 0.421 inch, which corresponds to a tube burst probability for one indication of only about 410 Thus differences between TRANFLO and MULTIFLEX are negligible for the TSP loads. The MULTIFLEX Code analyses support the absence of acoustic effects and independently support the TRANFLO Code analyses for the hydraulic loads on the TSPs in a SLB event.

When the reference hot standby loads are adjusted by the 2.00 uncertainty adjustment factor, the maximum TSP displacement at TSP 3 is increased from 0.350 inches to 0.544 inches (Case 11 of Table 2-2). Assuming a single throughwall indication is exposed by the 0.544 inch displacement, the burst probability for this indication would be about 1.610" This is significantly higher than the expected reference analysis (Case 1) burst probability of 1.710 "

! for the associated 0.35 inch displacement. Thus, only when the bounding TSP displacements are obtained by applying the conservative 2.0 load adjustment factor are the tube burst probabilities significant. Even when the 1.75 uncertainty adjustment factor is applied to the l

reference full power operation loads of Case 2, the maximum TSP displacement is a l negligible 0.258 inch. Therefore, full power operation does not result in significant TSP displacements or tube burst probabilities even when bounding uncertainties are applied.

Tube Burst Probability Analyses for Limited TSP Disolacement Tube burst probabilities calculated for the reference SLB analysis conditions, for the reference analysis loads adjusted by the uncertainty factor, and for the MULTIFLEX analyses are given in Table 2-2. For the reference hot standby conditions, the tube burst probability, assuming a throughwall crack at every hot leg TSP inte.rsection is exposed to become a free span crack the length of the local TSP displacement, is an extremely small 4.410-' The reference analysis for a SLB at full power conditions results in an even smaller tube burst probability.

When the conservative uncertainty adjustment factors of Cases 11 and 12 are applied with the assumed throughwall cracks at each TSP intersection, the burst probability for hot standby conditions is still small at 2.9108and for full power remains < 10-' Given the 73P,02 WPs 24 14evember 9,1994 t

conservatisms in the loads and the throughwall crack assumptions, these burst probabilities are negligibly small.

To obtain a more realistic estimate of the tube burst probability, a distribution of indications at the TSP intersections is required. This can be obtained by applying the actual distributions found at EOC 4 for Braidwood 1 and EOC 6 for Byron 1. Based on the APC analyses of tube burst probabilities for these EOC distributions, the resulting free span tube burst probabilities were greater than 10 2 Applying the limited TSP displacements to these EOC distribution results in burst probabilities of 2.4104 for Braidwood 1 EOC 4 and 1.110 5 for Byron 1 EOC 6. The application of the actual indication distributions reduces the burst probability by more than two orders of magnitude (from 2.910'5 to 1.110~5) compared to the bounding assumption of throughwall indications at all hot leg TSP intersections. The benefits oflimited TSP displacement are clearly seen by the reduction of about three orders of magnitude (> 10 2 to 1.110'5) in the tube burst probabilities compared to the free span APC analyses.

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Table 2-1. TRANFLO Hydraulic Analysis Matrix and Summary Results Case No. Analysis Conditions Summary Results E Break Brea Water TSP D C. Moody Un cer. Relative TSP AP(*)

Operat,ing Loc. k Level Loss Loss Disch. Adj.

Conditton Size Coeff Coeff. Coeff. Factor 3 (C) 7 (J)

Reference Analyses ,

1.110t standby S/G(') Guil. 487" Nom. Nom. 1.00 -

1.00 1.00

2. Full Power - 0.25 1.77 1.00(8) 1.00(')

3.110t standby O.C.(') - 0.87 0.82

4. Full Power - 1.00(') 1.03(')

Reference Analyses with Uncertainty Adjustment i1. Ilot standby S/G Guil. 487" Nom. Nom. 1.0 2.00"> 2.00 2.00

12. Full Power 1.75") 1.75(') 1.75(') .

Best Estimate Analyses

21. Ilot standby O.C. 1.5 ft' 487" Nom. Nom. 0.84 -

0.36 -0.64 TRANFLO Qualification Analyses (MULTIFLEX Code)

31. Ilot standby S/G Guil. 487" Nom. Nom. 1.0 - 1.38 1.69 TRANFLO Sensitivity Analyses ,

51.110t standby S/G 1.5 fl' 487" Nom. Nom. 1.00 - 0.92 0.97

52. liot standby Guil. 466" Nom. Nom. 1.00 - 1.25 1.33

53. Ilot standby 487" Max. Nom. - 1.29 1.28 54.110t standby Nom. Min. - 1.02 1.05 55.11ct standby Nom. 0.84 - 0.93 1.03

61. liot Standby S/G Guil. 466" Max. Min. 1.00 - 1.93 2.13

62. Full Power -

1.60(') 1.64(')

Notes: 1. For hot standby, ratios of mammum pressure drops for each case are given relanve to Case 1. For full power, ratios are relative to Case 2.

2. s/O break location is at cut of 90 nonje. O C. break location is outside containment penetration.

3. For full power cases 4,12, and 62, maximum pressure drop ratios are given relative to Case 2.

4. Uncertamty adjustment factor applied to all TRANFLO pressure drop time histories from Cases 1 and 2.

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• Table 2-2. SLB Displacement and Burst Proti 51ity Summary Results Relative TSP Max, TSP Displ. at Tube Burst Probability Case Operating Condition AP(" Tube Loc. (inch) 3 (C) 7 (J) 3 (C) 7 (J) Bounding Braidwood 1 Byron 1 Estim ate ( Actual EOC 4(" Actual EOC 6")

Reference Analyses 1 Ilot Standby - SLB et S/G Nozzle 1.00 1.00 0.350 < 0.120 4.4 10- < 10~ < 10-

2 Full Pown - SLB at S/G Nozzle 0.20 1.77 0.094 < 0.052 < 10- < 10- < 10

l.00") 1.00")

3 llot Standby - SLB Outside Containment 0.87 0.82 - - - - -

4 Full Power - SLB Outside Containment 1.00"' l.03") - - - - -

u Reference Analyses with Uncertainty Adjustment 4

11 liet Standby - SLB at S/G Nozzle 2.00(" 2.00"' O.544 0.268 2.9 10-5 2.4 10 1.1 10-5 12 Full Power - SLB st S/G Nozzle 1.75") 1.75") 0.203 0.258 < 10 < 10- < 10-'*

TRANFLO Qualification Analyses (MULTIFLEX Code) 4 31 Ilot Standby - SLD at 50 Nozzle 1.38 1.69 0.431 < 0.207 1.7 10 1.1 10' 7.0 10' Notes: 1. For hot standby, ratios of maximum pressure drops for each case are given relative to Case 1. For full power, ratios are relative to Case 2.

2. Very conservative bounding estimate based on assuming a throughwell crack at every hot leg TSP intersection is exposed as a free span crack oflength equal to the TSP SLB displacement at the tube location.

3. Burst probability calculated for actual EOC tube locations with TSP indications. Calculated as lower burst probability of free span burst from burst pressure versus voltage correlation or assumption of a throughwall crack length equal to the SLB TSP displacement at the tube location.

4. For full power cases 4 and 12, maximum pressure drop ratios are given relative to Case 2.

5. Uncertainty adjustment factor applied to all TRANFLO pressure drop time histories from Cases I and 2.

TSP _e2 WPS 2-7 "-.'a

3.0 MODEL D4 S/G DESIGN DESCRIPTION 3.1 Overall Design The Byron Unit I and Braidwood Unit I steam generators are of the Westinghouse Model D4 preheat steam generator design. Each steam generator (S/G) contains 4578 mill-annealed Alloy 600 U-tubes,0.75 inch OD x 0.043 inch wall, which provide 48,300 sq. ft. of heat transfer area per S/G. Figure 3-1 shows the steam generator layout; a detailed layout of the preheater region is shown in Figure 3-2. Primary coolant enters the hot leg channelhead and passes through the U-tubes, which transfer heat from the primary side to water on the secondary side, which is converted to steam. On the secondary side, about 10% of the feedwater flow at full power is bypassed to an auxiliary nozzle to enter the upper plenum in the region of the primary moisture separators. Most (about 90%) of the feedwater enters the S/G through the preheater inlet nozzle into the preheater region (see Figure 3-2). The feedwater flow entering the preheater nozzle is directed to the bottom of the preheater by a waterbox at the nozzle, from where it circulates upward through a series of preheater baffle ,

plates which discharge the flow upward into the tube bundle. As the secondary fluid passes through the tube bundle, it is converted to a water / steam mixture which passes upward through the transition cone region of the S/G shell, into the primary and secondary moisture separators in the upper shell region. Water is separated from the steam before the dry steam exits the S/G via the steam outlet nozzle. Water removed by the moisture separators flows down the annulus between the shell and the wrapper surrounding the tube bundle region.

Upon reaching the tubesheet, the water is once again directed upward through the flow distribution baffle into the tube bundle. A partilbn plate between plates B and L (see Figure 3-2) separates the preheater and hot leg sides t,f 'he S/G. Below plate B and above plate L, secondary flow can cross between the hot and cold leg sides of the S/G.

The S/G tubes pass through tube support plates (TSPs) which provide lateral support to the tubes and contain circulation holes through which the water / steam passes upward through the tube bundle. On the cold leg side in the preheater region, these support plates contain no circulation holes, and act to direct the flow across the tubes; therefore, these plates are also referred to as baffle plates. The flow distribution baffle, at an elevation of 6.0 inches above the top of the tubesheet, distributes flow across the tubesheet and upward through a cutout in the plate on the hot leg side.

During normal operation, a slight pressure drop exists across each TSP or baffle plate. This pressure drop causes small displacement of the TSPs relative to the tubes during normal operating conditions. At hot standby conditions, there is no secondary flow or pressure drop across the TSPs. However, during postulated accident conditions such as steam line break (SLB), pressure differentials across individual TSPs can act to displace unsupported regions of the TSPs in such a manner as to uncover degradation within the TSP crevice. The following sections provide specific design information concerning the Model D4 baffle and support 1

plates.

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3.2 Tube Support Plate Design The Model D4 steam generators at Byron 1 and Braidwood I utilize 0.75 inch thick carbon steel support plates with drilled (round) tube holes set on a square pitch of 1.0625 inches.  !

With the exception of the flow distribution baffle and preheater baffle plates (described below), the tube support plates also include flow circulation holes measuring 0.50 inch in diameter, set on a square pitch of 1.0625 inches within the tube hole array.

The tube support plates of the Byron 1 and Braidwood 1 S/Gs may be classified as one of three types: the flow distribution baffle, preheater baffle plates, and tube suppon plates. The flow distribution baffle (FDB), located 6.0 inches above the top of the tubesheet, is comprised of two halves, with the cold leg side containing no circulation holes or cutouts; a moon-shaped cutout on the hot leg side permits the secondary fluid to pass upward through the tube bundle. On the cold leg side of the FDB, the drilled tube holes measure 0.900" in diameter, compared to the 0.750" OD of the tube. Therefore, the FDB provides no lateral support for the tubes. On the hot leg side, the drilled tube holes inside a radius of 32" from center of the S/G measure 0.875" in diameter, and the tube holes measure 0.833" in diameter outside a 32" radius from the center of the S/G. Hence, the enlarged FDB holes allow some secondary fluid to pass upward through the tube /FDB crevices, but no lateral support is provided for the tubes at the FDB level due to the large tube to FDB clearances.

The preheater baffle plates, shown in Figure 3-3, contain 0.766" diameter drilled tube holes and no circulation holes. Their function is to provide lateral. support to the tubes and direct the flow back and forth across the tubes as the feedwater passes upward through the preheater. Letter designations are used by Westinghouse for the baffle plates at various elevations, with "A" representing the FDB and "B", "D", "E", "G", and "H" representing preheater baffle plates with no circulation holes.

On the hot leg side, two semi-circular plates ("C" and "F") with 0.766" diameter tube holes as well as 0.50" diameter circulation holes are located at the elevations of the "D" and "G" preheater baffle plates. These plates permit flow upward through the tube bundle and provide lateral support for the tubes. Plates "J" and "K", on the hot and cold leg side of the S/G at the top of the preheater, similarly contain 0.766" diameter tube holes and 0.50" diameter circulation holes. The remainder of the tube support plates, "L", "M", "N" and "P" are full ,

size circular plates with similar tube and circulation holes. In addition, the "L" through "P" plates contain central flow slots along the tube lane to enhance flow upward through the bundle.

At Braidwood I and Byron 1, number designations are used for the plates, counting upward from the FDB through the preheater to the top TSP. The correspondence with the Westinghouse letter designations are: 1 = A, 2 = B, 3 = C, 4 = E, 5 = F and G, 6 = H, 7 = J and K, 8 = L, 9 = M,10 = N and 11 = P.

T5P_03 WP5 3-2 Nav==bar 9.1994 l

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3.3 TSP Supports The FDB, TSPs and preheater baffle plates are supported vertically using several support mechanisms, including five tierods/ spacers in each half of the tube bundle. Preheater baffle plates C (3H), F (SH), and J (7H) are supported at their center by a vertical bar welded to the partition plate, while all of the TSPs above the preheater are supported at their center by a central tierod and spacer. Each of the baffle plates and TSPs are supported at the edges by vertical bars welded to the wrapper and/or partition plate immediately above the plates.

In-plane supports are provided by wedges located around the circumference of each plate.

The wedges are welded to the wrapper; their tapered design provides additional resistanc: to upward movement, in addition to in-plane support, due to the sloped face of the wedge.

Section 7.3 provides a detailed description of the Model D4 S/G TSP support system.

Figure 7-1 shows a schematic of the tube bundle region, and support locations for each of the plates are shown in Figures 7-2 to 7-10. Detailed descriptions of the support components, including the tierods, spacer bars, and wedges groups, are provided in Section 7.3.

3.4 Secondary System Considerations The steam generator secondary side consists of a natural circulation loop with feedwater inlets and a steam outlet. Most (~90%) of the feedwater enters the preheater of the generator through main feedwater nozzle. The feedwater then flows through four crossflow passes; it moves upward and leaves the preheater to join the flow from the hot leg side, after passing upward through the tube support plate L. A fraction of the feedwater flow comes down through the bottom baffle to meet with the flow from downcomer. The resulting flow then moves into the hot leg side via the tube lane.

A small fraction (-10%) of the feedwater enters the steam generator through an auxiliary feedwater nozzle in the upper shell. The feedwater from the auxiliary nozzle mixes with the separated water from the moisture separators. This takes place in the upper water reservoir.

The mixed water flows down the downcomer annulus, which is separated from the tube handle by the wrapper. The downcomer flow enters the tube bundle through the wrapper opening above the tubesheet. As the fluid approaches the first tube support plate in the hot leg side, axial flow becomes dominant. Boiling takes place and the flow moves upward along the hot leg side.

The hot and cold leg tube bundle flow meet above TSP L. The combined flow moves upward while boiling continues, leaving the tube bundle and entering the primary separators.

A large portion of the water is separated by the primary separators and retumed to the water reservoir. The steam with the remaining entrained moisture then enters the secondary separators. This entrained moisture is trapped by a system of hook and pocket vanes and retumed to the water reservoir. The steam then leaves the steam generator through the steam outlet nozzle.

T5P_03 WPS 3-3 November 9.1994

The Model D4 S/Gs utilize a venturi type flow limiter in the steam outlet nozzle. The venturi flow area at the throat is about 1.4 ft2 , while the steam line flow area is about 4.7 ft',

therefore, the critical discharge flow is controlled by the flow limiter throat area of 1.4 ft when a guillotine steam line break is postulated.

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Figure 3-3. Flow Distribution and Preheater Baffle Plates j November 9,1994 {

T5P_03 WP5 3-7 l 1

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r 4.0 THERMAL INDRAULIC MODELING A postulated steam line break (SLB) event results in EMwdown of steam and water. The fluid blowdown leads to depressurization of the seconc. y side. Pressure drops develop and exert hydraulic loads on the tube support plates (TSPs) baffle plates. These hydraulic loads were determined for the Model D4 steam generatc; ming the TRANFLO computer code.

4.1 TRANFLO Code Description The TRANFLO code uses an elemental control volume approach to calculate the thermal and hydraulic characteristics of a steam and water system undergoing rapid changes. Fluid conditions may be subcooled, two-phase or superheated. The code considers fluid flow as one-dimensional. It predicts the mass flow rate, pressure, pressure drop, fluid temperature, steam quality and void fraction.

Control volumes simulate the geometrical model, and flow connectors allow mass and energy exchange between control volumes. Each nodal volume has mass and energy that are uniform throughout the volume. Flow connectors account for flow and pressure drops. The system model allows for flow entering or leaving any control volume. This then allows that feedwater flows into the steam generator and steam flows out ofit. The system models also permit a heat source, which then can simulate the tube bundle with hot water flow.

TRANFLO solves for system conditions by satisfying mass, momentum and energy equations for all control volumes. It models the effects of two-phase flows on pressure losses. The code allows a variety of heat transfer correla6ons for the tube bundle. It covers all regimes from forced convection to subcooled liquid through boiling and forced convection to steam.

4.2 Model D4 TRANFLO Models The TRANFLO computer model for Model D4 steam generator is composed of a network of nodes and connectors that represent the secondary side fluid, tube metal heat transfer and primary coolant. Figures 4-1 and 4-2 show the nodal layout of the secondary side of the Model D4 steam generator. Figures 4-3 and 4-4 present the nodal network of the secondary fluid, primary fluid and tube metal. The computational model consists of the following elements:

1. 31 nodes (i.e., No's. 22 through 52) for secondary fluid.

2. 44 fluid connectors (i.e., No's. 23 through 66) for secondary fluid.

3. 21 nodes (i.e., No's. I through 21) for primary coolant.

4. 22 fluid connectors (i.e., No's. I through 22) for primary coolant.

5. 21 heat transfer nodes (i.e., No's. I through 21) for tubes.

6. 42 heat transfer connectors (i.e., No's. I through 42) from primary to secondary fluid.

4-1

For a postulated SLB event, the above model considers a break just outside the steam outlet nozzle. However, the break may be further downstream, outside the containment building.

Figure 4-5 illustrates such a model; it has three additional nodes and three more flow connectors to simulate the steamline from the steam nozzle to the containment. For the  :

Braidwood-l and Byron-1 plants, this length is about 120 feet with about three elbow turns.

The TRANFLO modeling reported in WCAP-14046, Rev.1 involved a coding error in flow connectors. Flow connector No. 22 is in the primary coolant side; its upstream node is No. 21 and its downstream node is No. 0 (an outside node, see Figure 4-3). Instead of using outside node No. O, node No. 22 was coded in the input. Note that node No. 22 is a secondary side node (see Figure 4-4). This error is corrected in this report. Results for the i hot standby case are similar between this report and WCAP-14046, Rev.l. However, there are significant differences between this report and WCAP-14046, Rev. I for the full power case. The error in the prior full power analyses leads to an overestimate of the TSP loads in WCAP-14046. The correction results in proper simulation of full power behavior. The results show that hot standby yields higher loads than full power, as expected since the water flashing due to depressurization is greater for hot standby than full power conditions.

4.3 Calculation of TSP Pressure Drop from Dynamic Analysis In the tube bundle area, the space between support plates or baffles forms a fluid node, and a flow connector links the adjacent nodes (see Figures 4-1 and 4-2). Pressure drops through support plates or baffles are calculated by the code for each flow connector, which includes a plate or baffle. The whole length of the flow connector is very long compared to the thin plate of less than an inch. For example, flow connector 39 links node 30 to node 29. Tube support plate P is the boundary between node 33 and node 32. Node 29 is the U-bend and the space below the inlet of riser barrels. The pressure drop through TSP P is calculated along flow connector 39.

For example, the pressure difference between the centroids of nodes 30 and 29 acts to accelerate the flow for the whole connector 39, including the inertia of all the fluid in the flow path. It would be highly unrealistic to apply the overall pressure difference between nodes or along the whole connector to the thin plate. Note that connector 39 has a length of 66.6 inches and that TSP P is only 0.75 inch thick. A correct approach is to apply only the form pressure drop to TSP P, since the fluid itself absorbs most of the unsteady pressure drop through inertia, and the momentum flux and friction terms are distributed through the fluid.

It em be shown that the force on an orifice plate (or a TSP) resulting from the transient, blowdown-type flow of a compressible fluid in a pipe can be calculated from a form loss.

The orifice pressure drop is equal to a loss coefficient times the fluid dynamic head. The TRANFLO code is a proven code for dynamic analysis of two-phase blowdown flow resulting i from a SLB event. The hydraulic loads on the TSPs as calculated by the TRANFLO code or l any other altemate code are based on the form loss of pressure through the TSP.

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l 4.4 Model D4 S/G Operating Conditions Both Braidwood Unit I and Byron Unit I are plants with Model D4 steam generators, hence, they have identical designs for the tube bundle and tube support plates.

Operating conditions at full power for the Braidwood Unit 1 S/Gs are as follows:

Thermal Power = 856.25 Megawatts Thermal Design Flow = 94,400 gpm Primary Inlet Temperature = 618.4 F Primary Operating Pressure = 2250 psia Feedwater Temperature = 440'F Steam Pressure = 997 psia Circulation Ratio = 2.35 Byron Unit I has similar operating conditions to those above, with the exception of the primary side temperature and secondary steam pressure. The plant is licensed to operate at a reduced primary temperature, as low as 600 F, which results in a lower steam pressure (824 psia).

In addition, both plants have identical hot standby conditions. The steam temperature at hot standby is 557 F. Both plants operate at the same water level. The steam piping layouts are similar between the two pl ants.

In summary, both plants have essentially the same geometrical and thermal-hydraulic conditions. The TRANFLO model developed for the Braidwood Unit 1 S/Gs can also be used for Byron Unit 1. The TRANFLO calculations were made at the Braidwood-l inlet temperature of 618.48'F. Differences in the dynamic transient due to a SLB event are considered insignificant between an initiation of 618.48 F or 600*F.

4.5 TSP Pressure Drop Data Laboratory tests were made to correlate the loss coefficient through a tube support plate (see Figure 5-2 in Section 5.2.4). As discussed in more detail in Section 5.2.4, Figure 5-2 shows the test data and correlation of the loss coefficient for determining the pressure drop through a TSP. The correlation constant ranges from 0.8 to71.4, and its best estimate is 1.1. This correlation has been incorporated in the GEN code, a steam generator performance code.

The circulation ratio depends on pressure drops through the circulation loop. The circulation loop consists of the downcomer, tube bundle and primary separator. The downcomer has a small pressure drop. The major pressure drops come from various TSPs and swirivanes of the primary separator.

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l 4.6 Acoustic Pressure Wave Considerations Acoustic pressure mn affect the fluid flow through the tube bundle and thus the pressure drop through the TSP if it is significant inside the steam generator, in particular, in the tube bundle. The steamline; break can generate a depressurization wave into the steam generator.

Because of hardware elements inside the steam generator, this decompression wave will reflect and not be transmitted into the tube bundle. The nature of tortuous paths from the steam nozzle to the tube bundle (see Figure 4-1) weakens the penetration of the wave into the tube bundle. In addition, since the secondary side is two-phase, there is ample l

compressibility of the mixture and the speed of sound is significantly reduced. The secondary side will respond to the depressurization by flashing of the two-phasc mixture to a higher void fraction. The pressure drops that will occur will be due to mainly the two-phase flow acceleration, friction and form loss of hydraulic motion.

If the steam generator is in a hot standby condition when the break occurs, the vapor space above the water level would provide compressibility for the flow. Also, as the secondary side begins to depressurize, vapor forms within the hot liquid due to flashing, which will also ,

provide additional compressibility to the mixture. )

The TRANFLO code uses the complete transient mass, momentum and energy conservation equations; the acoustic propagation of pressure waves through steam or water occurs naturally within the solution obtained from these equations. Results obtained from the TRANFLO code have been tested against analytical results of acoustic phenomena. Their comparison has j demonstrated that the TRANFLO code has the capability to simulate the acoustic effect. Its j simulation for a steamline break depressurization would be able to calculate the effect of a depressurization wave initiated at the steam nozzle, if not negligible. A smaller time increment would be appropriate to properly simulate the acoustic effect.

However, a finer nocalization than that shown in Figures 4-1 and 4-2 may be needed to adequately analyze the acoustic effect by the TRANFLO code. The MULTIFLEX code retains the same conservation equations and it uses the method of characteristics with an explicit numerical solution scheme. The method and numerical scheme thus minimize numerical diffusion. The effect of numerical diffusion is to stretch the wave out spatially.

For relatively thin structures like the tube support plate, the effect of numerical diffusion can be significant because the pressure difference across the plate due to an acoustic wave is underestimated. Therefore, MULTIFLEX provides the most accurate spatial representation of acoustic waves. In terms of two-phase flow modeling, the TRANFLO code uses the drift-flux model to consider the effect of flow slip between water and steam phases. The MULTIFLEX l code uses a homogeneous model without flow slip between phases.

MULTIFLEX has been used (Section 5) to simulate the SLB event with equivalent modeling  ;

to the TRANFLO model. Results are compared with those of the TRANFLO model, and the effect of acoustic waves are assessed. As noted in Section 5, the effects of acoustic waves are shown to be negligible, as expected, based on the above discussion.

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4.7 Balance of Plant Modeling A postulated steam line break could take place in any operating mode of the plant; it could occur at any location along the steam line, and the size of the break could be double ended (i.e., guillotine) or limited. These variables all lead to different blowdown flot rates and two-phase motion inside the steam generator. Therefore, they all result in different hydraulic .

loads on the tube support plates. For example, a dynamic transient initiated from full power 1

operation is different from that initiated from hot standby. A guillotine break will have higher blowdown flow than a limited break.

A complete assessment of these variables requires a parametric study. This assessment is given in Section 6; it covers reference cases, best estimates and a sensitivity study for the above variables in plant conditions, as well as for the TSP pressure drop.

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Figure 4-1. Secondary side nodes and tube support plates - Identification for Model D4 S/G (see Figure 4-2 for Preheater Detail) 4-6 ,

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a Figure 4-2. Preheater nodes and baffle identification for Model D4 S/G 4-7 ,

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a Figure 4-3. Primary fluid nodes and flow connectors, metal heat nodes and heat transfer connectors, and secondary fluid nodes within tube bundle (Model D4 S/G).

4-8

a Figure 4-4. Secondary side fluid nodes and flow connectors for Model D4 S/G -

l Model for SLB just outside steam outlet nozzle.

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Figure 4-5. Secondary side fluid nodes and flow connectors for Model D4 S/G -

Model for SLB just outside containment building.

4 - 10

5.0 QUALIFICATION OF TRANFLO CODE In the early 1970's, there was a need to accurately predict the steam generator behavior under transient conditions, such as a steam line break (SLB) event; a transient can develop thermal hydraulic loads on the internal components and shell of the steam generator. Structural analyses are required to analyze the adequacy of the individual components and the whole steam generator under various thermal and hydraulic loads. With the assistance of MPR Associates, Westinghouse developed and verified the TRANFLO computer code to conservatively model the thermal and hydraulic conditions within the steam generator under transient conditions.

Qualification of the TRANFLO code has been a continuing process in demonstrating the code's capability to accurately predict the hydraulic loads on intemal components of the steam generator.

5.1 Qualification Plan For application to Braidwood Unit I and Byron Unit 1, the qualification effort consists of two parts. Part 1 (Section 5.2) describes the historical verification effort. Part 2 (Sections 5.3 to 5.6) documents a specific effort for both units; this includes an evaluation of the acoustic effect on hydraulic loads on tube support plates. For Part 2, a verified computer program called MULTIFLEX has been applied as described below, together with analysis results and comparisons with TRANFLO code results.

5.2 Previously Reported Efforts The secondary side of the steam generator involves water boiling under high pressure during normal operating conditions. During a transient such as a SLB event, it may be subject to vapor generation due to rapid depressurization. Therefore, analysis methods have to recogr'ize this characteristic of two-phase fluid behavior. In the early stage of the computer code development and technology of two-phase flow, a homogeneous model was used. For current analyses, a more accurate slip flow model is used which takes into consideration the relative velocity between the liquid and vapor phases. Development of the TRANFLO code reflects this general trend of the two-phase flow modeling. The first version of TRAhTLO was a homogeneous model, and it was later updated to a drift flux model to simulate the effect of two-phase slip. Since the original issue of the code, Westinghouse has made several enhancements to the code and has performed the appropriate verification and validation of these changes. j 5.2.1 Acceptability of Application of TRANFLO l The original version of the TRANFLO code (Reference 5-1) was reviewed and approved by the NRC in Reference 5-2. TRANFLO was used as part of the Westinghouse mass and energy release / containment analysis methodology. Specifically, the code was used to predict 5-1

=

steam generator secondary side behavior following a spectrum of steam line breaks. Its output was the prediction of the quality of the steam at the break as a function of time. The quality is calculated as a function of power level, as well as break size. In order to assure that the TRANFLO code evaluates a conservatively high exit quality, Reference 5-2 states that the calculational sequences were reviewed for the determination of conditions prior to entering into the separation stages. The calculated rate, quality and energy content of the two-phase mixture entering the separation stages must be evaluated conservatively. This review was completed and found to be acceptable, as the NRC staff concludes in Reference 5-2 that the TRANFLO code is an acceptable code for calculating mass and energy release data following a postulated SLB. Therefore, it is concluded that the TRANFLO model is appropriate for predicting S/G behavior (including tube bundle region) under the range of SLB conditions. In particular, the NRC review concluded that the flow rate and quality entering the separation stages is adequately conservative. Therefore, if further review is required, the review would focus on the flow distribution within the tube bundle. The distribution within the tube bundle is principally influenced by the TSP loss coefficients, which are based on experimental data for Westinghouse manufactured S/Gs (see Section 5.2.3). To further address the flow distribution within the tube bundle, additional sensitivity analyses in TRANFLO pressure drop dependence on the TSP loss coefficients have been performed as described in Section 6.

For the current application, TRANFLO is used in conjunction with a structural analysis code to predict TSP movement following the same SLB event. The key data transferred between the transient code and the structural code is the pressure drop across the TSP as a function of time. This pressure drop calculation depends on the fluid conditions in the steam generator and on the adequacy of the loss coefficients along the flow paths. The conditions in the tube bundle as calculated by TRANFLO have been previously reviewed. Further justification of the adequacy of the pressure drop calculation is discussed in Section 5.2.3.

5.2.2 Different Versions'of TRANFLO The original version of the TRANFLO code has been reviewed and approved by the NRC.

Westinghouse has continued to update the code with new models that more accurately predict steam generator behavior. Four versions of the TRANFLO have been used in calculations.

The following are descriptions of each of them.

The Original Version (April 1974) ,

This is the original homogeneous model, which MPR Associates developed in April 1974.

The code predicts mass flow rate, pressure, pressure drop, fluid temperature, steam quality and void fraction. The code document includes results of TRANFLO calculations for a 51 Series steam generator subject to water and steam blowdown due to a SLB event. The document also presents code verification using blowdown test data from pressurized vessels.

Westinghouse documented this version in detail in September 1976, including code verification using vessel blowdown data (Reference 5-1). Sensitivity analyses were also performed and documented to show that the modelling was conservative.

5-2

The TRANFLO code uses an elemental control volume approach to calculate the thermal-hydraulics of a steam and water system undergoing rapid changes. Fluid conditions may be subcooled, two-phase or superheated. The code considers fluid flow being one-dimensional.

Control volumes simulate the geometrical model, and flow connectors allow mass and energy exchange between control volumes. Each nodal volume has mass and energy that are homogeneous throughout the volume. Flow connectors account for flow and pressurc drops.

The system model allows flow entering or leaving any control volume. This then allows that feedwater flows into a steam generator and steam flows out ofit. The system models also permit a heat source, which then can simulate the tube bundle with hot water flow.

TRANFLO solves for system conditions by satisfying mass, momentum and energy equations for all control volumes. It models the effects of two-phase flows on pressure losses. The code allows a variety of heat transfer correlations for the tube bundle. It covers all regimes from forced convection to subcooled liquid through boiling and forced convection to steam.

The Drift-Flux Version (November 1980)

This version implements a drift-flux model to better simulate relative flow velocity between I

water and steam. For example, it allows a realistic simulation of counter current flow of steam and water. It required modification of the mass, momentum and energy equations of the two-phase flow. A capability is provided for monitoring calculated variables for convenient examination of results.

TRANFLO Version 1.0 (November 1991)

This version accepts transient data of parameters as direct inputs, rather than supplying input subroutines, as used in the' drift-flux version. It also improves printouts and plots. This version maintains the drift-flux model, and includes the addition of thermal conductivity of Alloy 690 tubing.

TRANFLO Version 2.0 (January 1993)

This version provides an option for two inlets of feedwater flow into the steam generator. It involves minor changes to a subroutine for specifying feedwater flow. This version is used for separate inlets of simultaneous feedwater flow from the main and auxiliary feedwater nozzles.

5.2.3 Verification of Loop Pressure Drop Correlations As discu . sed earlier, an accurate prediction of mass and energy release from the vessel means that the TRANFLO code properly calculates local thermal-hydraulics in various nodes (i.e.,

elemenal control volume and flow connector). It is critical to accurately simulate the pressum drop inside a steam generator that consists of various components, such as the tube bundle .zith tube support plates, moisture separators, and downcomer. Hydraulic loads on l

5-3

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various components depend on accurate pressure drop calculations. Thus, it is important to verify the pressure drop calculations through the circulation loop.

The TRANFLO code uses the same pressure drop correlations as the Westinghouse GENF code, which is a performance program. The GENF code predicts one-dimensional steady state conditions, which include pressure drops along the circulation flow loop. Both laboratory tests and field data validate the accuracy of the GENF code. The GENF code is used extensively for steam generator performance analysis and has been shown to accurately 1 predict operating steam generator conditions.

When provided with all geometrical input and operating conditions, GENF calculates the steam pressure, steam flow rate, circulation ratio, pressure drops, and other thermal-hydraulic data. The circulation ratio is a ratio of total flow through the tube bundle to feedwater flow.

For a dry and saturated steam generator, there exists a hydrostatic head difference between the downcomer and the tube bundle. This head difference serves as the driving head to circulate flow between them (see Figure 5-1). The driving head is constant for given operating specifications, such as power level and water level. The total pressure drop through the circulation loop is equal to the driving head.

Pressure drops depend on loss coefficient and flow rate (i.e., velocity). Loss coefficient consists of friction loss and form loss; the majority of the loss is due to the form loss in the steam generator. Since the driving head is constant, a higher loss coefficient means a lower circulation flow rate and a lower circulation ratio. A lower loss coefficient yields a higher circulation ratio. Therefore, an accurate prediction of the circulation ratio depends on an accurate loss coefficient.

Model boiler and field tests are used in qualifying the loss coefficients in the flow loop of the steam generator. For example, the major contributors of the pressure drop are the primary separator and tube support plates. The loss coefficient of the primary separator has been verified using model boilers and field steam generators (Reference 5-3). Similarly, loss coefficients of tube support plates have been developed using test data; Figure 5-2 presents the correlation of the loss coefficient and test data.

Figure 5-3 shows a typical compalison between predicted and actual measured circulation ratios. There is good agreement in the circulation ratio between the prediction and measurement. _

The TRANFLO model uses the same loss coefficient correlations as the GENF code. This provides assurance in properly calculating the pressure drops throughout the steam generator.

5.3 MULTIFLEX Code Description The MULTIFLEX program is an engineering tool for calculation of pressure and mass flow distributions during rapid thermal-hydraulic transients caused by an imposed driving force on the system. The driving force is taken, throughout this report, to be a break of a  ;

5-4

secondary loop in a Pressurized Water Reactor (PWR) system. MULTIFLEX has been verified by comparison to test data and has been reviewed and approved by the United States Nuclear Regulatory Commission (USNRC).

The USNRC has reviewed Westinghouse WCAP-8708 (Proprietary) and WCAP-8709 (Non proprietary). The USNRC issued an approval letter on June 17,1977 together with a Topical Report Evaluation (Reference 5-4). It concludes that WCAP-8708 presents an acceptable computer program for use in calculations to evaluate the pressure time history of fluid within current Westinghouse reactor systems caused by subcooled decompression during a postulated loss of coolant accident. Based on this approval letter and comments in the attached Topical Report Evaluation, Westinghouse issued a revised WCAP-8708-PA-VI (i.e.,

Reference 5-5), which contained the Reference 5-4 Topical Report Evaluation.

The thermal-hydraulic portion of MULTIFLEX is based on the one-dimensional homogeneous model which is expressed in a set of mass, momentum, and energy conservation equations. Rese equations are quasi-linear first order partial differential equations which are solved by the method of characteristics. By the method of characteristics, the partial differential equations are reduced to ordinary differential equations. The formulation of the characteristic equations results in equations which include acoustic signal transportation.

The representation and analysis of complex hydraulic systems by MULTIFLEX is accomplished using a network of hydraulic flow paths and mathematical models of various hydraulic components, such as two- and multi-pipe joints, orifices, pumps, valves, etc.

MULTIFLEX has been used extensively to analyze, among other events, loss-of-coolant transients in both the primary and secondary sides of PWRs and waterhammer transients due to valve motion.

5.4 MULTIFLEX Models The MULTIFLEX model of the steam generator consists of a network of hydraulic flow paths, referred to as legs. Figure 5-4 shows the MULTIFLEX model in schematic form.

The entire hydraulic geometry of the secondary side of the steam generator is modeled with the legs by specifying appropriate values for flow area, flow length, elevation, loss coefficient, hydraulic diameter, and boundary coridition. He thermodynamic initial ,

conditions are specified with appropriate values for pressure, enthalpy, and mass flow.

The forcing function for the transient is described by a break model which simulates a double-ended, guillotine rupture of the main steamline, located at the outlet nozzle on the steam generator. The MULTIFLEX model, as depicted in Figure 5-4 and described by its associated input data, is equivalent to the TRANFLO model presented in Section 4.2.

5-5

5.5 Comparison of MULTIFLEX and TRANFLO Results Results of MULTIFLEX calculations for a SLB event initiated from a hot standby are shown in Figures 5-5 through 5-7. For the identical case, results of TRANFLO calculations are given in Figures 6-1 through 6-3 (see Section 6.2). From these figures, it is seen that neither MULTIFLEX nor TRANFLO results show a higher frequency oscillation in the TSP pressure drops as might be expected if acoustic wave effects were significant. Effect of the acoustic wave due te a SLB event is discussed in detail in Section 6.6.

TRANFLO predicts a flow split between TSP J and TSP L. That is, the transient flow at TSP J and below is in a downward direction, while in an upward direction at TSP L and above. MULTIFLEX gives a split between TSP L and TSP M. They differ by about one TSP span (i.e., about 43 inch tube span); this is in good agreement. It is shown in Section 8 that this shift in the flow split to above TSP L for MULTIFLEX has negligible influence on TSP displacement Table 5-1 presents the peak values for pressure drops over the transient for comparison. In general, MULTIFLEX yields higher values ranging from 5% to 90%, except TSP L, which is near the flow split. The time to the peak value is shorter for MULTIFLEX. Both codes predict the maximum drop to occur at TSP P, and their ratio is 1.24.

The differences in pressure drop between TRANFLO and MULTIFLEX are considered to be due to differences in modeling and solution schemes. Since the employed numerical method of solution in MULTIFLEX is the explicit scheme, numerical diffusion is minimized, thus providing an accurate spatial representation of the acoustic waves. TRANFLO could underpredict the acoustic effect because of the use of an implicit integration technique. An implicit numerical solution scheme tends to result in greater numerical diffusion than an explicit scheme. The effect of numerical diffusion on acoustic waves is to stretch the wave out spatially. For relatively thin structures like the TSPs, the effects of numerical diffusion can be significant because the pressure difference across the plate due to acoustic waves, if present, is underestimated.

There is a difference in the physical model for two-phase flow between TRANFLO and MULTIFLEX. TRANFLO uses drift-flux modeling, which accounts the effect of flow slip between liquid and vapor phases. MULTIFLEX considers a homogeneous model without slip. A homogeneous model tends to yield a higher void fraction than a drift-flux model. A drift-flux model would have a better estimate for void fraction. A higher void fraction could lead to a greater pressure drop. Thus, it could be expected that the MULTIFLEX homogeneous model would yield higher pressure drops than the TRANFLO d ift-flux model.

However, void fraction would not totally account for the difference observed.

5.6 Conclusions  !

This section presents a summary of the adequacy of the TRANFLO code for its current applications. Blowdown test data of simulated reactor vessels validate the adequacy of the 1

5-6 '

l

code in predicting the steam and water blowdown transient. The NRC has accepted the TRANFLO code in calculating mass and energy release to the containment during a steam generator blowdown due to feed or steam line break.

As part of its review, the NRC accepted the code's ability to accurately predict local thermal-hydraulics in the vessel. The calculated pressure agrees well with the measured vessel pressure. Flow through the internals of the steam generator depends on accurate prediction of pressure drops, which relies on the accuracy of the loss coefficients along the flow paths. Test data of pressure drops from model boiler and field steam generators have been applied to verify the correlations for the loss coefficients.

Westinghouse has made modifications to the code to better predict steam generator behavior following a SLB event. Westinghouse has performed the verification and validation consistent with the methods approved by the NRC staff for the original version.

To assess acoustic wave effects, the results between TRANFLO and MULTIFLEX have been compared. As expected for the steam environment and the diameter changes in the steam flow path (see Section 6.6), the MULTIFLEX results do not show any acoustic wave effects on the TSP loads. Thus acoustic wave effects do not need to be considered for the TSP loads. The minor differences between the TRANFLO and MULTIFLEX loads represent the sensitivity to the code used for the load analyses, and the differences between codes are bounded by the uncertainty factors applied to the TRANFLO TSP pressure drops.

In conclusion, the TRANFLO code is a verified program for adequately predicting thermal-hydraulic conditions during the blowdown transient of a steam generator due to a feed or steam line break.

5.7 References 5-1. R. E. Land, "TRANFLO Steam Generator Code Description," WCAP-8821, Westinghouse Nuclear Energy Systems, September 1976.

5-2. Memo from C. O. Thomas to E. P. Rahe, " Proprietary Content Review of SER on WCAP-8821 and WCAP-8822," October 14,1982.

5-3. P. W. Bird and P. J. Prabhu, " Review of}rimary Separator Loss Coefficient,"

WTD-TH-80-010, July 1980.

5-4. John F. Stolz, " Evaluation of Westinghouse Topical Reports WCAP-8708(P) and WCAP-8709(NP), United States Nuclear Regulatory Commission, June 17,1977.

5-5. K. Takeuchi, et. al., "MULTIFLEX, A FORTRAN-IV Computer Program for Analyzing Thermal-Hydraulic-Structure System Dynamics," WCAP-8708-PA- VI, Westinghouse PWR Systems, September 1977.

5-7

l i

Table 5-1 ,

Hot Standby Pressure Draps Calculated by TRANFLO and MULTIFLEX Ratio of Pressure Drop, psi MULTIFLEX f Plate TRANFLO MU., HFLEX - ig_TRANFLO i a,b A

C F

?

3 L ,

t M ,

N P

?

.b L

T t

D i

5-8

Steam Flow a

4 t 6 Steam

i +-

... :p :.

Primary Water Separator N g o, ,8j / Level l I ~-} o"*  % Recirculated

% o' g Water Feedwater , , ,N,

• l Flow ..

• * * , P oo **

Tube

[' j',*, Support ey oo e o!.p .

Plate so o ,

Downcomer . L ol a.! -

o l,  !

t a

• l 2. '

l e o* , Tube Bundle e o, l 7e  !

, de j 4 e o ,

. . l l s 1 ' 2,_-

Coolant In Coolant Out Figure 5-1. Diagram of Flow Circulation During Power Operation 5-9

- - - - _ v-__ ____ ___ ____ >,-,m~ . , _ w - - . - - - - w . ,_ m _ _ .

i a,b i

e i

t i

i t

t i

i

~  !

Figure 5 Counterbored Structural Quatrefoil Loss Coemeients .

h 5 - 10 i 6

i 1

a,b i 4

I i

i r

i i

i

?

e i

f

'i i

Figure 5-3. GENF Verification, Circulation Ratio Versus Load ,

i i

5 - 11 <

, ._ . ,e , __, -, . . . _ . _ _ . _ _ _ _ . . _ - _ _ _ _ . _ _ _ _ . . _ _ _ _ _ _ _ _ _ . _ ___ _ _ _ _ _ _ _ _ _ _ _

a J:

l 4

l I

)

l t

i f

I 5

I

?

I-D i

Figure 5-4. MULTIFLEX Model for Braidwood Unit 1 S/G {

4 h

i 5 - 12 y , -

m.- .,mm. -

7 , . _ , , - , , , _ _ - . ,, -

a h

i t

i t

t Figure 5-5. Pressure Drop Through Tube Support Plates M, N, and P 5 - 13

, +. , _ _ y.. a .- -

'I a

e 4

i i

i 1

t t

b f

I 5

F 3

-t i

r i

k i

J 1

Fi Fure 5-6. Pressure Drop Through Tube Support Plates F, J, and L 5 - 14

a  ;

b t

t F

E r

i 2

i 1

j Figure 5-7. Pressure Drop Through Tube Support Plates A and C 5 - 15 i

a K s ..ifA,_g 4- 4 +L 4--1 a w@ h 4 -- 4 & a- .+brs A --.;.- I A >a._ . A,_

i a

.\

I i

J ie t

i 9

1 a

h i

k o

e h

4 e

h f

?

L i

e 1

se e

b i

1 b

b L

?

. i k

l s

ti t

i 6.0 HYDRAULIC SLB LOADS ON TSPs 6.1 Analysis Plan ,

The hydraulic, pressure drop loads on the TSPs in a SLB event are calculated with the TRANFLO computer code as described in Section 4. An extensive set of TRANFLO analyses was performed and the results of these analyses are reported in this section. The TRANFLO analyses include: reference analyses for a guillotine SLB both inside and outside containment at both hot standby and full power conditions; and a series of analyses to assess the sensitivity of the TSP loads to the most influential input parameters including the break size for a limited break, the S/G water level, the TSP pressure drop loss coefficient, the S/G downcomer loss coefficient and the Moody discharge coefficient. The results of the sensitivity analyses are used to define a conservative or bounding uncertainty factor which is applied to the reference analyses pressure drops to obtain a bounding set of loads for the TSP displacement analyses. The reference analysis loads on the TSPs and the bounding reference loads adjusted by the uncertainty factors are used to develop the expected and bounding TSP displacements in Section 8.

The TRANFLO analysis matrix is given in Table 6-1 and is described below:

Reference 6nalyses The reference analyses, Cases I to 4 of Table 6-1, provide the expected TSP loads for a guillotine steamline break either inside containment at the exit of the S/G nozzle or at a location just outside of the containment building. The break at the S/G nozzle would result in a potential radiation release inside of containment, while the break outside containment represents containment bypass for the potential radiation release. In addition to the break size, the TRANFLO input parameters most influential on the TSP loads are the water level, the TSP loss coefficients, the downcomer loss coefficients and the Moody discharge coefficient. For the reference analyses, the expected values are used for the water level and the TSP /downcomer loss coefficients. The Moody discharge coefficient has been set at its maximum value (most conservative) of 1.0. The reference analyses thus lead to a moderately conservative representation of the expected TSP pressure drops, under hot standby and full power conditions, for a guillotine SLB inside and outside containment.

TRANFLO Srnsitivity Analyses TRANFLO sensitivity analyses were performed to assess the influence of the principal input parameters, individually and collectively, on the TSP pressure drops. The more limiting hot standby SLB event was used to determine the TSP load sensitivity to the break size (guillotine versus 1.5 ft 2), water level (expected 487" versus lower uncertainty level of about 466"), TSP loss coefficients (expected versus maximum values), downcomer loss coefficients (expected versus minimum values) and the Moody discharge coefficient (maximum 1.0 versus expected 0.84). These analyses are represented by Cases 51 to 55 in Table 6-1. From these analyses, it was determined that the reduction in water level, the increased TSP loss coefficients and the reduced downcomer loss coefficients resulted in an increase in the TSP TSP,06 WPS 6.] WW I. W

loads for the most limiting plates (plates C and J for TSP displacements) and typically for most other plates.

The above three input parameter changes that increase the TSP loads were then combined, in Cases 61 and 62, with the most limiting S/G guillotine break (with a Moody discharge coefficient of 1.0) to define a bounding set ofloads to develop an uncertainty adjustment factor for the reference analyses. The uncertainty factor is a constant factor on the loads obtained to reasonably bound the ratio of the peak pressure drop for each plate from the combined analyses of Cases 61 and 62 to the reference analyses of Cases 1 and 2.

Beference A_nalyses with Uncertainty Adiustment Cases 61 and 62 combine all the worst case or most limiting input parameters to define the combined uncertainty factor for each plate. It is extremely conservative to combine all the collective worst case input conditions in a single TSP load analysis. A more realistic bounding analysis would collectively combine intermediate (between expected and worst case) values for the input parameters in a single run, or would combine the resulting TSP loads by the square root of the sum of squares since all worst case conditions would not be expected to act simultaneously. However, it is the intent of these analyses to develop a very conservative set of bounding TSP loads, and the results of Cases 61 and 62 are used to define uncertainty factors. To further simplify the uncertainty factor approach, a constant uncertainty factor is developed based on the maximum ratio of the Case 61 to Case 1 and Case 62 to Case 2 pressure drops. As developed later in this section, this resulted in the Case 2, full power loads for all plates being increased by a factor of 1.75. For the Case 1, hot standby loads, the uncertainty factor applied was 2.0 for plates A to L (1 to 8) and 1.5 applied to plates M to P

( (9 to 11). Thus the reference analyses with the uncertainty adjustment provide a direct assessment of a constant factor increase in the loads. Cases 11 and 12 represent the adjusted load conditions and are considered to define an upper bound on the TSP pressure drops.

Case 21 provides a best estimate analysis for comparison with the reference analysis of Case 1 or the bounding analysis of Case 61. The best estimate conditions were selected to represent a more limited (than a guillotine break) SLB event with containment bypass and with the expected Moody discharge coefficient of 0.84. The analysis results, given later in this section, show that the best estimate loads of Case 21, representing a more likely SLB event, are typically about half the loads of Case 1 and about 1/4 the loads of the bounding Case 11. Since the loads of Case 11 are shown in Section 10 to result in acceptable tube burst probabilities for Braidwood-l and Byron-li it is clear that the more probable SLB event and expected S/G conditions would lead to small TSP displacements with a very low tube burst probability and would essentially eliminate tube burst as a concem for APC applications.

Cases 31 of Table 6-1 represents the MULTIFLEX computer code analysis described in Section 5 for qualification of the TRANFLO code. This analysis is used in Section 8 for TSP displacement analyses and is included in the table to define the complete analysis matrix.

TSP,06 WPS 62 w= A

• The TSP loads for Cases 1,2,11,12 and 31 are used in Section 8 to calculate the associated TSP displacements.

6.2 Reference Full Power and Hot Standby Loads Case 1 is a reference analysis for SLB at hot standby with a guillotine break at the S/G outlet nozzle. Its water level is at the normal setting of 487 inches above the top of the tubesheet, and both TSP and downcomer loss coefficients are at nominal values. The Moody discharge coefficient for critical blowdown flow is set to be unity. Figures 6-1,6-2 and 6-3 show the hydraulic loads through eight TSPs P, N, M, L, J, F, C and A.

Case 2 is equivalent to Case 1, except the SLB is initiated from full power operation.

Figures 6-4,6-5 and 6-6 show the hydraulic loads through eight TSPs P, N, M, L J, F C and A.

Case 3 is identical to Case 1, but the guillotine break at hot standby is outside the containment building. Figures 6-7,6-8 and 6-9 show the hydraulic loads through eight TSPs P, N, M, L, J F, C and A.

Case 4 is equivalent to Case 3, except its initial condition is full power operation, not hot standby. Figures 6-10,6-11 and 6-12 show the hydraulic loads through eight TSPs P, N, M, L, J, F, C and A.

For hot standby, a break at the S/G steam outlet nozzle generally yields higher loads than a break outside the containment. This is expected because of the additional flow resistance due to extra piping of about 120 feet and three 90* elbows between the steam outlet nozzle and the containment penetration. There is a flow split around the middle of tube bundle (between TSPs J and L). Flow is upward for TSP L and above (see Figure 4-1), and flow is downward for TSP J and below. Except for a low feedwater flow of 200 gpm through the auxiliary nozzle, the main feedwater nozzle is isolated during hot standby.

For full power operation, differences in hydraulic load.s are negligible between a break at the S/G steam outlet nozzle and outside of the containment building. Apparently, blowdown flow through the break has a negligible effect on the tube bundle flow for the full power mode, because there is not much water for flashing. During most of the transient, loads are more or less equal to those during steady state full power, because of continued feedwater flow through the main feedwater nozzle. Thus, there is no flow split within the tube bundle.

It is important to note that hot standby leads to higher loads on the TSPs than full power operation for both upper and lower TSPs, but middle TSPs tend to experience higher loads for full power than hot standby. Middle TSPs like J and L are where the flow split takes place for hot standby, and thus the flow rate in either direction tends to be small because they are not reaching fully developed conditions. Therefore, hot standby usually results in lower hydraulic loads at the middle TSPs than full power operation. Both upper and lower TSPs TSP,06 WP5 6-3 hh- t

• l 1

l experience higher loads for hot standby than for full power. This is due to the difference in amount of water available for flashing, as discussed below.

Once a SLB event begins, it triggers a rapid depressurization, which leads to water flashing. 1 The rapid water flashing generates water motion. Fluid velocity increases with an increase in the amount of water flashing. A higher water velocity will lead to a larger pressure drop across a TSP. In addition, when fluid moves in the tube bundle, water will exert a higher pressure drop across the TSP when compared to steam. Hot standby at zero power provides a i solid water pool in the tube bundle, while power operation generates a steam and water mixture. Therefore, hot standby yields higher loads than full power operation.

6.3 SLB Load Dependence on Water Level Sensitivity studies of water level on hydraulic loads on tube support plates have been performed and documented. Results show that the lower the water level, the larger the hydraulic loads on the TSPs. Figure 6-13 shows a relative load on the uppermost TSP with respect to the water level of a steam generator. This trend is expected, as explained below.

A postulated steam line break (SLB) event results in blowdown of steam and water out of the steam generator. The fluid blowdown depressurizes the secondary side fluid and thus causes the fluid to move. Fluid motion leads to pressure drops and thus hydraulic loads across the TSPs and baffle plates Depressurization triggers rapid water flashing, mainly across the water level during the early part of the transient. The rapid water flashing generates water and steam motion, and the closer the tube support plate is to the water level, the higher the flow rate and thus the higher the pressure drop.

Therefore, a chosen water level being lower than the normal setting is most conservative in determining and providing hydraulic loads as input to displacement analysis of the TSPs. The normal water level setting at hot standby for a Model D4 S/G, including Braidwood- 1, is about 492 inches above the top of the tubesheet. Braidwood Unit I actually controls at a level about 5 inches lower (or 487 inches above the top of the tubesheet); it is about 206 inches above the uppermost TSP. The span between the uppermost TSP and mid-deck is 264 inches, over which Braidwood Unit I has a 78% span.

The TRANFLO TSP loads at hot standby in WCAP-14046 are based on a water level at the uppermost TSP, which is excessively conservative. The control point water level of 487 inches above the tubesheet is the reference water level for the reference Braidwood-l TRANFLO analyses (i.e., Cases 1,2,3 and 4 of Table 6-1). The uppermost TSP peak pressure drop for the Braidwood-l updated analysis at the controlled water level is a factor of about 2.5 lower than the corresponding load given in WCAP-14046. The WCAP-14046 loads also include the adnitional conservatism of a feedwater transient coincident with the SLB.

This combined event is considered to be a low probability event and the feedwater transient is not considered in the Braidwood-l reference analyses for a SLB from hot standby.

T5P 06WPS 64 bebe 8. W

l l

6.4 Best Estimate Loads As stated earlier, the hydraulic load varies with initial and boundary conditions of a SLB event. The best estimate case (i.e., Case 21 of Table 6-1) consists of the following:

Break outside containment A limited break size of 1.5 ft' A nominal water level of 487 inches Nominal correlation constant of TSP loss coefficient Nominal downcomer loss coefficient A Moody discharge coefficient of 0.84.

There has been no steam line break event on Westinghouse designed PWRs. Two main '

1 feedline pipe breaks have occurred on Westinghouse designed PWRs. These breaks were outside containment.

A guillotine break of a steam line results in a double-ended break and a break area of the total pipe cross section. The probability of a guillotine break is extremely low. In reality, the steam line would leak before breaking, and the break takes some finite time to open up to its final size. Therefore, the blowdown rate is limited at the beginning since it depends on the break area. By the time it reaches its final break size, the system pressure drops, on which the blowdown flow rate depends. The maximum likelihood of the break size could be less than half of the pipe cross section.

The TRANFLO code does not simulate the break size as a function of time during the break of the steam line. It will take the final size and apply it to the beginning of the breaking.

Therefore, recognizing such a model and the nature of a non-guillotine break with gradual opening, we could consider a break size of 1.5 ft2 , about one third of the steam line flow area of 4.71 ft 2. Note that the throat area of the S/G steam nozzle is about 1.4 ft 2, If the break were to occur at the exit of the steam generator, it is assumed to be three feet from the top of the steam nozzle.

TRANFLO uses the Moody model for calculation of break (or critical discharge) flow.

Comparisons with measured break flows of pressure vessels have indicated the need to use a multiplier for the Moody model. The multiplier i,s less than unity, depending on the conditions of the two-phase flow and the geometry to and through the break. The multiplier can be as low as 0.55 and as high as 0.84.

Just the geometry alone, Moody model, like many critical flow models of two-phase flow discharge, considers a one-dimensional flow problem. In reality, it is not one dimensional flow. Consideration of two-dimensional flow demonstrates that a multiplier of about 0.84 is needed to achieve agreement with measured data. A multiplier of 0.84 is used as a best value for a SLB event for the Braidwood-1 and Byron-1 units. Figures 6-14,6-15 and 6-16 show l pressure drops through TSPs for this best estimate case. As expected, the pressure drops are TSP,06 WP5 6.$ he 8, W

smaller than those of Cases 1 and 3. The ratios to those of Case I range from about 0.4 to 0.8, depending on the individual TSP (see Table 6-2).

6.5 SLB Load Sensitivity Analyses 6.5.1 Break Size According to Section 6.4, a break size of 1.5 ft2 is considered at the exit of the S/G steam outlet nozzle.

The time histories of the pressure drops through various TSPs are similar to those of Figures 6-1,6-2 and 6-3 for Case 1, but smaller as expected for a limited break size.

Table 6-2 presents ratios of the peak values of the pressure drops between Case 51 and Case 1; the peaks are about 10 to 20% smaller for the limited break than a guillotine break.

6.5.2 Water Level The normal water level setting at hot standby and full power operation is 487 inches above the top of the tubesheet for the Braidwood-l steam generators. According to system control, the control is accurate within 1% of the 233 inch span of the narrow range taps. The water level for hot standby can drop below the normal setting due to steaming by the reactor residual heat. Ilowever, the water level is quickly restored within about 10 minutes at a typical flow rate of 200 gallons per minute. The cross sectional area of fluid space at this level is about 152 ft'. This implies that the water level can momentarily drop by about 21 inches. In other words, we consider a drop of the water level from 487 inches to 466 inches.

For the minimum water level, Case 52, the pressure drops through the various TSPs are similar to those of Figures 6-1,6-2 and 6-3 for Case 1, but higher as expected for a lower water level. Table 6-2 presents ratios of the peak value of pressure drops between Case 52 and Case 1; the peaks are about 10 to 30% higher for the water level of 466 inches than for 487 inches.

6.5.3 TSP Loss Coefficient As shown in Figure 5-2, the best value of the. correlation constant of the loss coefficient correlation is 1.1. Its upper and lower bounds ire 1.4 and 0.8, respectively. Reference ~ '

analyses of Cases 1 to 4 use the best estimate of 1.1. The upper bound of 1.4 is applied to assess the sensitivity of the TSP loss coefficient on pressure drops. Use of the upper bound of 1.4 leads to higher pressure drops.

For the maximum TSP loss coefficient, Case 53, the dynamic behavior of pressure drops through the various TSPs are similar to those of Figures 6-1 to 6-3 for Case 1, but higher in ,

value as expected. Table 6-2 presents ratios of peak value of pressure drops between Case 53 l and Case 1; the peaks are about 15 to 30% higher.

i I

rse_uwPs 6-6 Nn etw 8. W

6.5.4 Downcomer Loss Coefficient The loss coefficient of the downcomer consists of friction along the shell and wrapper, and form loss due to area changes and flow turns. The estimate of friction and form loss for the downcomer is straightforward, as it is within single phase (water) flow, and simple geometry; but, it can be subject to uncertainty, too. Its total drop is only about 0.6 psi during full power operation.

For sensitivity study, we will consider a possibility of a decrease of about 0.3 psi, or a 50% j decrease in the downcomer pressure loss. This decrease is achieved by a decrease in the  !

downcomer form loss coefficient. A decrease in downcomer loss will promote more tube bundle flow toward the tubesheet and then entering the tube bundle.

For the minimum downcomer loss coefficient, Case 54, the time histories of pressure drops through various TSPs are similar to those of Figures 6-1 to 6-3 for Case 1. Table 6-2 presents I i

ratios of peak value of pressure drops between Case 54 and Case 1; the ratios depict that downward flow increases by a few percent for TSPs A, C, F and J. Indeed, the upward l flow decreases slightly,2% for TSP N and 1% for TSP P. These are expected because the downcomer loss is very small compared to the total loss through the whole bundle and separator. l l

6.5.5 Moody Discharge Coefficient Pressure drop increases with an increase in flow rate, and vice versa. A decrease in break flow generally decteases the tube bundle flow and thus reduces the pressure drops across the TSPs. A Moody discharge coefficient of 0.84 reduces break flow compared to that of a coefficient of unity.

For the lower,0.84 Moody discharge coefficient, Case 55, the pressure drop transients through various TSPs are similar to those of Figures 6-1 to 6-3 for Case 1, but essentially lower. Table 6-2 presents ratios of peak value of pressure drops between Cane 55 and Case 1; the peaks are up to 20% smaller for Case 55 than for Case 1.

l 6.5.6 Combined Worst Conditions for Hot Standby Combined worst conditions are given in Table 6,1; they are:

Break outside steam outlet nozzle A guillotine break A lower water level of 466 inches I An upper bound for correlation constant of TSP loss coefficient Minimum loss coefficient A Moody discharge coefficient of 1.0.

Figures 6-17,6-18 and 619 depict pressure drops through the TSPs. As expected, they are higher than those of Cases 1 and 3. Their ratios to those of Case 1 range from about 1.4 to 2.1, depending on the individual TSP (see Table 6-2).

I T3P,06 WP5 67 Neba 8. W

I I

6.5.7 Combined Worst Conditions for Full Power  ;

The combined worst conditions for full power are identical to those for hot standby (see l Section 6.5.6). Figures 6-20,6-21 and 6-22 show pressure drops through TSPs. As expected, they are higher than those of Case 2 (see Table 6-2).

6.6 Acoustic Wave Consideration A steam line break does generate a decompression wave, which can enter the steam generator.

This pressure wave that starts at the break has to propagate through the tortuous path provided by the primary and secondary separators with their numerous area changes. It has to pass through the U-bend of the tube bundle. Before it reaches the uppermost or lowest TSP it will attenuate significantly along the path.

The MULTIFLEX calculation for Case 1 confirms that an acoustic wave due to a SLB event is indeed insignificant, as to be discussed below using the MULTIFLEX result.

In order to explain the results obtained, it is useful to first discuss what results would be obtained for a simplified, idealized case. The idealized case is a two foot long section of piping connected to the top of the steam generator. The area of the piping is 5 2ft while that of the steam generator is 99 ft' There is no flow restrictor, zero friction in the piping, constant fluid properties, the break occurs in zero time, and the pressure at the break plane remains constant. The break will generate a pressure wave of magnitude P = 1100 - 325 = 775 psi that has zero spatial length (a step change in pressure) and that will propagate at sonic speed to the steam generator. At an elapsed time of 0.00125 second (or t = L/c = 2 ft /1605 fps), the wave will arrive at the steam generator. Because of the area change at the steam generator, the initial blowdown wave will be decomposed into a penetration wave and a reflected wave.

The penetration coefficient is as follows:

C, = 2A,, / (A,, + Aso) u 2(5 ft2 ) / (5 ft2 + 99 ft' ) = 10%

Therefore, the reflection coefficient is 1 - C, = 90%. The penetration wave of magnitude P = 0.l(1100 - 325) = 78 psi will propagate into the steam generator, dropping the local pressure to a value of P = 1100 - 78 = 1022 psia. 'Ihe reflection wave of magnitude P = 0.9(1100-325) = 697 psi will propagate back toward the break, increasing the local pressure to a value of P = 325 + 697 = 1022 psia. Again at an elapsed time of 0.' J0125 sec, the initial reflection wave will arrive at the break and be reflected again, back toward the steam generator. Since the break plane is essentially a constant pressure boundary condition, any wave reflected from it will have the same magnitude as the incident wave but opposite sign. Thus, at time 2(0.00125) = 0.00250 second, there is a wave of magnitude 697 psi propagating toward the steam generator, reducing the local pressure to P = 1022 - 697 = 325 psia. When this wave arrives at the steam generator it is decomposed into a penetration wave of magnitude 0.l(1022 - 325) = 70 psi and a T5P,,06 WP3 68 habe 8. W

reflection wave of magnitude 0.9(1022 - 325) = 627 psi. This precession of waves will continue in an analogous manner until a quasi steady-state flow condition is achieved. From the perspective of the pressure just inside the steam generator it can be seen that the

, initial blowdown derssurization wave of magnitude 775 psi is reduced to a series of depressurization waves of much smaller and diminishing magnitude (i.e. first 78 psi, then 70 psi etc.), each separated by a time interval of 0.00250 second.

For the steam generator model analyzed in this report, the following factors '.erve to modify the above behavior for an idealized case: the presence of the flow restrictor, friction at the pipe wall, a 1 msec break opening time, and the presence of many area changes within the flow path inside the steam generator. The effect of the flow restrictor, within the context of this discussion, is to provide an additional point of wave decomposition. Each wave which arrives at the flow restrictor is split into two waves in a manner which is similar to that described above. The effect of friction at the pipe wall is to stretch the acoustic waves over an increasing spatial length, thus changing the ideal step change in pressure to a spatial ramp in pressure. The effect of the 1 msec break opening time is to change the initial blowdown wave from a step change to a ramp over 1 msec (and a corresponding spatial length). The effect of the additional area changes within the steam generator is to further decompose the acoustic waves traveling inside the steam generator.

' The combined effect of the above factors is to obscure the presence of individual acoustic waves, except for very early in the transient at locations relatively close to the break. After just a few milliseconds or at locations within the steam generator, the transient does not exhibit the characteristically rapid changes in pressure and mass flow associated with waterhammer, but rather appears as a relatively slow depressurization event. Due to friction primarily, the step changes in fluid parameters are transformed into linear ramps of finite spatial lenr.s Because of the complex network of area changes and superposition of the resulting l e number of individual acoustic waes; the individual linear ramps combine to resun in c, ..rinuous, smooth changes in fluid parameters rather than discrete step changes.

The transient is still driven by the propagation of acoustic waves, but individual acoustic waves are not discemable except as noted above.

The following figures, taken from MULTIFLEX results for the Hot Standby Case 1, illustrate the foregoing discussion: Figures 6-23 and 6-24 are the pressure at the break plane, plotted against two time scales: Figures 6-25 and 6-26 are the pressure at the top of the steam generator, again for tva .ime scales; and Figure 6-27 is the pressure at the upper surface of the uppermost tube support plate. Figure 6-24 shows that the initial blowdown wave is of magnitude P = (1106 - 330) = 776 psi with a duration of 1 msec. Figure 6-26 shows that the portion of the initial blowdown wave that penetrates the steam generator beginning at approximately 1 msec is of magnitude P = (1100 - 1080) = 26 psi, where 1080 is the pressure at time t = L/C + 0.001 = 0.00225 second (wave propagation time plus break opening time).

These waves are confined within the early few milliseconds of the transient, while the pressure drops reach their peak values at a much later time (at least 0.3 second), see Figures 5-5 to 5-7. This indicates that the dynamic transient is essentially a hydraulic process, not an acoustic one.

m.* wr5 6-9 Na'a=6ar 9.1994 I

b Figure 6-27 shows no discernable step change in pressure associated with the initial blowdown wave. Figures 6-25 and 6-27 show that, except for early in the transient, ,

individual acoustic waves (of the classic step-change type) from the blowdown event are not '

discernable inside the steam generator.

6.7 Conclusions Considering the above discussion, we can draw the following conclusions.

1. SLB at hot standby generates higher pressure drops across tube support plates than SLB at ,

full power operation. This trend is observed for most of TSPs, except for middle TSPs such as Plates J and L (see Reference Cases 1 to 4 in Table 6-2).

2. The best estimate case (i.e., Case 21) of hot standby SLB has limited break outside the containment and a Moody discharge coefficient of 0.84, and re .its in smaller pressure drops than Case 1.

3. Sensitivity studies indicate that 1) a limited break is less severe in pressure drop than a guillotine break,2) pressure drop increases with a decrease in water level,3) an increase in TSP loss coefficient increases the pressure drop,4) a decrease in downcomer loss coefficient has negligible effect on the pressure drop, and 5) a decrease in Moody discharge coefficient reduces the pressure drop.

4. Combined worst conditions were constructed from the sensitivity study. The sensitivity study has been applied to develop a bounding uncertainty factor on the TRANFLO TSP pressure drops. The resulting factors of 2.0 on hot standby loads and 1.75 on full power loads are applied to the reference Cases I and 2 loads for the TSP displacement analyses given in Section 8.

5. The effects of an acoustic wave initiated from the decompression at the break can penetrate into the steam generator but are less than 5% of the initial wave, and the acoustic wave effects dampen out within milliseconds. The penetrated waves do not travel deep into the tube bundle and are insignificant at the TSPs. The dynamic transient due to an SLB event is essentially a hydraulic process, not an acoustic one.

r Tse_os wes 6 - 10 Nmmba 9. W

Table 6-1. TRANRO Analysis Matrix Analysis Conditions Case Operating Condition snet Break w ater Tsr DC. Moody Lac. Size I m et Ims Ims Diech.

Coeff. Coeff. Coeff.

Reference Analyses 1 Reference Hot Standby, SLB at S/O Nonle S/Gm Guil. /87" Nom. Nom. 1.00 2 Reference Full Power SLB at S/O Nonle 3 Reference Hot Standby, SLB outside Containment O.C.m 4 Reference Full Power, SLB Outside Contamment Reference Analyses with Uncertainty Adj==*==at*

II Reference Hot Standby, Adjusted for Ui; certainties S/G Guil. 487' Nom. Nom. 1.00 )

_ j 12 Reference Full Power Adjusted for Uncertamues )

Best Fdmate Analyses 21 Best Estanate flot Standby 0.C. 1.5 ft' 487" Nom. Nom. 0.84 TRAhTLO Qualification Analyses (MULTIFLEX Code) {

i 31 Reference flot Standby, SLB at S/O Noule, S/G Guil. 487* Nom. Nom. 1.0 >

MUL11rLEX TRANH4 Renaltivity Analyses 31 liot Standby, Break Area Sensitmty S/G 1.5 ft' 487* Nom. Nom. 1.00 52 Hot Standby, Water Level Sensitivity Guil. 466' Nom. Nom. 1.00

$3 Hot Standby, TSP Loss Coeff. Sensitmty 487* Max. Nom.

$4 Hot Standby, Downcomer Loss Coctf. Sensiuvny

• Nom. Min.

$5 Hot Standoy, Moody Coeff. Sensarvey Nom. 0.84 61 Hot Standby, combined Worst Case Uncertainties S/G Guil. 466' Max. Min. 1.00 62 Full Power, Combmed Worst Case Unc.

Notes: 1. S/G break location is at exit of S/G nozzle. O.C. break location is outside containment penetration. l

2. An adjustment factor to increase the TSP loads is applied from these analyses based on the results of the TRANFLO sensitivity analyses.

l I

isr_o6 wr5 6 - 11 N==b- s. *

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

Table 6-2 ,

SLB Peak TSP Pressure Drops and Ratio of Each Case to Case 1 l Peak TSP Pressure Drop Elais . Cass.1 Cw.t2 cue 3- cue 4 p.ms.n cae51 cue 52 cae 53 Cas.14 cme 55 Gaarl1 Ems.62 .

s A 1.42 -0.28 1.27 -0.20 -0.60 1.29 -1.52 1.83 1.46 -130 2.04 -0.28  ;

l C 1.21 -030 1.05 -0.30 -0.43 -1.11 1.51 -1.56 -1.24 -1.13 233 0.48 ,

F -0.7) -0.53 -0 61 -0.53 032 -0 67 -0.91 '- 0.92 -0.74 -0.69 1.40 -0.81 {

i J -039 -0 69 -032 -0.71 0.25 -038 -0.52 -0.50 -0 41 -0 40 -0.83 1.13 ,

L 0.64 -0 84 0.71 -0.92 0.50 0.52 0 80 0.73 0.64 0.49 1.12 -134 M 0.58 -0.34 0.60 -0.38 0.48 0.50 0.71 0.70 0.58 0.48 0.85 -0.58 ,

N 1.02 0.33 0.99 -0.40 0.79 'O.92 1.15 1.26 1.00 0.92 139 -0.58 [

130 1.82 1.44 130 2.05 -0.85 t P 1.45 -0.41 1.46 -0.45 1.17 1.68 t

Note: The peak pressure drop minus the steady state pressure drop at time = 0 represents the change in pressure drop from the SLB event i and is judged to be the most appropnate quantity for companng the analysis cases. For an SLB at hot standby conditions, the steady state pressure drop is zero.

Ratio of Each Case to Case 1

?

i Elatt fatl Gaas 2 cwe3 cwe4 Catn Coe 51 cwe 52 case 53 cue 54 Canth Coe61 case 62 <

A 1.00 0.20 0.89 0 14 0 42 0.91 1.07 1.29 1.03 0.92 1.44 0.20 C 1.00 0.25 0.87 0.25 036 0.92 1.25 1.29 1.02 0.93 1.93 0.40 ,

F 1.00 0.75 0.86 0.75 -0.45 0.94 1.28 130 1.04 0.97 1.97 1.14 L

J 1.00 1.77 0.82- 1.82 -0.64 0.97 133 1.28 1.05 1.03 2.13 2.90 l l

L 1.00 131 1.11 1.44 0.78 0.81 1.25 1.14 1.00 0.77 1.75 2.09  ;

M 1.00 0 59 1.03 -0.66 0.83 0.86 1.22 1.21 .1.00 0.83 1.47 -1.00 N 1.00 C32 0.97 -039 0.77 0.90 1.13 1.24 0.98 0.90 1.36 -G.57 .j P 1.00 -0.28 1.01 -0.31 0.31 0.90 1.16 1.26 0.99 0.90 1.41 -0.59 P

r i

i TSP,06 WPS 6 12 Ne 8. l* j

, e .-n , ,. .

a 2

Figure 6-1. Pressure Drop Through Tube Support Plates M, N, and P - Case 1 T5P,06 %75 6 13 Novmbe8, M4

a l

l l

l Figure 6-2. Pressure Drop Through Tube Support Plates F, J, and L - Case 1 TSP,06 WP5 6 - }4 NamkAIm

I a  ;

-. i

.. l E

i t

5 i

r i

t a

i h

t f

t r

f e

t I

..f Figure 6-3. Pressure Drop Through Tube Support Plates A and C - Case 1 TSP _M WM 6. ]$ Nwek 8 N 1

l

I a

P 1

1 1

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i

.i i

t i

6 e

t I

t I

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Figure 6-4. Pressure Drop Through Tube Support Plates M, N, and P - Case 2 i

""* 6 - 16 9 ,w ,, im  ;

i

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a l

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i 1

l l

l 1

l l

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Figure 6-5. Pressure Drop Through Tube Support Plates F, J, and L - Case 2 TSP,06 WPS 6. )7 November 8,19M

a Figure 6-6. Pressure Drop Through Tube Support Plates A and C - Case 2 "P ""

- 6 18 ww w s. im

t i

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Figure 6-7. Pressure Drop Through Tube Support Plates M, N, and P - Case 3 rsr_mwrs 6 --19 N= =

• 8.

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Figure 6 8. Pressure Drop Through Tube Support Plates F, J, and L - Case 3 75P." wP5 6 - 20  %-w s. m

a l

l I

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i

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Figure 6-9. Pressure Drop Through Tube Suppcrt Plates A and C - Case 3 i t

Nwek s, im .

TSP,06 W 6 2}

a-5 6

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2  :

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i Figure 6-10. Pressure Drop Through Tube Support Plates M, N, and P - Case 4 f TSP,06 WP5 6 --22 Namh8.Im j t

4 s. -44 e- p 4 MA. -.AM M 4-a A A 4 J.k- e.+h- -- + -.en.4 2.4-- A I

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I b

i h

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l Figure 6-11. Pressure Drop Through Tube Support Plates F, J, and L - Case 4 TSP _06 WM 6 - 23 N==w . im

.I

i a

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l l

I Figure 6-12. Pressure Drop Through Tube Support Plates A and C - Case 4 TSP- "* 6 - 24 w..w . im

n ,. -n... -

l l

a l l

l t

i P

t f

i

-i r

F I

F t

i i

i i

Figure 6-13. Relative peak pressure drop across the uppermost TSP as a : l

- function of water level; water level ranging from the uppermost TSP  !

(280") to the mid-deck (544") - Model D4 S/G Novemk 3,1994 TSP.MWP5 6 2$

l l

l

a e

i Figure 6-14. Pressure Drop Through Tube Support Plates M, N, and P - Case 21 nr.m ures 6 - 26 N--w s. im

l.

a l-i

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l ,

I i

=

l l

l I

1 Pressure Drop Through Tube Support Plates F, J, and L - Case 21 l Figure 6-15. f i

Nwsunbar 8, IgN TSP,06 WPS 6 - 27 ,

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

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l a

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6 Figure 6-16. Pressure Drop Through Tube Support Plates A and C - Case 21 TSP,06 WP3 6 - 28 NamW B,im -  !

l r

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1 J

1 a

j f

3 a

S e

f 1 .

i 1

Figure 6-17. Pressure Drop Through Tube Support Plates M, N, and P - Case 61

[-

T5P,06 WP5 6 - 29 Na=w s. m h

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1 a ,

i 1

r i

i t

i 7

i P

r i

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1 Figure 6-18, Pressure Drop urough Tube Support Plates F, J, and L - Case 61 l TSP,06 WP5 6 - 30 Nwek I, W  ;

1 1

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a

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1 Figure 6-19, Pressure Drop Through Tube Support Plates A and C - Case 61 Novaunbar 8,1994 T5P,06 WP5 6 3}

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Figure 6-20. Pressure Drop Through Tube Support Plates M, N, and P - Case 62  ;

I TSP- 06WPS 6 - 32 w .w s. im

.. . _. -. ._ -.,- . _. . _ . . .- . . . .. - ._.-.=_ _..-- . -- _ .

I l

a +

1 i

v i

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+

1 l

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1

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l Figure 6-21. Pressure Drop Through Tube Support Plates F J, and L - Case 62 m_os e 6 - 33 w--w s. w I

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

_ . _ _ _ _ _ _ _ _ _ _ ~

a .

f i

-i t

l

.- 1 1

l l

Figure 6-22. Pressure Drop Through Tube Support Plates A and C - Case 62 -

i TSP.M WPS 6 34 Nwek8Im j l.

. -. ._ - . - = - - . - _ . = _ _ = . - . . . . .

i l

1 l

l a

I 1

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I 1

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'l T

E l

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p i

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Figure 6-23. Pressure at Steam Outlet Nozzle  :

I TSP,06

• 6 - 35 wa-w s. m ,

I g.- a ...7, ,.

a t

T P

=

Figure 6-24. Pressure at Steam Outlet Nozzle (Early Few Milliseconds) 1 75P,06 WP5 6 - 36 N==w s. im l

l

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

l l

l a

f 4

l i

i I

i i

i i

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5 I

I f

I i i

-l i i

i i

I i

t t

i i

T 5

1

?

1

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t f

A Figure 6-25. Pressure Immediately Upstream of Steam Outlet Nozzle i i

TSP M WM 6 - 37 wa=ws.7o- 'l 4

1 -

a I

1 l

l i

l 1

l i

Figure 6-26. Pressure Immediately Upstream of Steam Outlet Nozr!e (Early Few Milliseconds)

TSP- 06WP5 6-38 3,,,,,

a i

r

(

l Figure 6-27. Pressure at Tube Suppon Plate P ,

TSP,06 WP5 6 - 39 N bar s.1994

l l

7.0 STRUCTURAL MODELING FOR TSP DISPLACEhENTS 1

7.1 General Methodology Section 7.0 summarizes the structural modeling of the Model D4 tube bundle to determine relative tube / plate motions under steam line break loads. The analysis involves the  ;

preparation of a tube bundle model for the hot leg side of the tube bundle, consisting of the flow distribution baffle, seven tube support plates, tierods and spacers, channel head, lower shell, and tubesheet. The WECAN computer code, a general purpose fmite element code is used to develop the model. The model is composed mainly of shell elements, with beam elements used to model the tierods and spacers. Calculations are performed to define applicable dynamic degrees of freedom (DOF) for each piate. Once the DOF are defined, a global substructure is generated for the overall tube bundle. The dynamic response of the plates is then calculated using the special purpose computer program,pitdym.

Additional details for all aspects of the structural modeling are provided in the following sections.

t 7.2 Component Materials A specification of component materials is contained in Table 7-1, with the corresponding material properties summarized in Tables 7-2 through 7-7. The properties in these tables are taken from the 1971 edition (through summer 1972 addenda) of the ASME Code, the applicable code edition for Byron Unit I and Braidwood Unit 1. It should be noted that although the properties are provided over the temperature range 70*-700*F, the average temperature during the transient, based on the thermal hydraulic results is =550*F. Since temperature dependent properties cannot be used in substructures, properties for the finite element model correspond'to the values at 550*F. In addition, the material properties for the tubesheet and tube support plates must be modified to account for the tube penetrations and flow holes, and to account for added mass effects of the secondary side fluid. Additional ,

details of the material property modifications are provided later in Section 7.5.

7.3 TSP Support System The various TSPs and baffle plates are supported vertically using several support mechanisms.

A schematic of the tube bundle region is shown in Figure 7-1, with each of the plates identified. All of the plates are supported by three tierods/ spacers in each of the plate quadrants. In addition, plates C (3H), F (SH), and J (7H) (where the Braidwood-l and Byron-1 plate numbers are in parentheses),in the preheater region are supported at their center by a vertical bar welded to the partition plate, while the plates above the preheater, L (8H), M (9H), N (10H), and P (11H), are supported at their center by a central tierod and 7-1

spacer. Additional support is also provided to the plates by vertical bars welded to the wrapper and / or partition plate immediately above and below the plates.'

The support locations for the plates are shown in Figures 7-2 through 7-10. The location of the tierods and spacers is shown in Figure 7-2, while the plate / wrapper support locations are shown in Figures 7-3 through 7-10. The configuration of the tierod / spacer supports is shown in Figure 7-11, with the vertical bar / plate interface shown in Figure 7-12.

i The in-plane support for the TSPs is provided by wedges located around the circumference of each plate. In all cases, the wedges are welded to the wrapper. The wedges are intended to provide in-plane support for the plates. However, since the wedges are welded to the wrapper, there is resistance to upward vertical motion, due to the sloped face of the wedge (see Figure 7-12). In evaluating a oreliminary set of transients, vertical support provided by the wedges was not included. However, the resulting displacements showed that for the limiting plates, the maximum displacements occur near the tube lane at the outer edge of the plate. Thus, in evaluating subsequent sets of loads, wedge support (in the up direction only) is included at the 10' location for the plates showing the highest deflections in this area, Plates A (lH), C (3H), and J (7H). When the hydraulic laods are in the downward direction, the wedges at the corners of these plates do not provide vertical support and, as shown in Section 8, the maximum TSP displacements occur at the corners along the tubelane in the downward direction. All other plates have vertical bar supports at the corners and only smaller diplacements are found for these plates.

Regarding the tierods and spacers, the tierods are bars that are threaded into the tubesheet and run the full height of the tube bundle with a nut on the upper side of the top TSP, Plate P (llHL (For the central tierod, the lower end is threaded into a special coupling welded to the top of the partition plate at Plate L (8H).) Around the outside of the tierods are spacers which are located between each of the support plates. For the central tierod, the spacers are welded to each of the TSPs. For the spacers located at the periphery of the TSPs, however, there is no rigid link between the spacer and the support plates.

The lack of a rigid link between the spacers and TSPs for the outer tierods / spacers results in a non-linear dynamic system. The TSP / spacer non-linearities are included in the dynamic j solution. The tierods / spacers have different stiffness characteristics for upward and l downward loads, and these differences are incorporated into the model. For the up direction, the load path is through the spacers from the loaded plate to the one above it, up through the bundle to the top plate, where the load is transfe'rred to the tierod and down to the tubesheet.

For the down direction, the load path is through the spacers from the loaded plate to the one below it down through the bundle to the tubesheet. As a means of comparison, the combined

' For the Model D3 steam generator, the vertical bars are present only for some of the plates in the l preheater region, and only along the partition plate. Vertical support for the periphery of the plates is  !

I limited to the tierods and spacers, and in some cases local vertical support is provided by wedges welded to the wrapper. The wedges for the Model D3 steam generators are not welded to the wrapper in all cases. For a number of the plates, the wedges are welded to the TSPs and / or baflie plates. For cases where the wedges are not welded to the wrapper, they are unable to provide resistance to upward movement by the plates.

7-2 l

tierod / spacer stiffnesses for the outer tierod / spacers are summarized in Tables 7-8 and 7-9.

Table 7-8 provides the stiffness values for the " upward" load path, and Table 7-9 for the

" downward" load path. In general, the down load path is several times as stiff as the up load path. The combined tierod / spacer stiffnesses for the central tierod / spacer are summarized in Table 7-10. Recall that the central tierod / spacer runs from Plate L (8H) to Plate P (llH).

As will be shown later, the loads for these plates all act in the upward direction, so the combined stiffness is based on a load path up through the spacers to Plate P (11H), and then down to the coupling at the top of the partition plate,just below Plate L (8H). A summary of the spacer stiffnesses for the individual tube passes is provided in Table 7-11. These f stiffnesses are incorporated in the dyamic model for the non-linear TSP / spacer interface. l t

I As a result of the plate deflection and rotation under SLB loads, the potential exists for interaction between the TSP and the tubes. These interaction effects have been incorporated i

in the analysis. If the plate rotates locally such that the top surface of the plate contacts the tube on one side while the bottom surface of the plate contacts the tube on the other side, then the tube will bind up in the plate and restrict further deflection of the plate. The amount of rotation necessary to cause tube / plate interaction is summarized in Figure 7-13.

7.4 Finite Element Model The overall finite element model is shown in Figure 7-14. With the exception of the tierods, all of the structural components are modeled using three dimensional shell elements. The tierods are modeled using three-dimensional beam elements. The spacers are incorporated in the dynamics code through stiffnesses that are coupled to the various plate elements when the corresponding gaps are closed.

In modeling the plates, the various cutouts along the tubelane, the cutout for the FDB in the center of the hot-leg, and the cutouts at the outer edges of Plates N (10H) and P(llH), are accounted for. In terms of material properties, equivalent properties are specified only in the tubed regions of the plate. Actual plate properties are used along the tubelane and at the periphery of the plates.

l 7.5 Revised Material Properties As noted earlier, the material properties for the iubesheet and TSPs are modified to account for the tube penetrations, flow holes, and various cutouts. The properties that are modified are Young's modulus, Poisson's ratio, and the material density. In the case of the TSPs, the density is additionally modified to account for the added mass of the secondary side fluid.

Younc's Modulus / Poisson's Ra_ tic Due to the presence of flow holes in the TSPs, but not in the FDB, separate formulations are used to modify the material properties. Although different formulations are used for the two components, the same methodology is used in each case. Due to the square penetration pattern, different properties exist in the pitch and diagonal directions. T1e first step is to 7-3

establish equivalent parameters for Young's modulus and Poisson's ratio in the pitch and diagonal directions (E,*/E, E,*/E, v/, v,*), respectively. The equivalent Young's modulus for the overall plate is taken as the average of the pitch and diagonal directions. The next step in the process is to determine an equivalent value for the shear modulus, G'/G, for the plate.

This is done in a similar manner as for Young's modulus, starting with values in the pitch and diagonal directions, and then taking an average of the two values. The final equivalent value for Poisson's ratio is determined from the relationship between Young's modulus and the shear modulus.

A summary of the resulting effective plate properties for Young's Modulus and Poisson's ratio is shown in Table 7-12.

Material Density There are two aspects to revising the plate density. The first is based on a ratio of solid plate area to the modeled area. The second aspect corresponds to the plate moving through and displacing the secondary side fluid, creating an "added mass" effect. In calculating the added mass, the formulation shown below is used. j l

fj21 m, *pj Jg

( 4s d'

I d

I* R' 8d 'I ~' Q

where, pr = fluid density A, = solid area of plate A,, = flow area 1 = hole length (plate thickness) d = hole diameter b = hole pitch The first step in the process of calculating revised densities for the plates is to determine the applicable areas for the metal and the fluid. Summaries of the actual and modeled plate areas are provided in Tables 7-13. Summarized in Table 7-14 are the corresponding flow areas.

The resulting added mass is a direct function of the fluid density. Because the dynamic analysis cannot account for the change in fluid density with time, the analysis uses an average density value for the transient. Variations in the fluid density as a function of elevation in the tube bundle are accounted for by calculating an average fluid density for each plate.

Two sets of calculations are performed for added mass corresponding to whether the transient initiates from hot standby or full power. For transients initiating from full power, the secondary fluid has a much lower fluid density and a lower added mass. Summaries of the resulting fluid and structural masses, together with the resulting effective densities are summarized in Tables 7-15 and 7-16.

7-4

Note that the density modification for the tubesheet is based solely on an area ratio of the actual perforated tubesheet to the equivalent solid plate.

7.6 Dynamic Degrees of Freedom In setting up the global substructure, it is necessary to define the dynamic degrees of freedom. In order to define dynamic degrees of freedom for the TSPs, two sets of modal calculations are performed for each of the plates. The first set of calculations determine plate mode shapes and frequencies using a large number of degrees of freedom (approximately 120 per plate). The second set of calculations involves repeating the modal analysis, using a significantly reduced set of degrees of freedom (DOF). The reduced DOF are selected to predict all frequencies for a given plate below 50 hertz to within 10% of the frequencies for the large set of DOF. A frequency of 50 hertz was selected as a cutoff, as it is judged that higher frequencies will have a small energy content compared to the lower frequencies. This can be confirmed by noting that the frequency content for the dominant peak in the pressure time histories is typically less than 10 hertz.2 For each of the modal runs,in addition to symmetry boundary conditions along the "Y-axis", and vertical restraint at vertical bar locations, all the plates are assumed to be constrained vertically at tierod/ spacer locations.

A sample set of mode shape plots for Plate A (lH) for the full set of DOF are provided in Figures 7-15 through 7-17, while mode shapes for the reduced set of DOF are shown in Figures 7-18 through 7-20. A comparison of the natural frequencies for the full and reduced sets of DOF for the plates is provided in Table 7-17. Based on the tabular summary, the reduced set of DOF are concluded to provide a good approximation of the plate response.

Note that for Plate P (llH), the frequency for Modes 3 and 5 for the reduced set of DOF slightly exceeds the 10% objective for matching frequencies. These variations are not considered to be significant, and the selected DOF are judged to give an acceptable representation of the Plate P (llH) response.

Due to the possible interaction of Plates A (lH), C (3H), and J (7H) with the wedge at 10*,

these plates could respond with two different sets of mode shapes, depending on whether gaps at these locations are open or closed. Thus, two different sets of modal calculations were performed for Plates A (lH), C (3H), and J (7H). One set with the wedge support, and one set without the wedge support. A comparison of the results for the two cases is summarized in Table 7-18. (Note that the results reported in Table 7-17 are for the case without wedge support )

The reduced set of DOF consists of 8 - 13 DOF for each of the plates. Plots showing the resulting DOF for each of the plates are shown in Figures 7-21 through 7-28.

l

• Pressure time curves for each of the transient conditions are presented in Section 6.

{

7-5

i 7.7 Displacement Boundary Conditions The displacement boundary conditions for the substructure generation consist primarily of prescribing symmetry conditions along the "X" and "Y" axes for each of the components.

Vertical constraint is provided where the plates are constrained by the vertical bars welded to the partition plate and wrapper, and for the channel head at the locations corresponding to its interaction with the interfacing support structure. For the TSPs, rotations normal to the plate surface are also constrained, as required by the stiffness represeraation for the plate elements.

7.8 Integration Time Step / Structural Damping The dynamic time step used in evaluating the SLB transients is 0.0002 second. This time step was selected based on analyses using similar models and loadings where various time steps were considered, and 0.0002 second was shown to result in a converged solution. The analysis incorporates structural damping of 4%, which is judged to be a conservative value for the type of dynamic loading and response (movement of the plates through the secondary fluid) being considered.

7.9 Application of Pressure Loads The SLB pressure loads act on each of the TSPs. To accommodate the variation in load from plate to plate, load vectors are prescribed for each of the plates using a reference load of 1 psi. The reference loads are then scaled during the dynamic analysis to the actual time-history (transient) loading conditions. The transient pressures summarized in Section 6 are relative to the control volume for the thermal hydraulic analysis. The area over which the hydraulic pressure acts corresponds to the area inside the wrapper minus the tube area.

Before applying the pressure time histories to the structural model, they are scaled based on a ratio of the plate area in'the structural model to the control volume area in the hydraulic model.

i o

7-6  ;

I I

I

}

)

Table 7-1 Summary of Component Materials Com m t Material l .

[

Channel Head SA 216 Grade WCC Tubesheet SA-508 Class 2a Shell SA-533 Grade A Class 2 Tube Support Plate SA-285 Grade C Stayrod SA-106 Grade B Spacer SA-106 Grade B Tube Inconel600 5

f 5

l 1

l 7-7 l

i l

i

Table 7-2 Summary of Material Properties SA-285, Gr. C ll TEMPERATURE CODE ED. 70 200 300 400 500 600 700 PROPERTY 71 27.90 27.70 27.40 27.00 26.40 25.70 24.80 Young's Modulus 71 6.07 6.38 6.60 6.82 7.02 7.23 7.44 Coefficient of Thermal Expansion Density - 0.284 0.283 0.283 0.282 0.281 0.281 0.280 7.35 7.33 7.32 7.30 7.28 7.26 7.25 I PROPERTY UNITS l t Young's Modulus psix 1.0E06 l i

Coefficient of Thermal infm/deg. F x 1.0E-06 ,

Expansion i

Density IbTm^3 lb-sec^2/in^4 x 1.0E-4 l

1 7-8 l

l

Table 7-3 Summary of Material Properties S A-106, Gr. B ll TEMPERATURE PROPERTY CODEED. 70 200 300 400 500 600 700 I

Young's Modulus 71 27.90 27.70 27.40 27.00 26.40 25.70 24.80 71 6.07 6.38 6.60 6.82 7.02 7.23 7.44 Coefficient ofThermal Expansion Density - 0.284 0.283 0.283 0.282 0.281 0.281 0.280 7.35 7.33 7.32 7.30 7.28 7.26 7.25 PROPERTY UNITS l Young's Modulus psi x 1.0E06 CoefHeient ofThermal in/in/deg. F x 1.0E-06 Expansion Density Ib/in^3 lb-sec^2/in^4 x 1.0E-4

=

7-9 l

Table 7-4 Summary of Material Properties SB-166 l TEMPERATURE l PROPERTY CODE ED. 70 200 300 400 500 600 700 Young's Modulus 71 31.70 30.90 30.50 30.00 29.60 29.20 28.60 Coemeient ofThennal 71 7.13 7.40 7.56 7.70 7.80 7.90 8.00 ,

i Expansion Density - -- 0.306 0.305 0.305 0.304 0.303 0.302

- 7.923 7.905 7.886 7.867 7.847 7.828 I PROPERTY UNITS l Young's Modulus psi x 1.0E06 Coemeient ofThennal in/in/deg. F x 1.0E-06 Expansion Density lblin^3 lb-sec^2/in^4 x 1.0E-4 l

7 - 10 l

I Table 7-5 )

Summary of Masenial Properties l SA-216, Gr. WCC i

ll TEMPERATURE PROPERTY CODE ED. 70 200 300 400 500 600 700 l I

Young's Modulus 71 27.90 27.70 27.40 27.00 26.40 25.70 24.80 71 6.07 6.38 6.60 6.82 7.02 7.23 7.44 Coefficient of Thermal Expansion Density - 0.283 0.282 0.282 0.281 0.280 0.280 0.279 7.324 7.303 7.287 7.269 7.252 7.234 7.215 l PROPERTY UNITS l Young's Modulus psi x 1.0E06 Coefficient of Thermal in/in/deg. F x 1.0E-06 ,

Expansion i

Density lb/in^3 Ib-sec^2/in^4 x 1.0E-4 ,

I

• l 7-11

+

Table 7-6 Summary of Material Properties SA-508, Class 2a l TEMPERATURE l PROPERTY CODE ED. 70 200 300 400 500 600 700 Young's Modulus 71 29.90 29.50 29.00 28.60 28.00 27.40 26.60 Cocmcient ofThermal 71 6.07 6.38 6.60 6 82 7.02 7.23 7.44 ,

Expansion Density - 0.283 0.282 0.282 0.281 0.280 0.280 0.279 7.324 7.303 7.287 7.269 7.252 7.234 7.215 I

i I PROPERTY UNITS I Young's Modulus psi x 1.0E% i I

Cocmcient ofThermal in/in/deg. F x 1.0E-06 l Expansion Density Ib/in^3  ;

lb-sec^2/in^4 x 1.0E-4 I

i l

2

'l 7 - 12 l

Table 7-7

+

Summary of Material Propertie SA-533, Grade A Cass 2 i

i ll TEMPERATURE  :

! PROPERTY CODEED. 70 200 300 400 500 600 700

- Young's Modulus 71 29.90 29.50 29.00 28.60 28.00 27.40 26.60  !

Cocmcient of Thermal 71 6.07 6.38 6.60 6.82 7.02 7.23 7.44 Expansion i Density - 0.283 0.282 0.282 0.281 0.280 0.280 0.279 7.324 7 303 7.287 7.269 7.252 7.234 7.215  !

l PROPERTY l UNITS l l Young's Modulus psi x 1.0E06 f

Coemcient of Thennat in/in/deg. F x 1.0E-06 Expansion  !

Density Ib/in^3 l lb-sec^2/in^4 x 1.0E-4 l

7 - 13

1 I

l i

Table 7-8 Summary of Combined Tsered / Spacer Sdffnesses OuterTiered / Spacer Up Loads a

~ -

1 i

i h

. - ]

i 7 - 14

Table 7-9 Summary of Combined Tiemd / Spacer S6ffnesses Outr11emd / Spacer Down Loads a 9

i i

i 1

l l

l 7 - 15 i

i 1

l

TaWe 7-10 Summary of Combined Tiemd / Spacer Sdffnesses Central Tiemd / Spacer

.l l

A ,

)

r s

f r

i l

s

~$ t m

d 5

1 7 - 16 ,

[

P

i Table 7-11 Summary of SpacerStiffnesses

_ _ a 7 - 17 i

Table 7-12 Summary of Equivalent Plate Pmperties

_ a 7 - 18

i i

l Table 7-13 i

Summary of Place Amas ,

A

~

Structural Area i

}

i l

l

- l l

l i

7 - 19 l

1 i

i 1

I l

Table 7-14

, Summary of Flow Areas a

l l

t I

1 l

1 1

1 7 - 20

Table 7-15 Summary of Effecdve Densides ,

SLA laidsting from Hot Standby ,

t k

I f

I s

I e

t i

F I

i

?

i une l

)

i

+

[

l 6

's

)

i

?

I n

1 7 - 21

i i

Table 7-16 l l

Summary of Effective Densities SLB Inisting from Full Power

- , a 7 - 22

1 1

Table 7-17 .!

i I.

Comparison of Naement Frequencies l Full Versus Reduced DOF  !

s I

r 5

- m 3 2 i

6 a

2 L

i i

6

?

1

't I

(

I f'

f 1

r I

i

-e 1

I 1

l 1

W eums ,

i i

7 - 23 1 i

Table 7-18 Comparison of Natural Frequencies Full Versus Reduced DOF With and Without Wedge Suppons M

l l

M M

7 - 24

a

=

l l

Figum 7-1. Tube Bundle Geometry 7 - 25

_a

=

Hgure 7-2. Tiered / SpacerImcatiens 7-25

\

i

)

. a

=

Mgure 7-3. Plate A (IH) Support I4cadom 7 - 27

t i

.i

(

f i,

3

& t i

r t

4 1

i a

.i i

'?.

i

)

r

,1 t

1 i

f i

i f

k i

i 9

f i

i i

1 1

Mguse 7-4. Pisee C (3H) Support Imanions 7-28

a 8

=

Figure 7-5. Plate F (SH) Support locations 7 - 29

a

=

Hgure 7-6. Plate J (7H) Support Locations 7 - 30

l 4

-a l

l 1

i t

l I

1 Figure 7-7. Plate L (8H) Support Locations i

7-31 l

l

a

, 2 l

Figure 7-8. Plate M (9H) Support Imentions 7 - 32 l

a

=

Figure 7-9. Plate N (10H) Support Locations 7 - 33

_. a

=

l l

Hgure 7-10. Mae P (11H) Support Locations 7 - 34

a l

1

l

- \

)

i i

i i

Figure 7-11. Configuration of Tiered / SpacerSupports 7 - 35 i

a Figure 7-12. Configuration of Vertical Bar/ Wedge Supports 7 - 36

I

-* Phi d

= =  ?

4

(

.? 4 i

!e t

e

.i

. n

!=

i j t a U

i .

t i:

r! i i

D

= = -

i.

a T

t k

l Figure 7-13. Summuy of Plate Rotation to Cause l Two-Edged Contact l 7 -37 ,

I 1

M V

} '

\

h i

1 p guit l'

• g,,gt Enite De*** g,6el Geon*M 7-38 i

l

=_. -

i 1

... ,...................,......s

. ...... ..i 2

...... ..~..

. ...e o

t, ..,

...... s.....................

1 / i i 1 ',

...,. yq -

'. .. +

'., . \\ *

\ N . .. . .-

/ ....

. L Figure 7-15. Mode Shape Plot - Plate A (IH)

Full Set of DOF Mode 1 7 - 39

..2

.............--n.....................-

~.

-.. .~-

1 r

.- =

-l Mgure 7-16. Mode Shape Mot - Plate A (1H)

Full Set of DOF Mode 2 7 - 40 l I

- - - l

-o i 4

l 1

1 l

.a.

... ... - .. / '

\,NX ... -

~.

. . .. I t

I 1

r h

Mgure 7-17. Mode Shape Plot - Plate A (1H)

Full Set of DOF Mode 3 7 - 41

1 i

......'./ ,................... ......

..... . .,. g

.....1

...... 1 .

w ..................._

....- t Figure 7-18. Mode Shape Plot - Plate A (1H)

Reduced Set of DOF t

Mode 1 .

l 7 - 42 l

t

l l

l h

i J

1 s .

... ~.. .

.i - 3 t_ f ,,

,/

-/. ,

~-.

s.

x, .... ... -

VN' .

=

Mgure 7-19. Mode Shape Mot - Mate A (1H)

Reduced Set of DOF Mode 2 7 - 43 F

i 1

)

1 i

)

...*., n ................... ....*, ,

....... .......... s.

3

... s 1

..... ..s I

l l

Figure 7-20. Mode Shape Plot - Plate A (IH)

Reduced Set of DOF Mode 3 i

7 - 44 l

i t

i

a l

l I

=

l l

Figure 7-21. Dynamic Degrees of Freedom )

Plate A (IH) l I

7 - 45  !

1

,i

)

i i

,l i

e a.

i a

?

i i-

\

L T

I b

1

( TO I

, 1 i

t Figure 7-22. Dynamic Degrees of Freedom Plate C (3H) l 1

7 -!

L i

a t

1

-E e

e i

1 4

f 1

.. f i

l Figure 7-23. Dynamic Degrees of Freedom Mate F (SH)  :

7 - 47

i i

a .

i I

i 6

t i

E I

i

=

i Figure 7-24. Dynamic Degrees of Freedom Mase J (7H) l l

l 7 - 48

l t

i a

~

I i

i I

k 4

i k

1 4

9 i

1 5

=

Hgure 7-25. Dynamic Degrees of Freedom I Plam L(SH) 7 - 49 ,

1 l

1 l

.1 l

i I

l 1

l 1

a i

1 1

1 1

)

II l

l

=

i name Figure 7-26. Dy==mic Degrees of Freedoen Plate M (9H) 7 - 50

l 1

4 l

l l

-I j

a l

F I

i L

r t

i i

d 1

1 f

?

i t

t i

Figure 7-27. Dynamic Degrees of Freedom Plate N (IDH) 7 - 51 1

i j

- -- - .~ _ _ . . _ , ^' _ _

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

a I

l l

Figure 7-28. Dynamic Degrees of Freedom Mate P (11H) 7 - 52 l

l l

8.0 TSP DISPLACEMENT ANALYSIS RESULTS 8.1 Analysis Approach Section 8 provides a summary of the displacement results for the limiting SLB load cases presented in Section 6. Those cases, based on the pressure time loadings for the support plates are as follows:

Case Initial Break Location Uncertainty T/H Code Conditions Factor 1 Hot Standby S/G Nozzle ---- TRANFLO 2 Full Power S/G Nozzle ---- TRANFLO 11 Hot Standby S/G Nozzle 2.0 TRANFLO 12 Full Power S/G N'ozzle 1.75 TRANFLO 31 Hot Standby S/G Nozzle --- MULTIFLEX The objective of the time history analysis is to define the number of tube locations where displacement amplitudes exceed 0.100 inch. These results are then supplied as input to the evaluation to establish burst probabilities as a result of tube support plate motions under SLB loads. In evaluating the displacements, the first step is to generate displacement time histories for each plate for all of the degrees of freedom (DOF). Using the displacement time histories, ,

maximum displacements and corresponding times are determined for each of the support plates. Full plate displacements are then calculated for the limiting plates at the critical times j and superimposed on a tube map for local regions of the plates . Displacements are calculated at each tube location within the local region by interpolating the displacements of )

the closest nodes, and placed into groups based on the amplitude of the displacement. The (

number of tubes with a given relative tube / plate displacement amplitude is then determined I and provided as input to the tube burst analysis.

It is the relative plate / tube displacement that is of interest, with the tube and plate positions at the start of the SLB transient defined as the reference position. At hot standby, the TSP positions relative to cracks inside the TSP are essentially the same as at cold shutdown.

l Every known SG cold condition inspection shows ODSCC cracks within the non-dented TSP with a trend towards being centered within the TSP. Therefore, the cold condition TSP location relative to the tubes is essentially the same as for the full power condition where the cracks formed, which is also the position during hot shutdown. These inspections indicate that there is little relative movement between the tubes and plates throughout the operating cycle. Thus, thic analysis calculates relative tube / TSP motions based on the tube / plate positions at the initiation of the SLB transient.

8-1

Calculations are also performed to demonstrate the applicability of the elastic analysis approach in determining the resulting displacements. These calculations consist of showing that the tierods / spacers remain elastic throughout the transient, that significant yielding of .

i the tube support plates does not occur, and that the welds joining vertical bars (that provide vertical restraint for the plates) to the partition plate and wrapper remain intact.

8.2 Summary of Limiting Plate Displacements As discussed above, several sets of SLB loads have been considered. An overall summary of the limiting displacements for each of the plates is provided in Table 8-1. Three sets of displacement results are presented. The first set, the top set in Table 8-1, are the plate displacements relative to their initial installation (zero displacement) position. The second set of displacements, the middle set in TaHe 8-1, are the relative plate / tubesheet displacemenis (i.e. plate - tubesheet, since the tubests t displacements also represent the tube displacements at the TSP elevation). The third and final set of displacements, shown at the bottom of Table 8-1, are the change in the relative plate / tubesheet displacements from the start of tne transient. It is the final set of displacements that represent the relative plate / tube motion and -

the potential for uncovering cracks that could potentially exist in the tube inside the TSP.

Displacement time histories for each of the SLB cases, showing the displacement time history for the limiting location for each plate, are provided in Figures 8-1 through F-10. Two plots are provided er each transient. The first plot shows the response of the bottom four plates, Plates A(IH), C(3H), F(5H), and J(7H), while the second plot shows the response of the top four plates, Plates L(8H), M(9H), N(10H), and P(1IH). These plots show the change in relative plate / tubesheet displacements from the initiation of the transient. The analysis results show Case 11 (SLB initiating from hot standby with the break at the steam generator nozzle and an uncertainty factor of 2.0 applied) to be limiting. The results further show Plates A(IH), C(3H), and J(711) for Case 11 to experience the largest deflections. The results for the other plates are judged to not be .ngnificant relative to the limiting plates.

In comparing the results for the TRANFLO and MULTIFLEX cases, Cases 1 and 31, the MULTIFLEX case is found to result in somewhat higher displacements. Although the pressure loads were close in terms .findividual plate loads, the different frequency content of the loads results in displacements sat are somewhat higher for the MULTIFLEX case.

In reviewing the results for Case 11 relative to the other cases, it n notd for Plate C(3H) that the displacement for the limiting location does not retum to its inini sto; ting position at the end of the transient (or nearly so), as for the other plates. This is a result of the plate interacting with a tube. as discussed in Section 7.3. The response si the plate is such that once tube / plate interaction occurs due to local plate rotatior.s. the tube and plate remain in contact throughout the remainder of the transient. The nature of the overall plate response is shown in Figures 8-11 and 8-12, where the time history response of all of the Plate C(3H)

DOF is shown. (In conjunction with th.e ilots, nodal locations for Plate C(3H) are shown in I

1 8-2

~

l l

1 Figure 8-13.) Yhe tube / plate interaction occurs at node 736.' Referring to Figure 8-11, it is ,

observed that nodes adjacent to the tube contact location, Nodes 797 and 817, reach their j maximum deflection shortly after tube contact occurs, and begin to return to the unloaded position, which increases the local plate rotations, maintaining the plate / tube interaction. A plot of the local plate rotations is shown in Figure 8-14, and shows results consistent with the displacement pattern of Figure 8-11. Thus, once tube / plate contact is initiated, it remains throughout the remainder of the transient for this particular loading and plate geometry.

For the limiting transient, Case 11, plots are also provided for Plate C(3H) showing the plate displacements relative to the installation position in Figures 8-15 and 8-16. In order to compare the plate and tubesheet motions, the limiting location for Plate C is shown along with the corresponding tubesheet location in Figure 8-17. Also shown in Figure 8-17 is the differential plate / tubesheet motion. Note that each of the displacement curves in Figure 8-17 are relative to their position at installation. Finally, a plot showing the tubesheet motions for Case 11 at three tubesheet positions, center, mid-radius, and outer radius, is provided in Figure 8-18 to show the change in tubesheet displacement in going from the plate center to its outer radius.

The limiting plate displacements in all cases occur over a small region of the plate located near the tube lane at the outer edge of the plate, where the distance between vertical supports is greatest. Displaced geometry plots for Plates A(lH), C(3H), and J(7H) for the limiting set of loads, Case 11, are shown in Figures 8-19 through 8-21. The consistent displacement pattern is apparent for each of the plates. .

8.3 SLB Displacements By Tube Location In order to establish probabilities for tube burst as a result of relative plate / tube movement, calculations are performed to determine how many tubes are associated with a given displacement magnitude for a given plate. The plate displacements are categorized into groups, starting at 0.10 inch, and increasing in 0.05 inch increments to a maximum displacement > 0.55 inch.

1 The calculations to estimate the SLB tube burst probabilities are based on the change in the relative plate / tube position from their positions at the initiation of the SLB transient. The ODSCC indications are formed on the tube within the TSP at normal operating conditions.

Relative to cold shutdown conditions, the TSP is displaced (relative to a fixed location on the tube) by the net effect of the secondary fL s pressure difference across the plate plus the bow of the tubesheet. (Thermal expansion effects also marginally influence the TSP / tube displacements relative to cold shutdown, but these effects are negligible for SLB ,

i In referring to Figure 8-11, it is observed that Node 717 has a higher deflection than Node 736, and would also be expected to interact with tuves: Node 717, however, is located at the outer edge of Plate C(3H) and there are no tubes adjacent to this node. Node 736 represents the outermost node l

location that is adjacent to a tube. Thus, Node 736 was selected as the site for potential plate / tube j interaction

)

8-3 l 1

displacements at normal operating conditions.) The tubesheet bow displaces all axial locations on the tubes by the amount of the bow, while the TSP displacements closely match the bow of the tubesheet only at the locations of the tierods. The secondary flow pressure difference across the TSP tends to displace the p' tte in the upward direction relative to the j tube. The net displacement of the plate is the sum of the tubesheet bow interaction through  :

I the tierods and the pressure differential. Therefore, the movement of the plate relative to the tube is the difference between the net plate displacement and tubesheet bow. For the Model D4 SG, the relative plate to tube displacements at normal operating conditions are not large, typically s 0.06 inch, for the location with maximum SLB displacement. The relative displacements for hot standby conditions are not significantly different, which indicates that the full power TSP displacements due to the pressure differential across the plate are not large. Thus, it would be expected for the Model D4 SGs that the indications would be inside l the TSP at both cold and hot conditions, independent of whether or not the plates are ,

effectively clamped to the tubes as a result of crevice deposits. The net SLB displacement of l the plate relative to the ODSCC on the tube is then the change in relative plate to tube (or i tubesheet) displacement between a time in the SLB and time = 0.

The algorithm for calculating the relative dis'placements is as follows:

AD = (D, . 7 - D, . o ),w - (D, . 7 - D, . o ) row.w , where Dew = Plate Displacement Dr,w,w, = Tubesheet Displacement T = Time of maximum displacement from dynamic analys s m order to calculate the relative displacements across the full plate, displacement (stress) solutions are performed for the limiting plates at the times of maximum displacement.  ;

The displacement solutions are performed using the finite element representations for the '

plates. Displacements for the dynamic degrees of freedom for the limiting plates (and l tubesheet) are extracted at the times ofinterest from a file containing the DOF displacements  !

for the full transient. These displacements are applied to the finite element model as boundary conditions (along with any other appropriate boundary conditions representing i symmetry or ground locations), and displacements for the entire plate (and tubesheet) are ,

calculated. Using the algorithm above, the relative plate / tubesheet displacements are 2

calculated.  ;

i The next step in calculating displacements at individual node locations is to determine where ,

in the model node layout each tube is located. Shown in Figure 8-22 is an overlay of the tube pattem on the finite element grid for the critical plate region. Once the four nodes ,

surrounding a given tube location is known, the displacement at the tube location is calculated i by interpolating the nodal displacements using the following algorithm.

8-4

n y' - y* a 3, - d, +

7_7 (dc - d,)

,r C B;

' r y' - y^ r d2 = d, +

7 _7 A>

(do - d,)

3D

'ry - x',

d=d+

r 2 _ (4 -d) 3 where, A, B, C, D correspond to the four element nodes X ,, X3 ,. Y,, Y3, correspond to the coordinate locations of the nodes  ;

X,Y,.

1 7 correspond to the coordinate locations of the tube  ;

da, do, . correspond to the displacements at the four corner nodes .

d i , d 2, are the interpolated displacements in the Y-direction at the tube location dicorresponds to the final interpolated tube displacement As an example of the calculation methodology, sample results are provided for Plate C(3H) for the limiting set of SLB loads. Shown in Table 8-2 are the nodal displacements at time = 0.0, and at the limiting transient time,2.600 seconds, for Plate C(3H) and for the corresponding tubesheet nodes. Also shown in the far right hand column are the relative '

plate / tubesheet displacements with respect to time = 0.0, calculated using the above algorithm. A summary of the nodes associated with selected tube locations is provided in l Table 8-3. Referring to Table 8-3, on the left hand side of the table is the row and column number of the tube in question, and its coordinate location relative to the plate center. Next, for each of the four nodal locations surrounding the tube position are the node number, nodal i

coordinates, and the relative plate / tubesheet displacements from Table 8-2 (Note that C1, C2, C3, C4, in Table 8-3 correspond to the four element nodes A, B, C, and D in the above algorithm.) Finally, shown in Table 8-4 are the interpolated tube displacements for tube rows i 1 and 2 in the vicinity of the limiting plate displacement, and the categorization of the tube ,

displacements into the appropriate displacement group. In Table 8-4, di and d2 correspond to the intermediate interpolated displacements, with dr being the final interpolated displacement.

Similar calculations are performed for all of the tube locations in the vicinity of the maximum plate displ; cements, and an overall sum of the number of tubes in each displacement grouping is obtained. '

A summary of the number of tubes falling into each of the displacement groupings for the limiting plates for each of the load cases is provided in Table 8-5. Note that the numbers of tubes in Table 8-5 correspond to the full plate. The number of tubes in each plate quadrant is one-half of the values listed.

l 8-5

Summarized in Table 8-6 is a comparison of the maximum plate displacement to the plate j displacement at the limiting tube location (the tube having the highest displacement), RICl. l As can be observed in the displaced geometry plots in Figures 8-19 through 8-21, the  !

displacement gradients at the comer of the plate are high, so the maximum differential  !

displacement at RICI is less than the madmum plate displacement reported in Table 8-1.

In order to assess the influence of plate / tube interaction on limiting the plate displacement, one additional case was evaluated using a set ofloads that bounded all of the prior cases.

Using the loads for SLB from hot standby with a break at the steam generator nozzle as a reference, a factor of 3.0 was applied to the loads to arrive at loads for the overload condition. Thus, these loads are 50% higher than the loads for the limiting case, Case 11.

Comparing the displacements for Case 11 and for the overload condition shows only a small increase in the maximum displacements for Plate C(3H), where tube / plate interaction occurs.

The maximum plate displacements for the two cases are [ ]" for Case 11, and

[ ]" for the overload conditier which is only a [ ]" in the maximum plate displacement. Thus, once tube / plate contact occurs, any additional plate displacement is significantly limited.

l l

8.4 Summary of Stress Results Since the dynamic analysis is based on elastic response, calculations were performed to assure that the tierods, a significant support element for the plates remain elastic throughout the transient. The dynamics analysis results establish that the tierods do,in fact, remain elastic throughout the transient. The 0.2% offset yield point corresponds to an elongation of the outer tierods (that run from the tubesheet to the top TSP) of [ ]" inch, and an elongation of the central tierod (that runs from Plate L(8H) to Plate P(llH)) of [ ]" inch. The maximum elongations calculated during the limiting SLB (Case 11) are [ }" inch for the outer stayrods, and [ ]" inch for the central stayrods. In both instances, these elongations are well below the yield point for the stayrods. Similarly, for the spacers that are located on the outside of the tierods between the support plates, the maximum compressive stress in the tierods is on the order of [ ]" psi, which is well below the yield stress of 23,400 psi.

Also relevant in assessing the appropriateness of the elastic solution, are the stresses in the plates. Thus, in conjunction with the displacement results from the dynamic analysis, stresses are calculated for the hot leg plates at the times corresponding to the maximum plate displacements. The stresses are calculated by extracting displacements from the dynamic analysis for each plate degree of freedom, and tien applying those displacements to the finite element model. The finite element code then back-calculates the displacements and stresses for the overall plate model.

Additional boundary conditions corresponding to lines of symmetry and appropriate rotational constraints are also applied to the model. The finite element results give a set of displacement and stress results for the overall plate. The resulting plate stresses, however, correspond to the effective Young's modulus, and must be multiplied by the inverse ratio of effective-to actual Young's modulus to get the correct plate stresses.

8-6

Stresses are calculated for Plates A(IH), C(3H), and J(7H) for the limiting set ofloads. In order to interpret the stress results, stress contour plots for the maximum and minimum stress intensities have been made for each plate. Plots showing the maximum and minimum stress intensities are shown in Figures 8-23 to 8-28, for Plates A (lH), C (3H), and J (7H) respectively. These plots show the distribution of stress throughout the plate. As expected, the maximum stresses occur near the locations of vertical support, the tierod / spacers and vertical bars. The ASME Code minimum yield strength for the TSP material is 23.4 ksi. The resul;s for Pl. ate A show the stresses to be elastic throughout the plate. For Plates C(3H) and J(7H), there are local areas near the tierods, where the surface (bending) stresses slightly exc:ed the yield stress.

The plate stresses cannot be compared directly to the material yield strength, as these stresses correspond to an equivalent solid plate. In order to arrive at the plate ligament stresses, additional detailed stress analysis of the plates is required. Such an analysis is outside the scope of this program. The equivalent plate stresses do provide a general guideline as to those areas of the plate that are most limiting from a stress viewpoint. The plate stresses are meaningful in that they indicate that the stresses are generally low throughout the plate, and l that the elastic analysis is a good approximation of the transient plate response. For i Plate C(3H), which experiences the highest stresses, local yielding of the plate near the tierods will not lead to a significant change in the maximum displacements as they are limited by the tube / plate interaction.

Calculations have also been performed to determine the stresses in the welds between the venical bars and the partition plate and wrapper. The loads at the various support points are extracted from the static WECAN runs in the form of reaction forces at the times of maximum plate deflection. Loads have been extracted for the limiting plates (based on plate motions) for the limiting set of SLB loads. Calculations are also performed for the limiting plate above the preheater, Plate P (llH), as these plates have a somewhat different support arrangement than the lower plates.

The vertical bars are generally [ ]' inches in length, except for the bars tmderneath Plate A(lH), which are only [ ]' long. The bars are welded to the partition plate and / or wrapper using a full length [ ]' fillet weld along both edges of the bar. The stresses in the welds act in the shear direction. The weld throat area is simply the throat width times the height of the weld [ ]'. The corresponding stress intensity is twice the shear stress. _

l A summary of the reaction forces and corresponding stresses for each of the bar locations for the locations considered is provided in Table 8-7.. The results show all of the stresses to be low (<3 ksi) for a faulted event. The allowable stress for the welds is based on 2.4S,, x 1.5 x 0.35 (for fillet welds with visual examination) for carbon steel. S,, at 550"F is 15.5 ksi. The resulting allowable stress intensity is 19.53 ksi, and the weld stresses are acceptable.

One final set of calculations was performed for Case 11 to determine the stresses induced in the tube as a result of the tube / plate interaction. From the time of initial tube / plate 8-7

contact, the plate deflection increases by [ ] This also corresponds to the axial strain induced in the tube. The corresponding axial stress in the tube is [ ] The maximum primary-to-secondary pressure differential for the SLB event is on the order of

[ ] This corresponds to a tube axial stress of [ ] The combined tube stress is [ ] The tube yield stress at 650"F is 35 ksi. Thus, the tube remains elastic during the transient event. It should be noted that this analysis assumes contact with a single tube. However, it is likely that more than one tube would contact the plate, distributing the load, and potentially reducing the plate deflections.

Overall,it is concluded that the elastic analysis provides a good approximation of the dynamic response of the TSPs to the applied loading.

i l

8-8 l

l

1 l

l Table 8-1 Summary of Maximum 1SP Displacements for Postulated SLB Events for D4 Steam Generators I

a,c ;

~

ActualPlate Displacements k

h t

f f

'; i l

v l

i 8-9 l

?

Table 8-2 Summary of Relative Plate / Tubesheet Displacements Model D4 Steam Generator SLB initiating From Hot Standby Break at Steam Generator Nozzle Plate C a,c p

8 - 10

l l

Table 8-3 i Summary of Nodal Displacements Model D4 Steam Generator l SLB Initiating From Hot Standby Break at Steam Generator Nonle M are C a,c

=

l l

l 1

8 - 11

l l

i Table 8-4 Summary of Retsiive Plate / Tube Displacements Model D4 Steam Generator SLB Initiating From Hot Standby Break at Steam Generator Nozzle Mde C 8,C 1

r T

" 8 - 12 -

TaNe 8-5 Summary of Number of Tubes Having Different Displacement Magnitudes Model D4 Steam Generator Steam Une Break Imad Cases a,c i l

l 3

i

^5 _

l l

t 8 - 13 I

I

l TaWe 8-6 Comparison of Maximum Displacement at Plate Edge and at umiting Tube Location a,c

?

4

• ~~

8 - 14 i-

t Table 8-7 1

Sumniary of Vertical Bar Stresses e Model D4 Steam Generseers a,C p

i L

,e I

I e

S .

L I

8 - 15 ,

I

a,c f

'- g i

P

=

l Figure 8-1. Re:sthe '"Au t /1 ibesheet Displacement Time History Response SUI from Hot Standby  ;

Mr*.r.k 6 Steam Generator Nozzle l Pamies A(IH), C(3H), F(5H), J(7H) {

8 - 16 I

a,C i i

t 5

Figure 8-2. Reistive Plate / Tubesheet Displacement Time History Response SLB from Hot Standby Break at Steam Generator Nozzle Plates L(8H), M(9H), N(10H), P(11H) 8 - 17

i l

a,c k

i Figure 8-3. Relative Plate / Tubesheet Displacement Time History Response i SLB from Full Power Break at Steam Generator Nozzle Plates A(1H), C(3H), F(5H), J(7H) 8 - 18

a,c l

i Figurt 8-4. Displacement Time History Response SLB from Full Power Break at Steam Generator Nonle Plates 148H), M(9H), N(10H), P(11H) 8 - 19

a,c 1

)

4 l

2=

Figure 8-5. Relative Plate / Tubesheet Displacement Time History Response SLB from Hot Standby Break at Steam Generator Nonle Uncertainty Factor of 2.0 Plates A(1H), C(3H),, F(5H), J(7H)-

8 - 20

a,c

~

'b 1

h c:

i i

Figure 8-6. Relative Plate / Tubesheet Dispiscement Time History Response i SLB from Hot Standby Break at Steam Generator Nozzle Uncertainty Factor of 2.0 Plates L(8H), M(9H), N(10H), P(11H) 8 - 21 i

i

a,c i

I f

Figurt 8-7. Relative Plate / Tubesheet Displacement Time History Response SLB from Full Power Break at Steam Generator Nonle Uncertainty Factor of 1.75 Plates A(1H), C(3H), F(5H), J(7H) 8 - 22

' a,c l

l i

3 r

E i

Figure 8-8. Relative Plate / Tubesheet Displacement Time History Response SLB from Full Power  ;

Break at Steam Generator Nozr.le ,

Uncertainty Factor of 1.7f  !

Plates 148H), M(9H), N(10H), P(11H) 8 - 23 i

i a,c z  :

Figure 8-9. Relative Plate / Tubesbeet Displacement Time History Response ,

SLB from Hot Standby l Break at Steam Generator Nozzle MULTIFLEX Loads ,

Plates A(1H), C(3H), F(SH), J(7H) l 8 - 24

/

a,c l

a r

4 3

5 E  !

- i Figure 8-10. Relative Plate / Tubesheet Displacement Time History Response SLB from Hot Standby Break at Steam Generator Nozzle MULTIFLEX Loads Plates I48H), M(9H), N(10H), P(11H) 8 - 25

1 I

l a,C

~

P 1

r i

3 ,.

i l

Figure 8-11. Relative Plate / Tubesheet Displacement Time History Response SLB from Hot Standby Break at Steam Generator Nozzle Uncertainty Factor of 2.0 Plate C(3H) Degrees of Freedom 8 - 26

a,c l

t

-f.

Figure 8-12. Relative Plate / Tubesheet Displacement Time History Response SLB from Hot Standby l Break at Steam Generator Nozzle l Uncertainty Factor of 2.0 Plate C(3H) Degrees of Freedom 8 - 27

1882 1843 magg Y  % k iort

.. ,m, .ms = im<

1_ %, t 010

, m - n., - -, - ,

-i m - -, na - nn .,2 m, m, ms no 5

m n,e n,e min its m ti su m,i at, an a n,< nw I

am ans ano sin mi t ni, att 'atAel5 an. nne nrw ani am ani h

• ELgK

m. i n., m , , , _ - -, - n.n -, n., =4 u.%

d,.

n.. ., no N. y n., - n,n .,, m m .,,

, un un n.n att no an ata att t$6 utu, ut u. ut na.

,, r,a n,e asn asi att m nu nu 17 asi m m nu n,e rw am an. an, nna ano nin non mit att mia att dis 31 ani an, not and so, yet Tea ,e4 7%

9 va, vna v.n ve t

?mi,e, vat vna vet

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

l

..n ,,1 .., ,,, . ... . L, 71, 714 ,11 fia f 799,,, 7,1  ?,A T,4 PM  ?*, 7,8 t,0 FSA 711

~ = n2 fit un.EyY mm mm l,

n Hgure 8-13. Mate C(3H) Node Numben i

8 - 28 t

a,c 1

l

? -

Figun 8-14. Plate C(3H) Local Plate Rotations

• SLB from Hot Standby Break at Steam Generator Nonle Uncertainty Factor of 2.0 8 - 29

R.C t

6 t

i l

l l

3 Figure 8-15. Relative Plate / Tubesheet Displacement Time History Response SLB from Hot Standby Break at Slam Generator Nozzle 1 Uncertainty Factor of 2.0 Displacements Relative to lasta!!ation Position Plates A(IH), C(3H), F(5H), J(7H)  !

8 -30

i a,c t-l Figure 8-16. Relative Plate / Tubesheet Displacement Time History Response SLB from Hot Standby Break at Steam Generator Nozzle Uncertainty Factor of 2.0 Displacements Relative to lastallation Position Plates L(8H), M(9H), N(10H), P(11H) 8 - 31

i e

i a,c F

t b

k 1r i

Figure 8-17. Comparison of Plate and Tubesbeet Displacement Time Histories  !

SLB from Hot Standby Break at Steam Generator Nozzle .

t Uncertainty Factor of 2.0 Plate C ,

t 8 - 32

,. a.C i

{

i r

5 Figure 8-18. Tubesheet Displacement Time Histories as a Function of Tubesheet Radius SLB from Hot Standby Uncertainty Factor of 2.0  ;

Break at Steam Generator Nonle 8 - 33

. s. - .

• ~ - - ...

x xw--............-- --...__ .

-...' ~f .... ,,.. .

.- N N N % .

xx xx x

. . . . l T., '., , N N ... , ' -.  ;

l . ., x N xN xN x .,.,  :

. , x ... '

x x x x N . .- ..

'.'=x  :.,'. .

x y mx xx..v.x

'. '. \

~

\ x x x b_ .

. . . . s

. ~ x '

. A .,

. s. .

. .- ... s.s..

s

...1 Figme 8-19. Displaced Geometry Plate A(IH) : Tame = 2.572 see SLB from Hot Standby ,

Break at Steam Generator Nozzle Uncertainty Factor of 2.0 8 - 34

~. .

...=,,,,,..........-.......,,,,

..,*~.. . ,

.. ... .. .. =.*v.

- x s

~

1 m ....., 3 .

' . x x N' '

, xxxx - ..- -

. x .x x x x ... - . ,

, .x x x . .N '

N . \. . \ x \

'. \ g N N x \x \\ '. '.

e x x xx \

a e

', -- x- x -

..,........ ......s.. .

+

e

.. s. .

- ' . , .e e

'. g g

i

......- ..... ...... ...3

. s.

u;

~7. 1

e. :W

?

Figure 8-20. Displaced Geometry Plate C(3H) : 'nme = 2.600 see SLB from Hot Standby Break at Steam Generator Nozzle Uncertainty Factor of 2.0 8 - 35

- ...:-........ ...........~

xW -

x x x x x Y - .

.v < x x

x x

~

x- x x .

'x

'Q x_-N ,N N N \ *.

. v - x. 'x'x x x x x. .

%N \ ., i s.

x x .

s s N ',

~.

\ ..... .- .

.\.<....................

Figure 8 21. Displaced Geometry Plate .K7H) : Tiene = 2.656 see SLB front Hot Standby Break at Stemma Generuser Nozzle Uncertainty Factor of 2.0

'8 - 36

4 l

l I

,,, ,,, . Node Number CO C OO i., O000000000 000C 00 ,

,,, 20 0000000000 000C 00 0000000000 0000 000 m ggOgggoggg OggC ggg vvvvvvvvvv vvvv vvv g,,, '

0000000000 0000 000 15 0000000000000C 00000000000000 000 3 000 3 3{ ,  ;

0000000000 000C 000 30\m .8 i

.i 0000000000 OOOC 000 0000000000 DOOC 000 30 30 \ ,, 10 Z l i O000000000 OOOC 000 30 \ g ;

0000000000 000C 00000 ) rc

,i OOOOOOOOOO OOOC 00000 "

UUC >UUUI )UUU UUUC 000' )O0 0000000000 OOOC 000 300- -5

,, O000000000 OOOC 000 300 ,,

UU(7UUVIJUUU UUUU UUU JUU 0000000000 000C 000 300

,, 00000000000000C 000 300 - -1 ii l o lo u l i, I i.

l, i i.

20' l'5 l0 5 1 Column Number Esc Nede nuenbers eartsspond to lhte A Plase so-plate nede inereenent is 700 Figure 8-22. Tube Position Relative to Model Nede Location 8 - 37

l R,C l

l 2: _

M gute 8-23 Maximum Stess Intensity SIE from Hot Standby Break at Steam Genemtor Nozzle Uncertainty Facesr of 2.0 Place A(IH) 8 - 38

1

\

a,c

)

1 I

i l

l I

s ~~

Figme 8-24 Minimen Stress Innensity SLB from Hot Standby Break at Steam Generuser Neale Uncertnoty Factor of 2.0 Plate A(1H) 8 - 39

)

a,c 5

a Mgwe.8 25 Maximum Stress Intensity SLB from Het Standby Break at Seemn Generator Nozzle Uncertainty Factor of 2.0 Mase C(3H) a 40

\

1 R,C l

i I

t i

3 l Mgme 8-26 ,

Minimum Stress latensity SLB from Het Standby Break at Steam Generseer Nozzle l Uncertainty Factor of 2.0 i Plate C(3H) 8 - 41 i

l

I J

l A,C s

Figure 8-27 Maximum Stress latensity SIE from Hot Stmodby Brec.k ri Steam Generator Nozzle Tmertainty Factor of 2.0 Mate 47H) 8 - 42

1,C

~

\

1 1

I i

t 4

_= _

Hgme 8-28 Minimum Stren Intensity SIE from Hot Standby Break at 6 teen Generator Nozzle Uncertainty Factor of 2.0 l

Pisee J(7H) 8 - 43 I

i I

)

l I

9.0 SUI a TY OF BRAIDWOOD-1 AND BYRON-1 INSPECTION RESULTS The limited TSP displacement analyses of this report are applicable to all Model D4 SGs. The locations of significant displacements are generally limited to the comers of the TSPs adjacent ll to the tubelane for plates 3 and 7 (3H and 7H). Due to the small number of tube locations f subject to significant TSP displacements, it is useful to assume the frequency and size indications as a function of the TSP locations and elevation. These data are provided in this

-l section for the 1994 inspections at Braidwood I and Byron 1.  :

' 9.1 Braidwood Unit i 1994 (EOC 4) Inspection Summary Interim Plugging Criteria were applied for the first time to the bobbin inspection results during l

the 3/94 (AIR 04) inspection of the Braidwood 1 SGs; 100% full length inspection was l

performed, and all indications >l volt in flaw signal amplitude were subjected to l

re-examination with 3-coil RPC probes, as were those oflesser amplitude. RPC testing was performed to assess the consistency of the underlying tube condition with prior cases of TSP ODSCC, to support the indications as ODSCC within small dents, that the indications were within the TSP, and to confirm the absence of circumferential indications, as well as to l determine the extent of the indication with respect to repair criteria. He RPC inspection l l

confirmed that all indications were axial ODSCC located within the TSPs. The total num  ;

of TSP intersections RPC tested on the basis of possible ODSCC indications was 2775, distributed among the 4 SGs. The TSP bobbin indications confirmed by RPC testing numbered I 1582,486 in SG-A,76 in SG-B,642 in SG-C, and 378 in SG-D, representing a 57% rate of confirmation 'of the bobbin. calls. It is considered that 'only those intersections which exhibit {

detectable ODSCC with pancake coil inspections warrant scrutiny with respect to plugging i

criteria, whether under Tech. Spec. criteria or under the alternate basis represented by the Interim Plugging Criteria.  !

The distribution of the TSP ODSCC indications among the four SGs for the 1994 inspection is shown in Table 9-1, which tabulates the number of indications for each TSP elevation for l which indications were observed. For the D4 SGs of Braidwood I and Byron 1, the IH level l I

represents the Flow Distribution Baffle (FDB), a plate with oversize tube holes and no flow holes; for this censon the incidence of ODSCC is. expected to be low in the absence of unusual circumstances. In fact, none of the indications reported in Braidwood 1 occur at the FDB l l-elevation. He support levels above are numbered in the cold leg order; i.e., the next hot leg l l TSP is designated 3H since its height corresponds to the 3rd preheater plate. He remaining i

TSPs are designated SH,7H 8H,9H,10H, and llH. Thus though some probability of encountering ODSCC signals at the upper plates exists, it is expected that most of the indications will be observed in TSPs 3H, 5H, and 7H, since these levels are in the relatively ,

,, au

. n.,n .u u., e m_.... 9_1 ]

I l

hotter internal temperature zone of the tubes, maximum in the tubesheet and decreasing with elevation up to the apices of the U-be.As and thereafter decreasing as the cold leg elevation decreases.

Figure 91 presents an histogram of the numerical distributions of TSP ODSCC with respect to elevation. The bobbin amplitude distributions associated with the ODSCC indications are presented together with the cumulative distribution curves in Figure 9-2 for the composite of all 4 SGs. Figure 9-3 gives the RPC confirmation fraction of the bobbin indications as a function of bobbin voltage for each of the voltage bins; as expected the probability that the RPC probe will detect degradation increases with the bobbin voltage, which increases with the depth and length and number of cracks present. Table 9-2 provides detailed RPC confirmation statistics as a function of bobbin voltage for each of the individual SGs, as well as cumulative confirmation data for the 4 SG composite results.

9.2 Byron Unit 1 1994 (EOC 6) Inspection Summary The Cycle 6 refueling outage SG inspection encompassed 100% of the tubes, inspected full length with bobbin EC probes, and was performed under the IPC guidelines for calibration, identification, and measurement of signal amplitudes. RPC examination of all bobbin indications reported with amplitudes >l volt, as well as many smaller signals, was performed to permit disposition as required, i.e., repair of confirmed indications >l volt and all indications >2.7 volts. The population of bobbin indications observed in Byron I was consistent with the findings at Braidwood I; Figure 9-4 provides the distribution of the bobbin indications as a function of voltage. A slightly higher number of affected TSP intersections was seen - 3075 for Byron 1 vs. 2775 for Braidwood 1. Table 9-3 summarizes the Byron I results, from which it can be seen that of 967 indications which were RPC tested,717 were confirmed, a 74% confirmation rate. This compares with the 57% rate reported from Braidwood 1. Figure 9-5 shows the Byron 1 RPC confirmation rate for bobbin indications re-examined with RPC probes. The difference in confirmation rates is attributable to the much larger number of < 1 volt indwations which were RPC tested at Braidwood I, where all TSP flaw indications were RPC inspected. Comparing the confirmation rates for indications >l volt, Byron I had an 81% rate while Braidwood I saw a 91% rate. The distribution of TSP indications by distance from the hot leg tubesheet for Byron I is similar to that observed in Braidwood 1 (see Figure 9-6); 96% of all reported TSP flaw indications at Braidwood I were found in the first 3 TSPs, compared to 95% at Byron 1 (Table 9-4). Again no indications were reported at the elevation of the FDB (lH).

ei-. . v. >>..

. u ,,m...m ., .n m_.. .. 92

9.3 Indications at Regions of Significant TSP Displacement 9.3.1 Braidwood Unit 1 Tubes with ODSCC indications at TSP locations which have SLB dicplacements 2 0.35" are limited in number. Based on the results of section 8, the largest TSP displacements occur at TSPs 3 and 7 for a SLB at hot standby with uncertainty adjustment factor applied to the loads.

The displacements from the analysis can be used to identify indications found in the inspections at tube locations having significant (2 0.35 inch) TSP displacements. In Braidwood 1, SGs A and D have the most indications near the edges of the plate with maximum SLB displacements. Twenty-eight (28) out of 1590 indications at 3H, as summed over all four SGs, are located at positions with tube displacement between 0.35" and 0.507" (the maximum displacement at a location with an indication). Figures 9-7 and 9-8 show the 3H elevation distributions of bobbin indications for the Braidwood 1 SGs. The largest amplitude (> 4 volts) TSP indications for Braidwood I are plotted as a composite on Figure 9-9, and the corresponding voltages recorded for these locations are given in Table 9-5; 3 of these indications were found in the regions of significant TSP displacement, but none had a local TSP displacement 2 0.35"

~

9.3.2 Byron Unit I To review the displacement analysis results versus indication locations, the 3H and 7H indications for SG-A and for SG-C were illustrated in individual maps on Figures 9-10 and 9-11. It is seen that for SG-A, which has 416 indications at 3H, only 27 indications are situated in the comer areas at locations predicted to have 2 0.35" displacement, while none of the 96 indications at the 7H level are at such locations. Similarly at the 3H level in SG-C, the positions of predicted displacement 2 0.35" account for 21 of the 604 indications, while at the 7H level of SG-C there are no indications at these locations. The maximum TSP displacement for a tube location with an indication at Byron-1 is 0.522 inch. Figure 9-12 provides a composite plot for the four SGs of the locations of the largest amplitude (> 4 volts) TSP indications, and Table 9 6 provides the specific data for each of these indications. In Byron 1, only 1 of these tubes was found in the regions of significant TSP displacement, and has an associated local TSP displacement of about 0.47".

9.4 Conclusions The TSP ODSCC behaviors in the Braidwood I and Byron 1 SGs appear to be little different.

The numbers of tubes found with large voltages in regions of significant local TSP displacement are small in both plants.

, _ , ,, m .

. n. m .. m . m . n m _... m 93

Table 9-1 Braidwood #1 TSP ODSCC Indications (A1RO4) March,1994

= o.nn A se. o.n. a sm.= t 2 ,c si o.nw.in o An s=== o. = .

Volta Voets Otowth Volte volte Otowth Volts Vetts Growth Voets Volte Groudh Volte votes Growth s Idem. Ave. volte s Idem. Ave, weite s 14an. Ave. volte s Man. Ave. weite s tema. Ave. weite TSP 0 0 0 0 0 1H 4 90 0.88 0 37 134 4.25 0 88 0 13 700 2.38 0.72 0.21 403 8 82 0 82 03 1980 8 82 0 78 0 27 3H 344 v

238 8 33 0 95 05 116 2.75 0 50 0.23 235 2 46 0.82 0.16 100 10 44 0.75 0 28 770 10 44 0 75 0 26 SH 91 2.42 0.8 0.35 18 1.37 06 0.13 94 2 10 0.64 0 25 77 1.24 0.58 0 17 200 2 74 0 67 0.26 FM OH 29 3.46 0 83 0 44 3 0 44 0 41 0 05 20 0 92 0 55 0.24 23 1.50 0.57 0 25 74 3.48 0 08 0 31 9H 3 1 02 0 77 0 37 1 0 30 0 30 0 13 1 0 83 0 63 01 9 0 53 0.43 0.16 to 1 02 0 51 02 10H 2 0 76 0 83 0 22 0 1 0 45 0.45 0.18 1 0 67 0 67 0 43 4 0 76 06 0 26 0 0 0 1 0.52 0.52 1 0 S2 0 S2 11H l

l Table 11-2 l

Braidwood-1, SG-D Estimated Hot Standby SLB Burst Probability at EOC-4 i With Uncertainty Adjustment Free Span Burst Limited TSP Disp. Burst ")

Tubc/ Location TW Burst Applicable SLB localTSP TSP Volts Burst Prob. Displ. Probability Burst Prob. ")

Row Col

< 0.12 7.6E-23 - 0.0 5 10.44 2 4E-02 37 34 - 0.0

<0.10 9.2E-24 3 8 82 1.4E-02 23 12 0.235 2.5E-17 - 0.0 3 5.02 2.0E-03 12 9 -00 0.246 8.8E 17 3 4.28 1.1E-03 11 9 ~ 0.0 7.8E44 0.148 1.6E-21 7 3 3.95 19

<0.10 9.2E-24 - 0.0 3 3.83 7.0E44 33 20 ~ 0.0 5.6E 04 0.101 1.0E-23 29 3 3.62 35 0.197 3.6E 19 - 0.0 3 3.21 3.4E-04 11 12 -00 3.lE 04 <0.10 9.2E-24 42 3 3.12 16 ~ 0.0 2.lE 04 0.295 2.0E-14 105 3 2.84 2

0.077 8.5E-25 - 0.0 3 2.77 1.9E 04 8 20 ~ 0.0

<0.10 9.2E 24 3 2.60 1.4E 04 43 56 ~ 0.0

<0.12 7.6E-23 5 2.48 1.2E 04 32 27 ~ 0.0 8.6E-05 <0.10 9.2E-24 45 53 3 2.30

<0.10 9.2E-24 ~ 0.0 3 2.20 7.2E-05 31 20 ~ 0.0 6.9E-05 <0.10 9.2E-24 27 35 3 2.18

< 0.10 9.2E-24 ~ 0.0 3 2.16 6 6E-05 41 55 ~ 0.0 6.0E-05 < 0.10 9.2E-24 53 3 2.11 43 ~ 0.0 4.6E-05 0.126 1.5E-22 15 3 1.98 18 ~ 0.0 4.lE-05 <0.10 9.2E 24 29 3 1.93 18 - 0.0 4.0E-05 <0.12 7fE-23 21 110 5 1.92 0.507 8.8E 06 2.4E 07 113 3 0.59 2.4E-07 2 6.0E-09 2.2E-07 0.418 6.0E-09 110 3 0.58 3

0.250 1.4E 16 - 0.0 7 0.56 1.9E 07 2 113 ~ 0.0 3.6E 06 0.206 9.8E 19 108 7 1.10 2 - 0.0 3.4E 07 0.177 3.8E-20 105 7 0.64 2 ~ 0.0 5.6E 07 0.168 1.4E-20 II 7 0.72 2

0.172 2.2E-20 - 0.0 106 7 0.52 14E 07 4 - 0.0 6.0E-06 0.223 6.6E-18 4 112 7 1.24 0.214 2.4E-18 - 0.0 112 7 1.24 6.0E-06 5 ~ 0.0 2.4E 07 0.144 1.0E-21 7 105 7 0.59 0.128 1.8E-22 - 0.0 7 0 88 1.3E 06 13 110 2.4E-07 Total Burst F#=Niity 4.5E 02 Notes: h 1.

Analysis conservarrvely assumes the TSP 6? ~ ment exposes a throughwall crack equal to the

2. The applicable burst probability is the lesser of the free span and TW burst probabilities.

11/7/94

.n ., u ... n ,<= % ,,,,...t.

i

' Table 11-1: Probability of Burst of Tubes Having Different TSP Displacement Magnitudes I Byron Unit 1, Model D4 SG, TSP "C" SLB Load Cases Based on Throughwall Crack Length for AP of 2560 psi.

Plate SLB Displacement Range 0.25 0.3 0.35 0.4 0.45 0.5 0.55 0.6 Cumulative; ExfM Crack I2ngth ==> 0.15 0.2 1.lE-18 3.2E-16 8.2E-14 1.7E-I l 2.4E-09 2.0E-07 9.2E-06 2.2E-04 2.9E-03 Pr(Bursti Probability of Burst ==> 4.3E-21 48 26 0 0 0 0 0 4.43E-10 I C 168 114 76 No le

~

B at le Break at S/G No le 0' 200 126 60 8 0 0 0 0 0 6.79E-13 2.0 Uncertainty Factor J 0 0 0 7.0 M B at S/G zie 218 122 60 4 0 0 0 0 0 0 3.49E-13 1.75 Uncertainty Factor J Hot Standby SLB 108 86 64 46 28 8 0 0 0 1.65E-06 (MULT1 FLEX Loads) C 170 '

Break at S/G Nozzle Note: All remaining intersections of plate are assumed to undergo a displacement of 0.10".

iWK tim 94, t14FPM ICAESURST XLSlCAE De

Table 9-2 l

i Braidwood Unit 11994 TSP Inspection Summary 54 D " 'Pepenhen 54 A 54 5 ACC RPC BaM;ra APC Bseem RPC Betten ,

gegann myc assem RPC sees Md y h Ses[ spqN Sees % Card. esas apd esas %CmW. Sete epd $ses Md Said syd esas ed Said0spg 05 0 0 QA0 0 05 0 ISD 0 U 02 0 0 Sm 2 OM 0 Om 0.14 0 0.00 0 0 00 0 1 Om 0 0.00 0.1 1 0.78 1 14 2 3 9.e 0 0.00 21 0.M 1 4.79  !

8 cm 0 0.00 5 1.84 0 0A0 7 j

0.2 4 19 05 91 SR 13 28 8 31 4AD S 16.13 152 8.41 33 21.71 48 7As 11 22.45 21 8.80 0.3 e tin tu 17.20 41 mm a um 13 is m m2 18.30 s2 35.c 0.4 N irm 22 m.1 Si m21 as 34A2 402 32A2 102 40 2 8 30m 198 3231 GB 43Je 111 32 5 03 92 30 03 40 47 03 43 44.12 18 33.33 100 47.12 89 58.5 107 48R SS 52.34 407 47A2 217 33A2 08 82 43DS SS M.70 48- 41.75 SS #1 R S1 54 M 330 Bese 1N SSR .

as $2m e 70.50 27 71.5 8 20 m 142 00.53 88 00 98  !

O.7 4 24AD 117 71.38 80 73.90 es 71R di We 238 m24 19e s2m 0.8 45 St.07 28 57.78 25 80A0 20 73 2 304 78.71 100 7333 09 51 88.29 41 80.30 23 88.34 7 30.43 92 80JS 74 80 43 30 7 BAS i

1 33.33 N 88.31 57 80A8 34 81.84 29 88J9 145 Sta 125 08.21 44 72.52 38 88.38 3 90.44 1

4 82.85 5 83.33 41 90.18 31 75A1 41 87.75 38 87.00 122 88 49 97 79.51 ,

1.1 34 77.34 25 73.33 18 100A0 80 80.42 73 91 2  !

25 80 88 22 88.00 0 N.85 4 90A7 31 33.11 29 33.55 18 Som 1.2 2 08.50 0 0.00 18 N.81 17 N 44 14 M.38 13 RIAS SO 91 2 SS em 1.3 28 84.42 25 100.00 0 100.00 47 33JO 42 MJS 2 98.32 1 90.00 22 SSAS 18 81A2 9 31m 1.4 14 48 40 14 100A0 0 Sem 5 83.33 25 M.21 22 88A0 11 87.88 10 00A1 0 98 2 0 8 97A4 7 87A0 1.S 2 87.00 2 100 00 7 98.30 8 88.71 4 5.10 3 75A0 22 es se* 20 30.91 1A 9 GSJ4 0 100 m 8 100.00 22 .SSAS 22 We 0 57AS 0 3 98.98 3 100 00 8 NR 1.7 13 91.08 13 100m 2 esR is 98.41 15 33.75 1A 8 82.21 4 100.00 2 97 3 2 100.00 3 98A7 3 100m 3 98.4 i

0 3 98.15 3 100 00 4 SSR 4 100A0 14 SSA2 14 100A0 1.9 7 33JO 7 100.00 0 873 11 87.23 11 100A0 0 97.79 0 2 8034 2 100A0 3 97.41 3 100.00 2 8 tem 8 100.00 1 90.00 1 98.43 1 100 00 0 87.41 4 STA2 7 57.2 2.1 5 94M S 100.00 2 Se m 15 88.17 15 m.W 3 98.72 3 100.00 4 STAS 4 100.00 2.2 8 SS.8B 8 100.00 0 38.53 4 M AD 4 08.32 0 98.83 0 88.72 0 1 W.13 1 100.00 2.3 3 98.32 3 100 00 1 100.00 0 N 53 1 30.81 1 100 00 0 38.13 2 08 2 2 tem ,

2.4 1 98.45 1 NE 1 MAD 7 PSm 7 tem 4 97.03 4 100 00 1 88.00 1 100A0 1 00A1 1 100.00 2.S 1 100 m 3 98.75 3 100A0 2 100.00 0 90.00 0 SSA1 0 1 98.41 2.6 2 97.31 0 98 41 1 08.79 1 100.00 I 2.7 1 57.45 1 100.00 0 SSA0 0 98.91 0 '

1 80.28 1 100.00 1 100.00 1 90.00 1 08.95 1 100 m 3 98A0 3 mm 2A 0 57.45 0 1 98.70 1 100.00 1 SS.M i 100.00 r 2.9 0 97.48 0 0 SSJS 0 100A0 0 90JS 01002 0 M.70 3 mm 3 1 MAD l 3 3 97.88 3 100 00 1 100.00 4 W.19 4 15 2 ,

3.2 3 98.30 3 100.00 0 :.NJS 0 100m 1 Nm 0 90J8 0 100.00 1 SSAD 1 100.00 3 58.30 3 100m ,

3.4 2 SSAS 2 100.00 3 10040 0 100A0 0 98m 3 88.41 ,

30 3 SS 01 3 10040 0 88.28 1 100A0 3 W.52 3 100A0 1 88.15 1 100 00 1 90 83 1 100 00 0 100.00 1 W.14 38 4 90 87 4 100m 4 2 SS 43 2 100 00 0 90 83 0 100.00 2 W.42 2 1002 1 88.71 1 900A0 )

1 98.58 1 100.00 0 SSA3 0 100 00 0 W 42 42 2 88.75 2 100.00 i 4.4 0 98.58 0 1 100.00 1 100 00 0 100 00 1 WR 1 100 m 0 0 100.00 0 100.00 0 NA7 0 30.78 0 48 0 sem 0 30.78 0 48 0 98.98 0 0 100.00 0 100.00 0 MR 0 100.00 0 100A0 0 NR 1 SB22 1 15.00  ;

8 1 W.72 1 100.00 1 100A0 1 St.86 1 100.00 S.5 0 30.72 0 0 100 00 0 100A0 1 m.71 0 100 00 0 38.71 1 etm 1 1 5 .00 ,

0 1 Se SS 1 100.00 0 100.00  !

i 0 100A0 0 W.71 0 GBAB 0 l 8.S 0 SEAS 0 0 100A0 0 100A0 0 M.71 0 30.5 0  !

7 0 seas 0 0 100A0 '

0 100.00 0 190A0 0 m.71 0 sem 0 7.S 0 Seas 0 0100A0 0 100A0 0 W.71 0 SOm 0 S 0 SS OS 0 0 08.71 1 33A3 1 100.00  !

SA 1 100.00 1 100.00 0100A0 0 100A0 0 100m i mm 1 mm 1 mm i ,00 m 9 0 100m 0 0 100m 0 01002 0 100.00 0 Nm 0 W.SS 0 }

9.S 0 100D0 0 0 0 100A0 0 100.00 0 #AS 0 SOAS  !

to 0 100A0 1 tem 1 imm 1 tem 1 mm 10 3 0 100 00 0 0 100 00 0 100 00 78 togs est . Os 373 2731 705 470 272 I

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i Table 9-4 Byron-1: No. of EOC-6 ODSCC Indications vs. TSP Location

% ofTotal  :

SG-B SG-C SG-D Total '

TSP SG-A 605 303 1927 62.7 %

03H 416 603 253 113 810 26.3 %  :

05H 219 715 188- 6.1% l

% 47 17 28 07H 16 3 91 3.0%  ;

08H 42 30  :

6 7 45 1.5%

09H 23 9 l 0 3 9 0.3%

10H 6 0 2 3 5 0.2%

11H 0 0 914 899 460 3075 Total 802 ,

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l 10/24/94 3:01 PM CCE94RPT.XLS T9 4 GRDATA

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Table 9-5 Braidwood-1: Largest EOC-4 TSP Indication Voltages (Indications >4.0V) l Tube EOC-4 (1994) EOC-3 (1993) Cycle 4 f Support EOC-6 Analysis Re-analysis Voltage Row Column Plate No. Indication Voltage Voltage Growth S/G 34 5H DSI 10.44 0.68 9.76 D 37 3H DSI 8.82 0.76 8.06 D 23 12 5H DSI 8.33 1.66 6.67 A 45 41 5H DSI 5.54 0.50 5.04 A 18 23 9 3H DSI 5.02 0.76 4.26 D 12 ,

3H DSI 4.99 0.32 4.67 A 27 43 9 3H DSI 4.28 0.39 3.89 D 11 9 3H DSI 4.25 0.55 3.70 B 7 5H DSI 4.18 0.68 3.50 A 6 91 1

h 4

a 1

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l

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10/24/94 3:02 PM CCE94RPT.XLS T9_5 l

Figure 9-1 Braidwood Unit i Number of Indications as a Function of Elevation 800 700 600 E SIG A 3 500 .

ESIGB E

'N N SIG C O SIG D e

o

$ 300 -

E 200 100 - -

I I l- "I "

i 0- 1 11H SH 7H BH 9H 10H 1H 3H Elevation 10/25/94 9:45 AM C;WRDWDAPCCCESTA2.XLS F9_1

Table 9-6 Byron-1: Largest EOC-6 TSP Indication Voltages (Indications >4.0V)

EOC-6 (1994) EOC-5 (1993) Cycle 6 Tube 50C-6 Analysis Re-analysis Voltage Support Voltage Voltage Growth S/G Row Column Plate No. Indication 03H DSI 10.95 1.09 9.86 C 20 7 03H DSI 7.64 0.93 6.71 A 3 3 03H DSI 7.10 1.14 5.%

A 20 7 03H DSI 5.92 0.44 5.48 A 3 107 03H DSI 4.56 0.91 3.65 B 20 104 03H DSI 4.09 0.35 3.74 D 25 38 I

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l Figure 9-4. Byron-1 EOC-6 Bobbin Statistics (All SGs) 400 _ _  : :  ;;__;;;;. : : : : : : :

100 %

90 %

350 -

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300 - -

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/ k 8

M No. of Bobbin Indicctions j 250- -

60 %

-*-Cumulative Probability f

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f 50 -

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~~o2 es es os es er oa'es is u it is'is is is'ir'is is zo'2: 22 23 ' 2s as 2s 2r si n ur ssa sso sw ee sss ssa rio rss ioss Bobbin Signal Amplitude 10/24/94 3:16 PM T31 CHART.XLS BobbinCPDF

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

Mgure 9-5. Byron-1 EOC-6 RPC Inspection Statistics (All SGs)

: : : : : : : : : : : : : : : : = 100 %

200 90 %

180 80 %

160 140 p

)

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{a M + Percent RPC Confirmation .j j 50 % a o 100-- 0 5:

4 s.

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fu i 30% b 60 -

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40 10 %

20 --

II[II...........___,_____-_____

0 02_03IIII 04 OS OS OF OB 09 10 1.1 12 13 14 15 15 ti 14 19 20 2.1 22 23 24 25 26 27 31 32 3 31 364 3 00 398 40 0.,,

Bobbin Signal Amplitude 10'24/94 3:13 PM T31 CHART.XLS RPCConf

Figure 9-6. Byron-1: Number of ODSCC Indications vs. TSP 700 600 500 E \

E  !

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5SG-A l

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Figure 9-7 Lambons of N94 Bobbm indcatons at TSP 3 for SGs A and B  ;

I at Braidwood-1 I

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10,0 ANALYSIS METHODOLOGY FOR TUBE BURST PROBABILITY WITII LIMITED TSP DISPLACEMENTS Since the TSPs do not undergo any displacement relative to indications developed within the upper and lower planes of the TSPs during normal operation, tube bmst at pressures less than three times the normal operating differential pressure is obviated by the presence of the TSP.

Thus, the RG 1.121 requirement relative to 3 AP is inherently met. Hence, the considerations documented in this section are limited to determining the cnargin against tube burst during a postulated Steam Line Break (SLB).

10.1 General Description of Analysis Methods The essentials of the analyses consist of consideration of correlations of the burst pressure of throughwall cracks relative to crack length and the burst pressure of ODSCC TSP indications relative to the NDE ECT amplitude, i.e., the bobbin voltage. The concern for the potential of tube rupture during a SLB is based on the consideration that the pressure gradient in the SG will cause the TSPs to deform out of plane and expose TSP intersection tube ODSCC indications such that they behave as free-span indications without the constrain of the TSP.

The evaluation of the likelihood of tube rupture is based on the calculated deformations of the plates to determine the magnitude of potential exposure, the correlation of the burst pressure of tubes with free-span ODSCC indications to bobbin voltage, and the correlation of the burst pressure of tubes with free span axial cracks to crack length.

Most of the ODSCC TSP tube indications in a SG occur at row and column locations where the TSP undergoes relatively minor displacement, say s 0.1", during a SLB. Since the thickness of the TSP is 3/4", it is unrealistic to treat each indication as though it would be fully exposed during a SLB. In this case the expected burst pressure can be calculated by consider-ing a throughwall crack which is exposed by an amount equal to the deformation of the plate.

The probability of burst (PoB) for such a crack can be calculated based on the correlation of the burst pressure to the crack length. For larger deformations the PoB is calculated as the lesser value obtained by using both correlations. While the PoB may increase significantly for longer throughwall cracks, the actual PoB would be limited by that for a free-span ODSCC indication.

10.2 Burst Test Results for Cracks Extending Outside of the TSPs The results of burst testing of tubes with throughwall axial cracks has demonstrated a correla-tion between the burst pressure, or strength, and the crack length. For tubes in which a portion of the length of the crack was restrained in the radial and circumferential directions, i.e., as would exist within a hole in a TSP, the burst pressure correlates with the exposed crack length.

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This is because the local condition for burst is the achievement of a critical opening of the crack at the crack tip. For throughwall cracks in thin walled tubing the critical crack tip opening displacement (CTOD)is on the order of the thickness of the tube. He clearance between the OD of the tube and the ID of a TSP hole is not sufficient to permit the achieve-ment of the critical CTOD. Thus, the crack does not extend beyond that associated with minor blunting of the crack tip within the TSP.

10.3 Burst Probability as a Function of Throughwall Crack Length Recent analysis of burst test data for a variety of tube sizes indicates a strong correlation between the burst pressure, P., and the throughwall crack length, a, using an exponential relationship, i.e.,

P, =

[0.0615 + 0.534 e -a2n a], (10,1)

R where and R are the thickness and the mean radius of the tube, and or and og are the yield and ultimate tensile strength of the tube material. He term in brackets is usually referred to as the normalized or non-dimensionalized burst pressure, Py. The exponent term, A, is referred to n the normalized crack length, where a

A" (10.2) lR,t The coefficients of equation (10.1) were found by performing a non-linear regression of Py on 1, i.e, 2 ,1 Py= g, + g, e (10.3)

The index of determination of the regression was found to be 98.3% with a stipitrd error of 0.015. The p values for each of the coefficients was significantly less that 0.b The distribution of the residuals was found to be approximately normal. A plot of the resulting relation corresponding to the form of equation (10.1) is provided on Figure 10-1.

Before proceeding with a discussion of the evaluation of the probability of burst, it is noted that equation (10.1) yields estimates of the critical crack length for burst during SLB of 0.75" for the actual SLB differential pressure and 0.51" for a margin of 1.4 times the SLB differen-tial pressure for material considered to have lower tolerance limit (LTL) yield plus ultimate tensile strength at 650'F.

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t A generalized form of equation (10.1) can be written as, f  %

P, of , (10.4)

P, =

where o is the flow stress of the material (half of the sum of the yield and ultimate strengths).

f Equation (10.4) is used to estimate the probability of burst based on the variance of the estimate of Py about the regression equation and the variance of ofabout a large database of _'

measured properties of tubing installed in Westinghouse steam generators. An unbiased estimate of the variance, V, of P, is given by [

r i2 Py* V(o ) + o' V(P (10.5)

V(P,) = f y ) -V(of ) V(Py )

r s The standard deviation of the burst pressure, of,is taken as the square root of V(Py). If P, is  ;

an actual burst pressure, it is assumed that the statistic

( P" - P, ) -

t- , (10.6) o, ,

is distributed as a Student's t distribution with degrees of freedom equal to the degrees of ,

freedom used for the regression of Py on A.

1 Taking P, equal to the SLB pressure, a t variate is calculated from equation (10.6). The probability of randomly obtaining a t variate as large at that obtained is then calculated from  ;

the cumulative Student's i distribution. A plot of probability of burst as a function of crack length using this approach is provided as Figure 10-2. This is the probability of burst (PoB) during a postulated SLB for a tube with a throughwall crack lenga of a. In actuality, the  !

distribution of P, will not follow a Student's t distribution. Monte Carlo simul ations ofh te  ;

distribution of the burst pressure have resulted in# distributions which appear close to the form l of a Student's t distribution, but which have a longer tail in the higher burst pressure range, t.e.,  !

they are skewed right. Thus, the use of equation (10.6)is conservative. Comparison of individual 95% upper bound Monte Carlo results with predictions from equation (10.6) indicate l a small level of conservatism for high probabilities of burst, e.g'., PoB greater than 0.1, and an  !

4 order of magnitude difference for low probabilities of burst, e.g., on the order of 10 .

1 The above equations apply to calculating the PoB for a single throughwall crack or indication. ) '

For multiple indications the PoB of one or more of those indications is found as one minus the wee k3.tpH TSP 10 WP5 }Q . 3 l

- ..a- -- . - - - -.- ,, -

probability that none of them burst. The probability of no burst, or survival, for any single indication is one minus the PoB, thus the probability of burst of one or more of m indications is given as, PoB(m indications) = 1 -h (1 -PoB,) < f PoB,, (10.7)

& =) & =1 where PoB,is the probability of burst of the A* indication. In practice, the indications are segregated into crack length bins with all indications in one bin considered to have the length of the upper bound of the bin. Thus, the PoB for all indications in the same bin is the same.

By multiple application of equation (10.7), the PoB of one or more of all of the indications in the n bins is, PoB(n bins) < r, PoB , , (10.8) where r, is the number of indications in bin i and PoB, is now the probability of burst for an indication in the i* bin.

It is admitted that this practice omits consideration of the uncertainties of the parameters of the regression equation. However, formation of the normalized burst pressures from the test data includes an uncertainty of the material properties of the test specimens. Thus, the variance of an actual Py about its predicted Fy includes a contribution from the variation of material properties associated with repeated maasurements of tensile specimens from the same tube.

This, combined with the observation diat equation (10.6) always yields conservative results relative to Monte Carlo simulations is judged to outweigh the effect of omitting the uncertainty in the estimate of the coefficients.

10.4 Burst Probability as a Function of Bobbin Voltage The burst probability as a function of bobbin voltage is found in a manner similar to that for .

the burst probability as a function of crack length. Linear regression analysis of the burst pressure of model boiler specimens and sections of tubes re.noved from operating steam generators on the common logarithm of the bobbin amplitude ofindications in those specimens was performed. It was found that a prediction of the burst pressure, P,, as a function of the bobbin voltage, V, can be obtained from, P, = 7.858 - 3.137 log (V), (10.9) muowes 10 - 4 w = = w e.tve4

for material with a flow stress of 75 ksi. The number of data pairs, N, used in the regression was 82. The database consisted of all data as specified in WCAP-14046, with model boiler specimen 598-1 omitted as an outlier, plus the data obtained from testing of the Braidwood 1 removed tube sections, on itting one result for a specimen meeting the EPRI data exclusion criteria. The index of determination of the correlation was 81.2%. The p values for the intercept and slope coefficients were infinitesimal (< 10-). The standard error of the residuals, o r, was 0.94 ksi, and the residuals were found to be normally distributed about the regression line. A plot of the correlation relation and the database is provided on Figure 10-3.

For an individual indication with voltage V, the effective standard deviation of the burst pressure, o,, about the regression prediction of the burst pressure is 1

0, " Or I+y+ (log[y(log V -TogP)'

(10.10)

V,-IoiP)

The predicted burst pressure for a tube with an indication is obtained from equation (10.9) by adjusting for the actual flow stress of the tube material, i.e., by multiplying the result from equation (10.9) by the ratio of the actual flow stress, of, to the reference value used for the regression. Likewise, the standard deviation of the burst pressure about the prediction can be obtained by multiplying o, by the same ratio. To account for the variation in material properties in estimating the variance of the burst pressure for a random tube in the SG, an equation identical to equation (10.5) is used, except that the burst pressure predicted by equation (10.9) is substituted for Py and the ratio of the mean flow stress for the population of tubes to the reference flow stress is used instead of of. The probability of burst is then obtained by forming a Student's I distribution variate a la equation (10.6) and calculating the likelihood of randomly obtaining a i value as large or larger than the calculated variate. A plot of the probability of burst as a function of bobbin amplitude using the obtained correlation is provided on Figure 10-4.

The structural limits for tubes with a LTL material flow stress, based on a 95% lower bound prediction obtained from equation (10.9) using the effective standard deviation of equation (10.10), are 4.6V and 11.5V respectively for 1.4 times the SLB differential pressure, per RG 1.121 requirements, and the actual predicted SLB differential pressure. As noted previously,  !

the RG 1.121 requirement relative to three times the normal operating pressure differential is inherently met for ODSCC tube indications at TSP elevations. ,

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10.5 Modelling for Burst Probability with Limited TSP Displacements Two models were considered for the evaluation of the burst probability given the relative displacements of the TSPs during a postulated SLB. The first method is based only on the displacements. The second method is based on evaluating the largest indications found and comparing the PoB obtained from the free-span correlation of burst pressure to the bobbin amplitude and the PoB obtained from the displacement associated with those indications.

TSP Displacement Only Model An initial, and very conservative, estimate of the PoB can be made by assuming every intersection to have a throughwall crack equal to the thickness of the TSP. Thus, every intersection is considered to have a throughwall crack exposed by the magnitude of the displacement at each intersection. A summary of the PoB for the indications at selected TSPs is given in Table Il-1. The PoB for each plate was calculated using equation (10.8). The exposed crack length for each of the indications was assumed to be the upper bound for the bin. In addition, since displacements less than 0.1" were not binned,it was assumed that each of the remaining intersections of the plate, regardless of plugging, contained a crack which was exposed by 0.1" The maximum probability of burst,2.910, was obtained for plate "C" utilizing a factor of 2.0 to account for uncertainties in the SLB hydraulic / structural response analysis. The probability for a burst to occur at plate "A", the FDB, was not included in the evaluation since no indications were found at that plate. For plate "J" under the same assumptions the PoB was found to be 6.810'". Hence, the combined probability for plates "C" and "J" is then on the order of 2.910 Thus, under very conservative assumptions regarding the number of indications in the SG, the PoB is less than one-third of the limit specified in the NRC's draft Generic Letter on Alternate Plugging Criteria. Since the total number of indications is really a small fraction of the total number of intersections in the SG the magnitude of the conservat:sm would be expected to be multiple orders of magnitude.

A more realistic estimate of the PoB could be obtained by considering only the estimated number of indications and the spatial distribution of those indications in a steam generator.

Furthermore, assessment of the PoB of a tube afSLB conditions for limited TSP displacement only requires an estimate of the probability of a large indication occurring at the corners of the TSP where the TSP displacements are significant. Only plates 3 and 7 have significant TSP displacements such that a burst assessment is appropriate.

Although the flow distribution baffle (FDB, plate 1) has a few TSP intersections with signifi-cant displacements, no bobbin indications have been found at the FDB. The FDB in the Model D4 SGs has large tube to FDB gaps (nominally 100 mils diametral clearance toward the center of the plate and 88 mils with radialized holes for the outer region). Thus, there is a  ;

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Figure 10-1: Burst Pressure vs. Crack Length 0.750" x 0.043", Alloy 600 M A Steam Generator Tubes @ 650 F with or = 71.6 ksi ,,,

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RFK: 11/7/94,8 09 PM

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significantly lower likelihood of packed crevices with associated tube corrosion at the FDB intersections. Since no indications at the FDB have been found at either Braidwood 1 or Byron 1, the FDB is not included in the tube burst assessment.

Largest Indications Model Results of the PoB analyses for Braidwood I and Byron I for the largest indications in each plant and taking into account the row and column locations of the indications are summarized in Table 11-2 and Table 11-3 respectively. The details of the analyses are presented in Section 11 of this report. Results are presented therein which indicate that the probability of burst during a postulated SLB for a selection of the largest indications found at Braidwood I and Byron 1 is on the order of 2.410 4and 1.1 10-5 respectively. The analysis methodology consists of calculating a pubabliity of burst for each indication by first assuming the indication to be a throughwall crack with an exposed length equal to the TSP displacement at the location of the indication, and then by considering the actual indication to be completely exposed during a SLB. The lesser probability of burst obtained by both calculations is the expected, or applicable, probability of burst for that indication. The probability of burst of one or more of the group of indications is conservatively taken as the sum of the individual probabilities per equation (10.7).

i l

10.6 Conclusions A conservative estimate of the probability of burst of one or more indications in a SG during a ,

postulated SLB event indicates a likelihood of burst less than that required in the NRC's draft Generic Letter on Alternate Plugging Criteria for tube ODSCC indications at the elevations of the TSPs. Consideration of the aggregate of the largest individual indications in the Braid-l wood I and Byron 1 SGs results in an expected probability of burst of one or more tubes at either plant of approximately five and three orders of magnitude less than the requirement of the draft Generic Letter respectively.

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Figure 10-3: Burst Pressure vs Bobbin Amplitude 3/4" x 0.043" Alloy 600 MA SG Tubes (NRC Database @ 650 F) _

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RFK: 11/7/94. 8 24 PM I34PV_NRC.XLS) NRC P vs V

Figure 10-2: Probability of Burst @ 2560 psid vs. Crack Length 0.750" by 0.043" Alloy 600 MA SG Tubes @ 650 F a.e l

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RFK: 11/744,824 PM (34PV.NRC.XLS] PoS NRC

11.0 TUBE BURST PROBABILITIES FOR LIMITED TSP DISPLACEMENTS l 11.1 General Approach to the Tube Burst Probability Assessment l l

The tube burst probability assessment approach applied to support Braidwood I and Byron 1 is based on demonstrating limited TSP displacement in a SLB event to reduce the likelihood of a tube burst to negligible levels. The structural analysis results of Section 8 for TSP displacements in a SLB event are applied to a conservative assumption that the displacements expose a throughwall crack length equal to the TSP displacement. By applying the burst pressure versus throughwall crack length correlation of Section 10.2, both deterministic and probabilistic burst assessments are made for the assumed exposed throughwall crack length.

That is, the burst capability is a function of the exposed throughwall crack length. This analysis is equivalent to assuming that the indication at the TSP has a throughwall crack length approximately equal to the TSP thickness. This is an extremely conservative assumption since the bobbin voltages associated with such long throughwall cracks would be in the many tens of volts and much higher than that found at Braidwood I and Byron 1, which are bounded by a maximum indication of 10.9 volts. Consequently, the conservatism of this bounding burst assessment envelopes any realistic growth rate for Braidwood I or Byron 1 and the burtt margins obtained, based on limited SLB TSP displacements, are essentially independent of growth rates.

The conservatism of the bounding burst probability based on postulated throughwall cracks at all hot leg TSP intersections can be quantified by applying the limited displacement analyses to the actual Braidwood 1 EOC 4 and Byron 1 EOC 6 indication distributions. This assessment utilizes the actual tube locations for indications at TSPs and the associated TSP displacement at the indication location. Based on the free span assumption of the current APC methodology, the calculated burst probabilities for the actual EOC distributions exceeded 10 2 The benefits of limited TSP displacement in reducing the free span burst probability and in reducing the burst probability associated with the throughwall crack assumption are demonstrated by the EOC analyses for Braidwood 1 and Byron 1 given in this section.

The limited SLB TSP displacement would result in most of the crack length for indications at l TSPs covered by the TSPs and associated crevice deposits. This effect would tend to reduce leakage below that of free span indications which is the basis for data developed to support the SLB leak rate correlations used for leakage analyfes. EdF has performed system leak rate measurements on French S/Gs at pressure differentials exceeding SLB conditions. Bobbin voltage levels in the French units at the time of these tests exceeded that found at Braidwood I l and Byron 1, In addition, the French units included axial free span cracks in the roll transition at the top of the tubesheet which are left in service per repair criteria implemented by EdF, The total system leakage at pressure differentials typical of SLB conditions from these tests was on the order of a few gpm. This leak rate is much lower than would be predicted by the free span leak rate correlations considering only the indications at TSPs and ignoring the roll transition indications. Thus the EdF tests demonstrate the conservatism of the free span nr.n wn 11 . I we w

correlations particularly when the indications are within the packed crevice of the TSP.

However, the objectives of this report are to address tube burst probability and the benefits of limited TSP displacement on SLB leakage is not further addressed in this report.  :

i 11.2 Deterministic Burst Margin' Assessment ,

i As shown in Section 8, TSP displacements for a SLB at normal operaths conditions are small .;

for most plates and bounded by a maximum TSP displacement of 0.26 inch even when the full power SLB loads are adjusted by the 1.75 uncertainty factor. Only about 4 TSP intersections '

at the tubelane corners of the 7th TSP are subject to TSP displacements exceeding 0.25 inch.

As shown in Section 10, a throughwall crack length of about 0.51 inch (lower tolerance limit material properties - LTL) corresponds to a burst capability of 1.43 APst.s = 3660 psi. Thus, for a SLB at normal operating conditions, the maximum exposed potential throughwall crack ,

length of 0.26 inch is less than the R.G.1.121 structural limit of 0.51 inch. It is shown in l Section 8.6.4 that this corresponds to an extremely low probability of burst. It can also be noted in Section 10 that the free span throughwall crack length for burst at SLB conditions of 2560 psi is about 0.75 inch. Thus a free span throughwall crack with a length equal to the TSP thickness is required for burst at SLB conditions.

The exposed crack lengths associated with the maximum TSP displacements exceed the R.G.

1.121 structural limit of 0.51 inch only for a small number.of TSP intersections for the  ;

i bounding SLB at hot standby conditions. The reference hot standby SLB has a maximum TSP displacement of 0.35 inch and the bounding hot standby loads adjusted by the conservatively i developed 2.0 uncertainty factor has a maximum TSP displacement of 0.544 inch. Thus the i expected hot standby loads have displacements less than 0.51 inch and the bounding, adjusted hot standby loads result in only 12 tube locations at the tubelane corners of the 3rd TSP that-have SLB displacements exceeding 0.51 inch. Only these 12 tube locations, under the assumptions that a throughwall crack of > 0.51 inch occurs at these locations with the lower crack tip coincident with the bottom of the TSP, have the potential for not satisfying the 1.43 AP t., burst margin of R. G.1.121. No pulled tube to date, including 7 indications destructively examined with bobbin voltages between 8 and 23 volts have had a throughwall crack length as large as 0.51 inch. Thus the likelihood of even one indication with a throughwall crack length > 0.51 inch occurring,at the 12 TSP corner locations with displacements up to 0.544 inch is negligible and it can be expected that all actual indications would satisfy the burst margin requirements. The largest TSP displacement associated with an actual indication at Braidwood I or Byron 1 is 0.522 inch and this was a low,0.94 volt indication. Even if a 0.544 inch throughwall crack would occur at one of the 12 corner l d

locations, the tube burst probability for this indication would be a negligible 1.610 . I Overall, it is concluded that with the limited SLB TSP displacements for Braidwood I and Byron 1, the R. G.1.121 deterministic burst margin requirement for a burst capability of 1.43 AP,3., would be satisfied.

tse,nfrs 11 2 Ne.

• 11.3 Bounding Tube Burst Probability with Limited TSP Displacements A bounding estimate of the probability of burst of one or more tubes in a SG during a postulated SLB can be made by assuming every tube to have a throughwall crack equal to the thickness of the TSP at every TSP intersection. Thus, during a postulated SLB, every intersection is considered to have a throughwall crack exposed by the magnitude of the displacement at that intersection. The probability of burst of one or more tubes is then calculated as the sum of the probabilities of burst of each indication. The PoB for each plate was calculated using equation (10.8) of Section 10 of this report. The exposed crack length for each of the indications was assumed to be the upper bound of the displacement bin as specified in Section 8 of this report. Since displacements less than 0.1" were not binned, it was assumed that each of the remaining intersections of the plate, regardless of plugging, contained a crack which was exposed by 0.1" Finally, while likely not significant, it is noted that this approach conservatively treats the potential for the occurrence of multiple ruptures in the same tube as independent events, i.e., the rupture of multiple tubes.

A summary of the PoB for the indications at selected TSPs is given in Table 11-1. He maximum probability of burst,2.910, was obtained for plate "C" utilizing a factor of 2.0 to account for uncertainties in the SLB hydraulic analysis for a SLB at the exit of the SG nozzle initiating at hot standby conditions. For plate "J" under the same assumptions the PoB was found to be 6.810'" The probability for a burst to occur at-plate "A", the FDB, was not included in the evaluation since no indications were found at that plate. For plates with no exposed indications greater than 0.1" the probability of burst of an indication is negligible.

Hence, the total probability for the SG, i.e., the combined probability for plates "C" and "J", is on the order of 2.910~' Thus, under very conservative assumptions regarding the number of indications in the SG, the PoB is less than one-third of the limit specified in the NRC's draft Generic Letter on Alternate Plugging Criteria. Since the total number of indications is really a small fraction of the total number of intersections in the SG, the probability of burst is likely to be multiple orders of magnitude less than this result. Consideration of the row / column distribution of the indications would tend to reduce the result even further. His supposition is supported by the results presented in the following sections of this report. Therein the likelihood of burst for the population of actual Braidwood I and Byron 1 indications consisting i of the largest voltr.ge indications and the indications at the largest expected plate displacements  !

~

is evaluated. j l

11.4 Probability of Burst for Braidwood 1 EOC 4 TSP Indications The bounding tube burst probability given above is based on the assumption of throughwall crack indications at all hot leg TSP intersections with the throughwall length at least equal to the SLB TSP displacement. To obtain a more realistic estimate of the tube burst probability, a distribution ofindications at the TSP intersections is required. This can be obtained by rir.n wrs 11 3 mm l

applying the actual distributions found at EOC 4 for Braidwood 1. Based on the APC analyses of tube burst probabilities for this EOC distribution, the resulting free span tube burst probability was greater than 10 for S/G D, the limiting S/G at Braidwood 1.

The burst probability for S/G D at EOC 4 with limited TSP displacement can be obtained directly from the indications found and the TSP displacement at each specific indication.

Table 11-2 identifies the indications found in the inspection for the larger voltage indications and for indications at locations having the largest TSP displacements in the hot standby SLB l event. The bobbin voltage and free span burst probability at the given voltage level are provided for each indication. Also given in the table are the local TSP displacement (bounding l hot standby SLB displacements including the 2.0 uncertainty factor on the TSP loads) and the burst probability for a throughwall crack length equal to the TSP displacement (conservatively assumed exposed throughwall crack length). The applicable SLB burst probability column shows the lower of the free span or throughwall burst probability for each indication. The lower of the two burst probabilities is the appropriate value since the limited TSP displacement can reduce the free span burst probability but the free span probability cannot be exceeded.

The throughwall burst probability can exceed the free span value only because it is conservatively calculated for a throughwall crack while the free span value, based on bobbin voltage, is more realistically based on the actual crack morphology as reflected in the voltage amplitude.

For the S/G D indications given in Table Il-2, the total burst probability calculated assuming 2

free span (very large SLB TSP displacements) conditions is 4.510 which is dominated by the two largest voltage indications of 10.4 and 8.82 volts. Accounting for the limited SLB TSP displacements at the locations of the indications, the total burst probability is 2.410-' It can be noted that none of the high voltage indications occurred at locations of high TSP displacement and the TSP constraint reduces the burst probability for these high voltage indications to approximately zero. Only the small voltage indications found at the corners of plate 7, where SLB displacements are significant, contribute to the burst probability. The 0.59 volt indication at R2Cll3, TSP 3 had the largest local TSP displacement (0.507 inch) of any of the Braidwood 1 EOC 4 indications. For this indication and tube location, the free span burst probability is limiting and the associated 2.410 burst probability is the only significant contribution of all indications in the S/G.

l The results of Table 11-2 show the effectivenes of limited TSP displacements in reducing the tube burst probability to small values and also show that Braidwood I had an acceptably low burst probability at EOC 4. The application of the actual indication distributions reduces the

burst probability by four orders of magnitude (from 2.910' to 2.410) compared to the bounding assumption of throughwall indications at all hot leg TSP intersections. The benefits ,

of limited TSP displacement are clearly seen by the reduction of about five orders of magnitude (4.510-' to 2.410-') in the tube burst probabilities compared to the free span APC j analyses.

I i

TSP,Il WP5.4 November 9.1994

11.5 Probability of Burst for Byron 1 EOC 6 TSP Indications The analysis given above for Braidwood I can also be performed for the actual EOC 6 TSP indications at Byron 1. S/G C was the most limiting S/G for Byron 1 EOC 6. Table 11-3 providQ the same information for Byron 1, S/G C as given in Table 11-2 for Braidwood 1. Again, none of the large voltage ( > 2.0 volts in this case) indications occurred at tube locations with significant SLB TSP displacements. The indications at R5Cl (1.19 volt), RIC3 (1.74 volt), R3Cl (0.94 volt) and R3Cl14 (0.51 volt) occur at tube locations with significant TSP displacements of 0.492 to 0.522 inch and dominate the contribution to the limited TSP displacement burst probability of 1.110 For the indientions given in Table 11-3, the free span burst probability is 3.010 2 The results of Table 11-3, like the above results for Braidwood 1, show the effectiveness of limited TSP displacements in reducing the tube burst probability to small values and also show that Byron I had an acceptably low burst probability at EOC 6. The application of the actual indication distributions reduces the burst probability by more than two orders of magnitude (from 2.910 to 1.110 5) compared to the bounding assumption of throughwall indications at all hot leg TSP intersections. The benefits oflimited TSP displacement are also seen by the reduction of about three orders of magnitude ( 3.010 2 to 1.110) in the tube burst probabilities compared to the free span APC analyses. I1.6 Conclusions The results of the Braidwood I and Byron 1 tube burst assessment with limited SLB TSP displacements can be summarized as follows: The R.G 1.121 deterministic burst margin requirement for a burst capability of 1.43 AP3t3 (corresponding to a 0.51 inch throughwall crack) can be expected to be satisfied by the limited SLB TSP displacements for Braidwood I and Byron 1. TSP displacements exceeding 0.50 inch occur only at 12 TSP 3 tube locations for a hot standby SLB with the reference loads adjusted by a conservatively developed 2.0 uncertainty factor. The likelihood of a

  > 0.51 inch throughwall crack occurring at one of these 12 TSP 3 locations is very small.

The largest TSP displacement for any tube location with an indication at Braidwood 1 or Byron I was 0.522 inch and this was a low voltage (0.94 volt) indication. Even if it is postulated that a throughwall crack ~ was present at every hot leg TSP intersection, the tube burst probability for the expected or reference hot standby SLB loads would be 4.410 end for the reference full power SLB loads would be < 10 With this bounding throughwall crack assumption and with the bounding hot standby SLB loads (uncertainty factor of 2.0 applied to the reference loads), the tube burst probability would remain small at 2.910 Even when both of these bounding conditions are applied to the full power SLB, the tube burst probability remains < 10 787.i: 5P5 11 - 5 he. * *

 . If the bounding hot standby burst probability of 2.910'is weighted by the 4% of the typical fuel cycle with operation at hot standby conditions (3.8% evaluated for Braidwood-1, average of 1.7% for the last three cycles at Byron-1), the burst probability is only 1.210" per full      ,

operating cycle.

 . When the bounding hot standby SLB TSP displacements are applied to the actual Braidwood 1 EOC 4 and Byron 1 EOC 6 indications, the associated tube burst probabilities are very small at 2.4104and 1.1 10 5, ,,3p.ctively. When calculated as free span probabilities based on the licensed APC methodology for these EOC distributions, the tube burst probabilities were > 104. Thus, the limited TSP displacements reduce the tube burst probability by three orders of magnitude compared to free span conditions. In addition, the application of the bounding hot standby TSP displacements to actual EOC indications reduce the tube burst probability by two to four orders of magnitude compared to the assumption of throughwall indications at all hot leg TSP intersections.

( . The TSP loads adjusted by the 2.0 and 1.75 uncertainty adjustment factors for hot standby ! and full power operation represent loads that bound the analysis uncertainties in TRANFLO and the differences between codes such as the MULTIFLEX Code. These loads would be the appropriately conservative loads to apply for assessments of tube expansion at TSP intersections as a means of further reducing the SLB TSP displacements to negligible levels.

 . The results of this report demonstrate that, even under worst case uncertainties, tube burst probabilities are acceptable based on the limited TSP displacements and that tube burst should not be limiting for the Braidwood I and Byron 1 S/Gs. Based on these low tube burst probabilities for the most conservative possible tube degradation assumption, it is concluded that deterministic structural limits and low 1.0 tube repair limits to preclude tube burst are not required for the Braidwood I and Byron 1 APC applications. Even with the large voltage growth rates found for a few indications at the last Braidwood I and Byron 1 inspections, the S/Gs can operate to the planned, full cycle refueling outage and achieve low EOC tube burst probabilities due to the limited SLB TSP displacements for these S/Gs.

a l l T*P 1Iwrs 11 6  %=he 9. **H

Q\ Tcble 1103 Byron-1. SG-C Estimated Hot Standby SLB Burst Probability at EOC-6 With Uncertainty Adjustment Tube / location Free Span Burst Limited TSP Disp. Burst m LocalTSP TW Burst Applicable SLB Row Col TSP Volts Burst Prob. Displ. Probability Burst Prob.

• 7 3 10.95 2.7E 02 0.128 1.8E-22 ~ 0.0 20 13 5 3 80 6.7E 04 < 0.10 9.2E-24 ~ 0.0 25 32 30 3 1.64 5.7E-04 0.100 9.2E-24 ~00 27 98 3 2.68 1.6E 04 0 090 3.2E-24 -0.0 47 35 3 2.68 1.6E-04 < 0.10 9.2E-24 ~ 0.0 20 102 3 2.58 1.4E-04 0.118 6.2E-23 -00 21 98 3 2 50 1.2E 04 0.111 2.9E-23 - 0.0 27 59 3 2 45 1.lE-04 <0.10 9.2E-24 ~ 0.0 16 59 3 2.38 1.0E 04 < 0.10 9.2E-24 ~ 0.0 27 93 3 2.29 8.5E-05 0.116 5.0E-23 ~0.0 35 26 3 2.14 6.4E-05 0.105 1.6E-23 ~ 0.0 12 3 3 2.10 5 9E-05 0.344 4.0E-12 - 0.0 46 52 5 2.02 5.0E 05 < 0.10 9.2E 24 - 0.0 9 3 3 4.7E-05 0.394 6.2E-10 6.2E 10 1.9_9 46 70 5 1.99 4.7E-05 <0.10 9.2E-24 ~ 0.0

[ 18 42 3 1.97 4.5E-05 < 0.10 9.2E-24 ~ 0.0 25 10 3 1.95 4.3E-05 < 0.10 9.2E-24 ~ 0.0 40 78 5 1 E8 3.7E-05 <0.10 ~ 9.2E-24 ~ 0.0 40 65 3 1 85 3.4E-05 < 0.10 9.2E-24 - 0.0 12 74 5 1.81 3. lE 05 < 0.10 9.2E-24 ~0.0 5 1 3 1.19 5 0E-06 0.500 5.3E-06 5.0E4 1 3 3 1.74 2.6E-05 0 492 2.9E-06 2.9E-06 3 1 3 0.94 1.8E 06 0.522 2.5E-05 1.8E4 2 2 3 0.70 5.0E 07 0.507 8.8E 06 5.0E 07 7 2 3 1M 2.4E 06 0.446 7.2E-08 7.2E 08 l 8 2 3 0.65 3.6E 07 0.431 1.9E 08 1.9E-08 ) 1 6 3 0.72 5.6E 07 0 412 3.4E-09 3.4E 09 2 6 3 0.68 4.4E 07 0.402 1.3E-09 1.3E-09 l 8 4 3 0.71 5.3E-07 0.385 2.6E 10 2.6E 10 l 5 5 3 0.72 5.6E-07 0.397 8.3E 10 8.3E 10 6 4 3 1.00 2.4E-06 1 0.411 3.1E-09 3.1E-09 4 112 3 0.56 1.9E 07 0 458 2.0E-07 1.9E-07 3 114 3 0.51 1.3E47 0.522 2.5E 05 1.3E-07 4 111 3 0.44 6.7E 08 0 433 2.3E-08 2.3E-08 4 110 3 1. I 1 3.7E4 0.408 2.4E-09 2.4E-09 2 109 3 0.74 6.3E-07 0.402 1.3E 09 1.3E 09 8 111 3 1. I 1 3.7E-06 0.385 2.6E 10 2.6E-10 Total Burst Probabihty 3.0E 02 1.lE-05 Notes: I

1. Analysts conservatsvely assumes the TSP dispin=nent exposes a throughwall crack equal to the Ap1=~

2. The applicable burst probability is the lesser of the free span and 1W burst probabilities.

C 3iarcisce94\r f t\burspret.als 11/1/94

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