ML20211F244

Whipjet Program Rept

ML20211F244
Person / Time
Site:	Beaver Valley
Issue date:	10/17/1986
From:	ELECTRIC POWER RESEARCH INSTITUTE, ROBERT L. CLOUD ASSOCIATES, INC., STONE & WEBSTER ENGINEERING CORP.
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ML19292G154	List: ML20211E960 ML20211F244
References
NUDOCS 8610310110
Download: ML20211F244 (182)

Text

. _ __.

l BEAVER VALLEY POWER STATION UNIT NUMBER 2 i

NHIPJET PROGRAM REPORT i

OCTOBER 17, 1986 I

l Prepared for Duquesne Light Company by:

Stone and Webster Engineering Corporation Robert L. Cloud Associates, Incorporated Electric Power Research Institute -

j B610310110 861023 PDR ADOCK 05000412 l A PDR i

TABLE OF CONTENTS SECTION PAGE

1. Introduction 1-1

2. Summary of Results 2-1

3. Program Scope 3.1 High energy systems 3-1 within NHIPJET 3.2 System descriptions 3-4

4. Screening Considerations 4.1 Anomalous conditions 4-1 4.2 Fluid transient 4-3 4.3 Stress corrosion cracking 4-6

5. Crack Propagation Analysis 5.1 Stress determination 5-1 5.2 Crack growth law 5-6 5.3 Stress intensity factor 5-7 calculation 5.4 Initial crack size 5-8 5.5 Acceptance criteria for 5-8 fatigue crack growth 5.6 Example 5-9 5.7 Fatigue crack growth results 5-12 6 Material Property Data 6.1 Austenitic stainless steel 6-1 lines 6.2 Ferritic steel lines 6-1 6.3 Highest stresses and lowest 6-2 material properties 7 Leak Detection 7.1 Introduction 7-1 7.2 Leak detection systems 7-2 (inside containment) 7.3 Leakage diagnosis 7-3 7.4 Actions 7-3 7.5 Leak detection outside 7-5 containment

8. Leak Rate Calculation Summary 8-1

9. Crack Stability Calculations 9-1 (For Normal + SSE Loads) 10 Crack Stability,Under 10-1 Excessively High Loads i

i Table of Contents (continued)

APPENDICES A. NUREG-0582 Fluid Transients B. Stainless Steel Material Properties C. Ferritic Steel Material Properties D. Nelding Procedure E. Leak Rate Calculation Methodology F. BVPS-2 Leak Rate Curves G. FLET Verification Calculations 2

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I SECTION 1 INTRODUCTION WHIPJET is an engineering program which assures that essential structures. systems, and equipment are given protection at least equivalent to that conventionally provided by designing hardware to protect against the dynamic effects of postulated pipe whip and jet impingement. WHIPJET demonstrates that the fluid leakage from a postulated defect at a postulated break location in a high energy piping line can be detected well before the rupture of the pipe. Adequate time is assured to both detect and repair the leak and to bring the plant to a safe, controlled shutdown. The program has been applied to the Beaver Valley Power Station. Unit 2 (BVPS-2) for balance-of-plant (BOP) piping (i.e. other than the Primary Reactor Coolant Loop). Since HHIPJET is based on the leak-before-break (LBB) approach, using elastic-plastic fracture mechanics for assessing the potential for pipe rupture. WHIPJET is consistent with the procedural recommendations and analytical criteria found in NUREG-1061 Volume 3. [1.1] The basis for assurance of piping design and construction quality is presented in the BVPS-2 Final Safety Analysis Report (FSAR).

The HHIPJET program is limited in scope to only the pipe rupture locations which have been previously determined to be a threat to essential surrounding safe shutdown equipment through potential damage caused by either pipe whip impacts or det impingement loadings. Pipe break locations were determined following NUREG-0800 [1.2) Standard Review Plan (SRP) Section 3.6.2 except that arbitrary intermediate breaks were previously eliminated as approved by the Nuclear Regulatory Commission (NRC). Some pipe break locations do not require any protection hardware on the basis of detailed target and hazard analyses. Remaining system locations were then screened to eliminate piping for which crack growth mechanisms (such as corrosion or fatigue) and other failure potential (eg., water hammer) could not be ruled out or there was no economic benefit to include this piping in the program. Each remaining line was also analyzed for fatique crack growth possibilities for a postulated part-through wall crack of an initial size equal to the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code.Section XI acceptance standards: this defect was shown not to grow to a size which could lead to a through-wall crack during the normal operating life of the piping system.

l Lines were checked to determine which break locations were most critical in terms of highest loads and minimum material properties: these break locations were then used in the elastic-plastic fracture mechanics calculations for assuring LBB, The i pertinent system locations were analyzed to determine the stability of postulated through-wall flaws in the base metal or weld given the local load and stress conditions in the pipe

} [(dead weight plus thermal plus pressure plus Safe Shutdown l

l l 1-1

Earthquake (DH + TH + P + SSE)). Leak rate versus crack size was determined for break locations considering normal operating loads (DH + TH + F ) and fluid conditions. A limiting leak rate based upon the detectable leakage limit multiplied by a margin was used to determine a leak rate crack length. This leakage size crack length was analyzed for stability using a further margin on crack size and piping loads equal to normal + SSE. Crack stability was further assured by evaluating the limiting leakage crack size for excessively high piping loads equal to 1.414 (normal + SSE).

If these engineering criteria were satisfied. protection hardware will not be required and need not be installed. No relaxation of environmental qualification requirements was intended with this approach. In fact. the overall integrity of the BVPS-2 plant was improved since much more is now known about the piping than was the case before the application of HHIPJET.

In summary. NHIPJET addresses an alternative engineering approach for the provision of protection from the mechanistic effects of postulated pipe rupture. Figure 1.1 shows the steps required in the HHIPJET program to satisfy the leak-before-break approach for eliminating pipe rupture hardware.

REFERENCES 1.1 Report of the U u S m Nuclear Regulatory Commission Piping Review Committee: Evaluation of Potential Pipe Breaks. NUREG-1061. Vol. 3. November 1984, 1.2 Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants. LHR Edition. NUREG-0800 July 1981 1-2

Postulated SRP 3.6.2 Pipe (Except arbitrary Breaks ,

intermediate breaks)

Target and Interaction Evaluations Screening:

Industry Experience, Economics Leak Detectability Margin Material Properties Fatigue Crack Growth Evaluation i

Limiting Detectable Leak Rate Leak Rate x Margin Calculations Normal Loads l

Margin on Crack Stability Normal +SSE Crack Size Evaluation Loads Stability Check For Excessively 1.414(Nor+SSE)

High Loads Loads FIGURE 1.1 WHIPJET PROGRAM 1-3

i SECTION 1

SUMMARY

OF RESULTS Fracture mechanics technology has advanced to the point that an engineering approach using the concept of leak-before-break in lieu of postulating double-ended pipe rupture is now possible, i

The results for such a program (termed NHIPJET) has been successfully applied at BVPS-2. The overall results indicate i

' that pipe rupture hardware is not necessary for lines inside containment greater than or equal to 6-in. nominal pipe size in

, systems that have passed initial screening. Outside containment, the only lines analyzed were ferritic carbon steel, 3-in. in diameter, and these lines were also judged to be exempt from hardware requirements.

' The screening process mentioned above included evaluation for the potential for stress corrosion cracking, excessive fatique, water 1 hammer, pipe.

or other conditions which could result in failure of a i Also because of economic considerations, some lines were

! not evaluated using the NHIPJET methodology. The main steam system lines fell into this category due to the advanced state of construction and installation of the pipe rupture hardware. Some other lines in the steam generator blowdown system and reactor coolant system were deleted from NHIPJET analysis due to accessibility or projected low fluid leakage. Fatigue crack i

growth was also evaluated, and none of the lines failed the crack growth acceptance criteria (i.e., the extent of crack growth was minimal over the 40-year life of the plant). "

)

The evaluation for leak-before-break involved three key i processes: leak rate calculations for normal operating loads, crack stability analysis for normal + SSE inertia loads, and an excessive load case where stability is assessed for. loads much higher than normal + SSE. Three margins were assessed related to l these analyses: (1) a cargin on leak detection of 10 for inside containment and 4 outside containment on the minimum

! detectability limits of 0.5 spm inside containment and 0.05 spm outside containment (where visual inspection is possible); these result in 5 gpm and 0.2 gpm values for assessing stability crack sizes; (2) a margin on crack size for asrcesing stability of at least 1.8; and (3) a margin on loads of the square root of two (1.414) for the excessively high load stability check. The intent of all of these margins as required by the NRC staff was met for the lines analyzed for BVPS-2. ,

I The application of the NHIPJET leak-before-break program at BVPS-2 has led to the improvement in overall safety and integrity of the plant. Not only are unwanted and unnecessary pipe rupture hardware devices eliminated, but much more is now known about the piping and its capabilities than was the case before NHIPJET was applied.

2-1

SECTION 3 PROGRAM SCOPE 3.1 HIGH ENERGY SYSTEMS HITHIN HHIPJET The BVPS-2 pipe rupture evaluation program identifies the high energy. pipe break locations which are potentially hazardous to adjacent plant equipment which are required to mitigate the effects of the postulated event. The high energy systems and the associated abbreviations are:

ASS Auxiliary Steam System BDG Steam Generator Blowdown System BRS Boron Recovery System CHS Chemical and Volume Control System DGS Gaseous Drains System FHE Auxiliary Feedwater System FHS Main Feedwater System GNS Gaseous Nitrogen System MSS Main Steam System RCS Reactor Coolant System RHS Residual Heat Removal System SIS Safety Injection System These twelve systems were reviewed to establish operating pipe stress levels and then were subjected to break postulation criteria (in accordance with NUREG-0800 [3.1]). Further, the resulting pipe whip and jet impingement trajectories were delineated in order to identify any interactions that could occur with adjacent safety-related structures, systems, and components.

If the interactions were judged severe enough to cause significant impairment of the target's safety function, then the resulting damage was evaluated in order to determine whether loss of essential safe plant shutdown function would result.

This review indicated that there are no essential targets for postulated pipe breaks in several systems, and consequently no pipe rupture protection hardware needs to be provided. These systems were therefore eliminated from further consideration by the NHIPJET program. Note that several of the detailed targeting analyses were performed concurrently with the HHIPJET program and in some cases resulted in the corresponding breaks being i eliminated based upon targeting.

The remaining systems which have breaks that target essential safety equipment are:

l 3-1 l

4 BDG Steam Generator Blowdown System FMS Main Feedwater System MSS Main Steam System I RCS Reactor Coolant System RHS Residual Heat Removal System

, SIS Safety Injection System Mithin some of these systems, some breaks have been evaluated as not requiring hardware based upon detailed target analyses; these systems and affected pipe sizes include:

i RCS -- 6-in. inside containment RHS -- 10-in. inside containment (only one of two lines)

SIS - and 3-in. inside containment 3- and 4-in, outside containment BDG -- 1.5- and 2-in. outside containment 1

The above relevant systems and lines were reviewed against nuclear industry experience with pipe cracking as documented in NUREG-0691 [3.2] where several occurrences of thermal fatigue

induced cracking in main feedwater piping were noted. The feedwater system also has potential for water hammer. Based on this information and discussions with the NRC staff and its consultants, it was decided to exclude FHS from further review I within the NHIPJET program.

The NUREG-0691 review also noted that a number of cases of vibration-induced fatigue cracking were experienced at socket weld fittings for small lines in particular. Except f or the 2-

in. RCS lines, there are no postulated break locations in smcIl l lines which target essential shutdown equipment. Accordingly, all 2-in. lines were eliminated from further consideration.

t l The material type used for the remaining systems were reviewed i for documented toughness properties. Carbon steel was used for t

the MSS piping and for some lines in the BDG system. The MSS pipe whip restraints have been fabricated and installation is l near completion. Therefore, there was no economic benefit to l include this system in the MHIPJET review. Consequently, a I

carbon steel testing program, planned originally to establish toughness properties of actual BVPS-2 materials, was determined not to be cost-effective.

Instead, the evaluation of the remaining carbon steel piping was performed using material toughness test results provided by others. The balance of the systems use austenitic stainless steel which was judged to be sufficiently tough for inclusion in the NHIPJET program. Again, industry data for the stainless steel materials were utilized in the NHIPJET analysis. Steam Generator Blowdown piping located in the pipe tunnel area was considered too remote and inaccessible to be visually inspected for leaks. Accordingly . protective hardware is provided for the breaks in this area.

3-2

I Additionally, initial calculation of leakage rates using the NUREG-1061 Vol. 3 [3.3) safety margins of 10 on leakage and 2 on crack size suggested that some small lines would not be applicable for LBB. Therefore, these lines were not included in the NHIPJET program for eliminating pipe rupture hardware. After final target analyses, only one 4-in RCS line was ultimately eliminated from NHIPJET by this low leakage determination.

The piping deleted after screening and consideration of other factors (i.e., economics and leakage' are listed in Table 3.1.

The piping systems remaining and the associated hardware are shown in Table 3.2. An indirect benefit due to increased awareness of the piping systems and loads also resulted -- many other pipe break locations and associated hardware have been determined to be unnecessary due to detailed targeting and interaction studies. Many of these studies would not have been performed at this level of detail if the HHIPJET program had not prompted them.

Also indicated in Table 3.2 are the piping materials and sizes.

As shown, the pipe size ranges from 6 to 14-in, nominal pipe size for inside containment austenitic stainless steel (Types 304 and 316) and 3-in. for ferritic (carbon) steel in the BDG line outside containment (SA 106, Grade B). All field welds were made using the shielded metal arc welding (SMAH) process. A few shop welds utilized submerged arc welding (SAH). Both the stainless and carbon steel welds in these systems are used in the as-welded condition. Post-weld heat treatment was not performed as it was not required in accordance with the rules of ASME Section III.

The Beaver Valley Project uses the following nomenclature for piping:

Line Number 2RCS-008-020-1 5 6 d h $

Pipe Class

} Designates BVPS-2 piping-

- Line Number Piping System Abbreviation -- Nominal Piping Size (e.g., 8-in, diameter)

Table 3.3 presents the HHIPJET Program piping by line number and t

lists the piping dimensions, material type, operating conditions, insulation type and thickness, and indicates the plant area code were the piping is located. CS202, CS204, and CS206 are the j three reactor coolant loop cubicles. CS203 is the pressurizer cubicle which contains the surge line. VC104 is the area which contains the steam generator blowdown piping outside containment.

3-3 l -- -- --- ~ - - ~ ~ ' ' - ~ ~ '~

Table 3.4 lists the break locations for each of the lines listed

)_ in Table 3.3 where.

l 2RCS-004-C-C il 4 si l Break Number 1

Location Code Break Geometry C = Coolant Loop Cubicle C = Circumferential O = Outside Containment L = Longitudinal Split i

Figures 3.1 through 3.7 show the location of these breaks.

i 3.2 SYSTEM DESCRIPTIONS The following is a description of each MHIPJET system as to its primary function and its location within safety related areas.

3.2.1 Steam Generator Blowdown System (BDG)

The BDG system is designed to remove water f rom the secondary side of the steam generator and send it to .the blowdown flash l tank in order to maintain steam generator water chemistry and to limit the buildup of corrosion products. The lines are maintained at high pressure by the steam generator both inside and outside containment.

i 3.2.2 Reactor Coolant System (RCS) '

l The RCS transports heated water from the reactor core to the steam generators where heat is transferred to the main steam

system. Leak-before-break for the main reactor coolant loop has i been demonstrated by Hestinghouse; primary loop breaks are excluded because of this work and the new limited scope rule change to GDC-4. The RCS piping remaining in the MHIPJET program 1

consists of the pressurizer surge line and the reactor coolant i loop bypass lines. The piping is directly connected to the primary reactor coolant loops and is part of the reactor coolant system pressure boundary.

3.2.3 Residual Heat Removal System (RHS)

! The RHS transfers heat f rom the reactor coolant system to the primary plant component cooling water system in order to reduce i the fluid temperature of the reactor coolant system to the cold shutdown temperature. The high energy portion of the piping is i located directly adjacent to the connection to the reactor j coolant loop and is part of the reactor coolant system pressure boundary.

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3.2.4 Srfaty Inicction System (3IS)

The SIS is part of the emergency core cooling system (ECCS). The short term function of the SIS is the prompt deliver >* of borated water to the reactor core following a loss-of-coolant accident.

The system is divided into two parts: the low pressure safety injection system piping between the accumulator tanks and the RCS cold leg. and the high pressure injection system piping between the charging pumps and the primary loop. The high energy portion of the piping is located directly adjacent to the reactor coolant loopand is part of the reactor coolant system pressure boundary.

REFERENCES 3.1 Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants, LWR Edition, NUREG-0800, July 1981.

3.2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors. NUREG-0691, September 1980.

3.3 Report of the U. S. Nuclear Regulatory Commission Piping Review Committee; Evaluation of Potential for Pipe Breaks, NUREG-1061. Vol. 3, November 1984.

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3-5 l _ _ _ _ _ _ _. - - - - . _ -.

TABLE 3.1 PIPING ELIMINATED FROM HHIPJET SCOPE BY SCREENING OR OTHER CONSIDERATIONS SYSTEM RUPTURE PIPE RETAINED REASON FOR LOCATION SIZE HARDWARE ELIMINATION (IN)

FHS CONTAINMENT 16 5 INDUSTRY EXPERIENCE FHS OUTSIDE 16 18 INDUSTRY EXPERIENCE CONTAINMENT RCS CONTAINMENT 1.5, 2 2 INDUSTRY EXPERIENCE MSS CONTAINMENT 32 9 ECONOMICS MSS OUTSIDE 32 3 ECONOMICS CONTAINMENT BDG TUNNEL 3 5 INACCESSIBILITY RCS CONTAINMENT 4 i LEAKAGE MARGIN ' 10 TOTAL 43 3-6

TABLE 3.2 FINAL PIPING SYSTEMS FOR MHIPJET ANALYSIS PIPING PIPE MATERIAL BREAKS HARDWARE SYSTEM SIZE (IN) PRR JIS (1) (2)

BDG 3 SA 106, 4 3 0 GRADE B RCS 8 SA 376, 6 6 0 TYPE 304 RCS 14 SA 376, 13 8 0 TYPE 304 RHS 10 SA 376, 1 0 (3) 0 TYPE 316 RHS 12 SA 376 1 1 0 TYPE 316 SIS 6 SA 376, 20 8 4 TYPE 316 SIS 12 SA 376, 28 10 4 TYPE 316 TOTAL 73 36 8 TOTAL 44 NOTES:

(1) PRR = Pipe rupture restraint (2) JIS = Jet impingement shield

-(3) Break requires SIS restraint 3-7

TABLE 3.3 WRIPJET PROGRAM PIPH7; DATA LINE NUMBER 0.D. WALL MATERIAL TEMP PRESS INSULATION THK LOCATION 2800-003-222-4 3.500 0.300 SA106 0rade B 517 775 CaSilicate 3.0 YC104 2BDG-003- 223-4 3.500 0.300 SA106 0rede B 517 775 CaSilicate 3.0 VC104 2BDG-003-224-4 3.500 0.300 SA106 0rede B 517 775 Cast l1mte 3.0 YC104 2RCS-008-020-1 8.625 0.906 SA376 Type 304 613 2327 Renective 3.3 CS206 2RCS-008-021-I 8.625 0.906 Sn376 Iype 304 545 2327 Redective ' 3.3 CS206 2RCS-008-040-1 8.625 0.906 M376 Type 304 613 2327 Renective 3.3 CS204 2RCS-008-041-1 8.625 0.906 SA376 Type 304 545 2327 ReDective 3.3 CS204 2RCS-008-060-1 8.625 0.906 SA376 Type 304 613 2327 Redective 3.3 CS202 2RCS-008-061-1 8.625 0.906 SA376 Type 304 545 2327 Renective 3.3 CS202 2RCS-014-084-1 14.000 1.406 M376 Type 304 656 2327 ReDective 3.7 CS203 2RHS-010-024-1 10.750 1.125 SA376 Type 316 140 640 Renective 3.7 CS202 2RHS-012-001-1 12.750 1.312 M376 Type 316 613 2327 Renective 3.7 CS206 2 SIS-006-012- 1 6.625 0.718 SA376 Type 316 545 2327 Renective 3.5 CS206 2515-006-015-1 6.625 0.718 SA376 Type 316 545 2327 Renective i 3.5 CS204 2 SIS-006-016- 1 6.625 0.718 SA376 Type 316 545 2327 Renective 3.5 CS202 2 SIS-006-024- 1 6.625 0.718 M376 lyM 316 613 2327 Renective 3.5 CS204 2 SIS-006-025- 1 6.625 0.718 SA376 Type 316 613 2327 Renective 3.5 CS202 2 SIS-012-287- 1 12.750 1.312 SA376 Type 316 545 2327 Redective 3.7 CS202 2 SIS-012-288- 1 12.750 1.312 M376 Type 316 545 2327 Renective 3.7 CS204 2 SIS-012-289- 1 12.750 1.312 SA376 Type 316 545 2327 Renective 3.7 CS206 NOTE:

CASILICATE HAS A 0.010 INCH THICK TYPE 304 STAINLESS STEEL OUTER JACKET PrFLECTIVE INSULATION HAS BOTH AN INNER AND OUTER STAINLESS STEEL LAYER 3-5

TABLE 3.4 WRIPJET PROGRAM PIPE BREAKS LINE NUMBER BREAK NUMBER TYPE WELD LINE NUMBER BREAK NUMBER TYPE WELD 28D0-003-222-4 28D0-005-0-C EA FW 251S-006-015-1 2 SIS- 162-C-L TE FW 2800-003-222-4 28D0-006-0-C EA FW 2 SIS- 164-C-L EA FW 2BDG-003-223-4 28D0-022-0-0 EA FW 2 SIS- 166-C-L EA FW 2BD0-003-224-4 28D0-034-0-C EA FW 2 SIS- 168-C-L EA FW 2RCS-008-020-1 2RCS-004-C-C TE FW 2 SIS- 183-C-L EA FW 2RCS-008-021-1 2RCS-001-C-C TE FW 2 SIS-006-016- 1 2 SIS- 177-C-C TE FW 2RCS-008-040- 1 2RCS-008-C-C TE FW 2 SIS- 179-C-C EA FW 2RCS-008-041-1 2RCS-005-C-C TE FW 2 SIS-006-024- 1 2 SIS- 118-C-C TE FW 2RCS-008-060- 1 2RCS-012-C-C TE FW 2 SIS- 119-C-C EA FW 2RCS-006-061-1 2RCS-009-C-C TE FW 2 SIS-006-025- 1 2 SIS- 147-C-C TE FW 2RCS-014-064-1 2RCS-241-C-C TE FW 2 SIS- 148-C-C EA FW 2RCS-242-C-C EA SW 2 SIS-012-287- 1 2 SIS-032-C-L EA SW 2RCS-243-C-L EA SW 2 SIS-034-C-L EA SW 2RCS-244-C-C EA SW 2 SIS-035-C-C EA SW 2RCS-245-C-L EA SW 2 SIS-036-C-L EA SW 2RCS-246-C-C EA BM 2 SIS-037-C-C EA FW 2RCS-247-C-L EA BM 2 SIS-038-C-L EA FW 2RCS-248-C-C EA BM 2 SIS-039-C-C EA BM 2RCS-249-C-L EA BM 2 SIS-040-C-L EA BM 2RCS-250-C-C EA BM 2 SIS-041-C-C EA SW 2RCS-252-C-C EA BM 2 SIS-042-C-L EA SW 2RCS-254-C-C EA FW 2 SIS-043-C-C EA SW 2RCS-256-C-C TE FW 2 SIS-044-C-L EA SW 2RHS-010-024-1 2RHS-002-C-C TE FW 2 SIS-012-288- 1 2 SIS-016-C-C TE FW 2RHS-012-001-1 2RHS-003-C-C TE FW 2 SIS-017-C-C EA SW 2 SIS-006-012- 1 2 SIS-091-C-C TE FW 2 SIS-018-C-L EA SW 2 SIS- 151 -C-C EA FW 2 SIS-020-C-L EA SW 2 SIS- 152-C-L EA FW 2 SIS-022-C-L EA SW 2 SIS- 153-C-C EA FW 2 SIS-024-C-L EA SW 2 SIS-006-012- 1 2 SIS- 154-C-L EA FW 2 SIS-026-C-L EA BM 2 SIS- 155-C-C EA FW 2 SIS-028-C-L EA BM 2 SIS- 156-C-L EA FW 2 SIS-012-289- 1 2 SIS-003-C-L EA BM 2 SIS-158-C-L EA FW 2 SIS-005-C-L EA BM 2 SIS- 181 -C-L EA, FW 2StS-007-C-L EA BM 2 SIS-009-C-L EA SW CODE:

2 SIS-010-C-C EA SW 2 SIS-013-C-L EA FW TE = TERMINAL END BREAK 2 SIS-011 -C-L EA SW l 2 SIS-012-C-C EA FW EA = EXCEEDS THE STP.ESS AND/0R FATIGUE FACTOR THRESH 0LD FOR BREAK EA 2 SIS-013-C-L FW l l I I ~l FW = A FIELD FABRICATED WELD IS MADE AT THIS POSTULATED PIPE RUPTURE LOCATION I- 1 I I I i

SW = A SHOP FABRICATED WELD IS MADE AT THIS POSTULATED PIPE RUPTURE LOCATION l l I I I l

BM = BASE METAL. THERE IS NO WELD AT THIS POSTULATED PIPE RUPTURE LOCATION 3-9

200G-003-223 4 2800-034 0-Cg[/280G-003 224-4 2DDG-022 0.c / 20DG-003-222 4 (2dDG-006-0-c OUTSIDE CONTAINMENT

((

/

/ 25DG-005-0-C

/ l Figure 3-1.steem Generator Blowdown Pipe Breaks

)

l l

! - -- - -_--__- __ _. ___ _____ _ ' ~ ~ ~ ' - _ _ _ . _ _ _ _ _ . _ _ _ _ _ _

2RCS-008-021-1 2RCS-006-020-1 2RCS-001 C-C- - 2RCS-004-C-C TERMWAL DC TERNMAL END O-COLD LEO

' - ^

-O HOT LEG 2RCS-008-041-1 2RCS-008-040-1 8

2RCS-005-C-C- -2RCS-006-C-C TERMNAL De TERMNAL END O-COLD LEG

'~"

-O HOT LEG 2RCS-006-061-1 2RCS-006-060-1 2RCS-009-C-C- -2RCS-012-C-C TERMNAL De TERMWAL END LOOP C COLD LEG HOT LEG Figure 3.2 Reactor Coolant Loop Bypass Line Pipe Breaks 3-11

2 SIS- 155-C-C 2 SIS- 156-C-L - 2515- 181 -C-L 2 SIS- 153-C-C ,

, 251S-006-012-1 2 SIS- 154-C-L 2 SIS- 156-C-L 2 SIS- 151 -C-C 2 SIS- 152-C-L -

2 SIS-091-C-C LOOP A COLD LEG 2RCS-275-3-1 2 SIS- 166-C-L 2 SIS- 183-C-L

/ /

2 SIS- 166-C-L '

%  ; 2515-006-015-1 2 SIS-164-C-L 2 SIS- 162-C-L LOOP B COLD LEG

@ 2RCS-275-6-1 l

2 SIS- 177-C-C l

( '; 2515-006-016-1 I

2 SIS- 179-C-C LOOP C COLD LEG

$ 2RCS-275-9-1 Figure 3.3 Sofety injection Line Pipe Breaks 3-12

2 SIS- 1 18-C-C TERMINAL END LOOPB l 2 SIS- 1 19-C-C 2R S 029-4-1 2 SIS-120-C-L 2 SIS-006-024-1 2 SIS- 148-C-C 2515-006-025-I N

2 SIS- 147-C-C TERMINAL END LOOP C HOT LEG 2RCS-029-7-1 l

l-l Figure 3.4 Safety Injection Line Pipe Breaks 3-13 t

I - -- - - - --

2 SIS-012-C-C  % 2 SIS-010-C-C 2 SIS-013-C-L - - 2 SIS-011-C-L 2 SIS-009-C-L 2 SIS-007-C-L 2 SIS-005-C-L- '

2 SIS-003-C-L 2 SIS-012-289- 1 LOOP A COLD LEG l

2RCS-275-3-1 2 SIS-012-288- 1 2 SIS-028-C-L 2 SIS-026-C-L 2 SIS-024-C-L 2 SIS-022-C-L

[ 2 SIS-020-C-L 2 SIS-017-C-C N

2 SIS-018-C-L J 0

0 2 SIS-016-C-C TERMINAL END 2RCS-275-6-1 2 SIS-043-C-C 2 SIS-044-C-L 2 SIS-041-C-C 2 SIS-042-C-L 231S-012-287-1 2 SIS-039-C-C 2 SIS-040-C-L 2 SIS-037i C-C 2 SIS-038-C-L 2 SIS-035-C-C 2 SIS-036-C-L 2 SIS-034-C-L /

/

2 SIS-032-C-L u

LOOP C COLD LEG l

2RCS-275-9-1 ,

Figure 3.5 Accumulator injeClion Line Pipe BreekS 3-14

PRESSURIZER SURGE LINE LOOP C HOT LEG 2RCS-241-C-C 2RCS-029-7-1 TERMINAL END 2RCS-242-C-C 2RCS-014-084-1 4

2RCS-243-C-L 2RCS-244-C-C 2RCS-245-C-L 2RCS-246-C-C N2RCS-256-C-C TERMWAL E@

2RCS-247-C-L 2RCS-246-C-C 2RCS-249-C-L 2RCS-250-C-C d b2RCS-254-C-C

- 2RCS-252-C-C l

l Figure 3.6 Pressurizer Surge Line Pipe Breaks 3-15

9 2RHS-010-024-1 E

M 2RHS-002-C-C g FROM ACCUMULATOR PRIMARY LOOP LOOP A HOT LEG 2RCS-029-3-3 F 0 9

2RHS-003-C-C TERMWAL END

/

2RHS-0l2-001-g Figure 3.7 Residual Heat Removal Line Breaks 3-16

SBCTION 4 SCREENING CONSIDERATIONS Five systems (Steam Generator Blowdown, Feedwater, Reactor Coolant, Residual Heat Removal, and Safety Injection) were evaluated through a screening process designed to eliminate from consideration any piping which exhibits the potential for:

o Ancmalous conditions: excessive vibration, flow stratification, failure of pipe fittings, failure of equipment supports, and uneven mixing o Fluid transients (water hammer) o Stress corrosion cracking The overall results were summarized in Section 3. The following sub sections discuss each of these items in detail for the four systems remaining in the NHIPJET program.

4.1 ANOMALOUS CONDITICNS 4.1.1 Flow Stratification Flow stratification is a thermal fluid phenomenon identified in IE Notice 84-87. It potentially occurs in large diameter fluid systems where the following circumstances exist: (1) the piping system is long and horizontally oriented. (2) the pipe is filled with hot (or cold) fluid flowing very slowly, and (3) much colder (or hotter) fluid is introduced at some point upstream of the horizontal run at a rate by which mixing does not occur.

' This phenomenon was evaluated for those systems within the final scope of NHIPJET systems and was concluded to have no potential for application. The safety injection system is a predominantly

' cooler, small diameter, high flow system. The steam generator blowdown system has a smaller diameter than other saturated steam and water systems, and there are no small sources of cold water which can bleed into the system.

It is appropriate to note that, while the feedwater system is not in the final NHIPJET program, this system was reviewed in detail i

for specific potential for flow stratification and concluded not l to have the sufficient conditions to be a concern. The. potential i

' for backflow from the steam generators is minimized by the main feedwater isolation valve (FIV) and check valve in each line.

Each system within the HHIPJET program was reviewed for its susceptibility to fluid flow stratification. It is concluded that no potential for flow stratification exists. Previous relant experience has s'hown exceptionally high and turbulent flow rates at primary loop branch connections. The Safety Injectioh System is predominant y smaller piping which experience cooler, high 4-1

volume fluid flow rates. The Steam Generator Blowdown piping in-the NHIPJET program is 3 inches in diameter which precludes flow stratification. )

Induced System Vibration  !

4.1.2 Positive displacement pumps are potentially a source of vibration and fatigue failure. BVPS-2 systems remaining ire the NHIPJET program have no positive displacement pumps in service with the exception of the hydrostatic test pump. Vibration from this pump is insignificant because of its low flow and infrequent use.

General system vibration is addressed in the BVPS-2 startup testing program.

4.1.3 Pipe Fittings Consideration f or LBB Analysis A review was conducted to determine the presence of any cast materials for systems evaluated in HHIPJET. Table 4.1 summarizes fitting material specifications. Other than for the large bore primary coolant piping, no cast material is used for fittings within the NHIPJET program high energy systems. Typical PHR piping system fittings include elbows, tees, and branch connections.

The most often expressed concern regarding wrought fittings is for weld cracking in elbows. Elbows may be either seamless or welded. Seamless elbows are generally forged, but may also be made by bending seamless pipe. Helded elbows are usually made from plate materials which are either rolled and welded into pipe sections before being bent into elbows or formed into halves of elbows which are then welded together by longitudinal welds.

Service experience with elbows has been excellent except for a few instances of cracks caused by a defective longitudinal elbow weld.

All of the BVPS-2 lines remaining in the HHIPJET program have only forged seamless fittings. Radiographic, ultrasonic, or liquid penetrant examinations were performed by the manufacturer in accordance with ASME Section III. NB-2550 for Class i seamless materials, thus assuring material soundness. Satisfactory experience combined with the type of elbows and NDE checks.

reduce the probability of any longitudinal cracking of the elbows. Also, the material properties for the fittings are co:nparable to those in the wrought pipe itself.

Additionally, each line of piping was evaluated to determine limiting stress locations (i.e.. potential break locations).

Thus, the limiting locations where NHIPJET is being applied correspond to those determined to be the highest stress points as dictated by the ASME Code and standard piping analysis procedures. These procedures consider the entire piping system including pipe fittings.

4-2

4.1.4 Equipment Supports Evaluation The potential for the failure of equipment supports resulting in the subsequent failure (rupture) of attached high energy piping exists if the supports have not been seismically qualified.

Since ASME Section III piping systems terminate at ASME Section III Seismic Category I components, seismic qualification is assured. For non-nuclear safety (NNS) high energy piping systems, terminations may be to non-seismic qualified components.

i However, the BDG high energy system at BVPS-2 is qualified f or

' normal plus safe shutdown earthquake (SSE) inertia loadings. The portions of this system within NHIPJET scope had no component terminations, and no component support qualification has been required.

4.1.5 Uneven Mixing Thermal fatigue due to water temperature fluctuations caused by slow, continuous mixing of water flows with different temperatures, has been reported to occur for certain BNR plants. '

BVPS-2 HHIPJET systems were reviewed and concluded to be free of causative factors leading to this phenomenon.

4.2 FLUID TRANSIENTS 4.2.1 Introduction BVPS-2 evaluated the potential for flow transient events in accordance with applicable Stone and Hebster Engineering Corporation (SWEC) and NRC quidelines. A review by fluid system was conducted. Applicable flow transients were identified and the transient description and design information used in the pipe stress evaluation. Potential water hammer sources considered for the design of BVPS-2 piping systems were based on industry experience and the concerns presented in various NUREG's, including specifically NUREG-0582 and NUREG-0927 [4.1, 4. 5 ].

Hater hammer caused by steam voids, condensation, flashing and thermal mixing are controlled through system design, operating procedures, and operator training.

Since NUREG-0582 was issued, additional NUREG*s that address the issue of water hammer were published [4.2 through 4.6]. These documents recognize that water hammer events are inevitable and provide guidance to eliminate or reduce the frequency and effects of these events. They also acknowledge that water hammer is not as significant a safety issue as previously anticipated since the resulting damage from water hammer events has been limited to 4-3

piping and equipment supports. Table 4.2 is reproduced from Table 3-2 of NUREG-0927 and provides specific PHR system water hammer causes and preventive measures. The detailed discussion below addresses the BVPS-2 provisions for minimizing steam / water hammer effects and, where applicable, addresses specific details associated with that guidance for minimizing water hammer.

Specific flow transients for BVPS-2 were identified subsequent to a comprehensive review of NRC Guidelines and plant experience.

Guidance was obtained from NUREG-0582 [4.1], which is a comprehensive summary of water hammer and fluid system transients. Appendix A contains a working summary of and comments on fluid transients evaluated against NUREG-0582, and as appropriate, included in the Plant's design basis for pipe stress analysis.

4.2.2. Provisions for Minimizing Water Hammer Effects Systems within Hertinghouse scope of supply are not in general susceptible to water hammer. The reactor coolant, and residual heat removal systems have been specifically designed to preclude water hammer. Operating experience at other plants with Hestinghouse systems have verified this design approach.

Nestinghouse has conducted a number of investigations into the causes and consequences of water hammer events. The results of these investigations have been reported to Hestinghouse operating plant customers and have been reflected in design interface requirements to the BOP designer for plants under construction, to assure that water hammer events initiated in the secondary systems do not compromise the performance of the Hestinghouse-supplied safety-related systems and components.

In general, the approaches taken, individually or in combination, to address water hammer concern were to prevent or minimize water hammer effects through system design features and operating procedures.

Potential water hammer sources to be considered were based on industry experience and the concerns presented in various NUREG's. The following discusses in more detail the i potential water hammer sources, if any, that were considered in the design of the subject systems and the actions taken to minimize and prevent water hammer effects.

4.2.2.1 Reactor Coolant System (RCS)

There is a very low potential for water hammer in the subcooled j water solid portions of RCS since these portions of RCS are designed to preclude void formation. Relief valve discharge loads associated with the pressurizer have been specifically identified and analyzed for BVPS-2. (NUREG-0927) t 4-4

4 4.2.2.2 Safety Injection System (SIS)

As discussed below, it is considered unlikely that water hammer could occur in the Safety Injection System. The low temperature

, SIS lines, which are normally-water solid, have a very low probability of steam void formation. Proper initial fill and venting ensures that low and high head safety injection system piping remains filled. In addition, the head of water provided by the RHST provides a continuous mechanism for ensuring that the low head safety injection system lines remain full.

For the SIS lines which are part of the Residual Heat Removal System return flow path, operating procedures for RHS minimize the potential for water hammer in these lines.

For the SIS lines which are part of the Reactor Coolant Pressure Boundary to the first isolation valve, there is a very low potential for water hammer as indicated in the above RCS discussion.

4.2.2.3 Residual Heat Removal System (RHS)

Portions of the RHS piping is high energy because it is normally pressurized by the RCS or SIS during normal plant operating conditions. When RHS is operating (i.e., the short operational period), valve closure times and operating procedures minimize

! the potential for water hammer. Proper fill and venting will initially ensure that air does not become trapped in any part of the RHS during start-up. Additionally, just prior to RHS 4

initiation, the RHS will be cross-connected with the Chemical and Volume Control System (CHS). This action utilizes the pressure head in the CHS to collapse any voids (should they remain) prior

!, to opening the RHS suction valves from the RCS.

J Nhen the RHS system is not operating, the normally pressurized portions of the system are water solid and are either at a low temperature or subcooled. RHS voiding, therefore, has been addressed by a combination of operator training and startup procedures which provide for the complete filling and venting of

! the system before operation. (NUREG-0927) 4.2.2.4 Steam Generator Blowdown (BDG)

The BDG system is designed to preclude flow transient events.

The isolation valves are air-operated and have an opening / closing time of 10 seconds which would create negligible flow transient loads.

I l

4-5

l l

4.3 STRESS CORROSION CRACKING 4

4.3.1 Introduction The pipe break locations which were determined to have essential targets and passed the material and fatigue screening criteria were reviewed further to determine if they were susceptible to stress corrosion cracking (SCC). These break locations are either in the primary side, reactor coolant environment, or, in the secondary side, steam generator blowdown environment.

Austenitic stainless steel is used in the primary side environment systems (RCS, SIS, RHS) while ferritic (carbon) steel, is used in the secondary side environment. The review concluded that based on specified chemistry, cleanliness, fabrication, and operating controls, and successful operating ,

experience, SCC is not expected to occur at these postulated pipe break locations. Figure 4.1 presents an overview of the corrosion-related review conducted f or BVPS-2, 4.3.2 Primary Side Environment The NHIPJET program pipe break locations in the primary side environment are in the RCS, RHS, and SIS branch connections to the reactor coolant loop piping. Type 304 and Type 316 austenitic stainless steel piping materials are used at these locations. These materials when used in other than the solution annealed condition are susceptible to SCC when exposed to three simultaneous conditions: high tensile stress, high temperature, and a corrosive environment.

Controls were used in the manufacture and fabrication of this

piping to minimize the material susceptibility to SCC by limiting l

both cold work effects. (i.e., strain hardening) and sensitization effects (i.e. chromium-carbide precipitation in the grain boundaries). Cold working of the material was controlled t

by requiring the original manufacturing process of the pipe to L

include solution annealing and by also requiring subsequent l forming operations such as pipe bending to be done with bend radii greater than or equal to five pipe diameters for cold i

bending and by re-solution annealing all hot bent piping. Piping sizes greater than 2-inch nominal pipe size included in the NHIPJET program were hot bent. Hence, the subject stainless steel piping was furnished in the solution annealed condition prior to welding. Helding-induced sensitization was controlled by limiting the weld interpass temperature and weld heat input.

The weld interpass temperature was limited to 350 degrees F maximum, and the weld heat input was limited to 50 kJ/ inch maximum. These controls were evaluated to the requirements of ASTM A708 and were judged to be effective in minimizing weld-induced sensitization. In summary, the measures taken to minimize materi&1 susceptibility to SCC by controlling the material condition followed the recommendations of Regul'atory Guide 1.44 and help to prevent SCC from occurring.

l 4-6 l

Since welding does result in a limited degree of material sensitization further controls were used to prevent SCC from occurring. Of the three simultaneous conditions required for SCC, preventing a corrosive environment was used at BVPS-2.

The primary controls are provided through adoption of proven water quality standards and pipe inside diameter (ID) and outside diameter (OD) cleanliness requirements. The reactor coolant chemistry is described in the FSAR, Table 5.2-5 which is included in this report as Table 4.3. This chemistry control has proven to be effective in preventing SCC due to the control of oxygen to less than 0.1 ppm when operating at temperatures above 180 F and by controlling chlorides to less than 0.15 ppm at all times. The oxygen control is accomplished by using hydrazine to control the initial oxygen content to less than 0.1 ppm when operating above 180 F and by a hydrogen overpressure during normal plant operation. Chloride control is accomplished by using strict purity requirements for the procurement of reactor coolant chemical additives. Monitoring of the reactor coolant chemistry assures a corrosive environment will not be formed.

Pipe ID and OD cleanliness is controlled in accordance with the recommendations of Regulatory Guides 1.37, 1.38,and 1.39. Swipe testing of the pipe OD for contaminants provides assurance against externally initiated SCC. In addition, the use of thermal insulation in accordance with the recommendations of Regulatory Guide 1.36 helps assure against OD initiated SCC.

Industry experience with these materials has been reviewed in NUREG-0679, NUREG-0691, and NUREG-1061 [4. 7 to 4. 9 ]. SCC has neither been reported nor is it expected under the specified operating conditions.

4.3.3 Secondary Side Environment The NHIPJET program pipe break locations in the secondary side environment are in the high energy portions of the BDG piping.

SA 106, Grade B ferritic mild carbon steel is used at these locations. The BDG fluid is supplied from the main feedwater system and is described in the FSAR Table 10.4-13 which is included in this report as Table 4.4. A higher solids content in the BDG fluid is expected as the system's purpose is to minimize solids accumulation on the steam generator tube sheets. The carbon steel has been proven to be resistant to stress corrosion cracking in this environment due to these strict fluid chemistry requirements (Iow oxygen, low chloride, low caustic, and high pH), and by adherence to strict cleanliness controls as recommended in Regulatory Guides 1.36, 1.37, and 1.38 as previously discussed.

Industry experience with this material has been reviewed in NUREG-0679 and NUREG-0691. SCC has neither been reported nor is it expected under the specified operating conditions.

4-7

4.3.4 Conclusion Based on the strict fluid chemistry, cleanliness, fabrication, and operating controls, and successful plant operating experience, the subject break locations are judged to be not susceptible to stress corrosion cracking. Therefore, these break locations pass this WHIPJET program screening criteria.

REFERENCES 4.1 Water Hammer in Nuclear Plants. NUREG-0582. July 1973.

4.2 Evaluation of Water Hammer in Nuclear Power Plants.

NUREG-0927 March 1984, 4,3 Compilation of Data Concerning Known and Suspected Water Hammer Events in Nuclear Power Plants. NURE0/CR-2059 May 1982, 4,4 Evaluation of Water Hammer Events in Light Water -

Reactor Plants. NUREG/CR-2781 July 1982, 4,5 Prevention and Mitigation of Steam Generator Wate-Water Hammer Events in PNR Plants. NUREG-0918 Nc; ember 1982, 4,6 Regulatory Analysis fer USI A-1. ' Water Hammer'. VJREG-0993 March 1984 4,7 Pipe Cracking Experience in Light Water Reactors. NUREG-T 0679 August 1980, 4,8 Investigation and Evaluation of Cracking Incidente in Piping in Pressurized Water Reactors. NUREG-0691 September 1980, 4,9 Report of the U m S , Nuclear Regulatory Commission Piping Review Committee Evaluation of Potential fcr Pipe Breaks. NUREG-1061. Vol. 3 November 1984, 4-8

TABLE 4.1 PIPE FITTING MATERIAL USED FOR HHIPJET SYSTEMS SYSTEM PIPE CLASS FITTING SPECIFICATION BDG 901 SA234NPB RCS 1502 SA403HP316 RCS 2501R SA403HP316 SIS 1502 SA403HP316 RHS 1502 SA403NP316 SA234HPB Carbon steel, wrought fittings SA403MP316 Stainless steel, wrought fittings e

4-9

TABLE 4.2 PWR WATER HAIMER vf5?EM PRIMARY CAUSES 0F WATER HAMMER PREVENTIVE HEASURES (*)

DESIGN PLANT OPERATION Feed- Feedwater Cortrol FCV Design veri-w a ter valve (FCV) Over-sizing & Instability fication (3.6)

Unknown and Operator Operating Procedures Error Induced Stees (3.12), Operator Subble Collapse Training (3.11)

Main 6 team Hammer (Value Steam include Valve Closure) Closure Loads in Pipe Support and Component Design Basis (3.9)

Relief valve include Relief Olsenarge Valve Discharge Loads in Pipe Sup-port and Components Design Basis (3.10)

Steam Water Entrain- Operating Proceeures ment, Unknown (3.12), Operator Train-

^ ing (3.11)

Reactor Relief Valve include Relief Valve Coolant Olsenarge Olsenarge Loads in (Pres- Pipe Support and surizer) Components Design Basis (3.10)

RHR voiding Venting (3.3) Operating Precedures (3.12), Operator Training (3.11)

ECCS Voiding venting (3.3). Operating Procedures void Detection (3.1) (3.12), Operator Training (3.11)

CNCS Steam Subble Col- Operating Procecures lapse or Vibration (3.12). Operator Training (3.11)

Essen- Voiding Venting (3.1),

tial Filling Essential Cool-Filling Essential ing Water (3.4), Oper-Cooling Cooling water Water ating Procedures (3.12).

(3.4), Analysis Operator Training (3.4) (3.11)

Stean Line voicing ETP AS8 10-2 Gener- Follcwed by Steam BTP ASS 10-2 Provisions ator Bubble Collapse Provisions (3.13): (3.13): Testing. Keep-Top Disenarge. ing Line Full. Auto-Short Line matic AFW InfT14 tion Lengths, External Header (B&W Only) i l

(*) Refers to section of this report providing details of oreventive measures.

4-10

(Fran BVPS-2 FSAF)

TABLE 4.3 REACTOR COOLANT CHEMISTRY SPECIFICATION Electrical conductivity Determined by the concentration of boric acid and alkali present.

Expected range is <1 to 40 yMhos/cm at 258C.

Solution pH Determined by the concentration of boric acid and alkali present.

Expected values rangg between 4.2 (high boric acid concentration) and 10.5 (low boric acid concentration) at 25sC.

Oxygen'18 0.005 ppm, maximum Chloride (28 0.15 ppm, maximum Fluoride r2) 0.15 ppm, maximum Hydrogen'38 25-50 cc (STP)/kg H 2O Suspended solids 848 1.0 ppm, maximum pH control agent (Li70H)'58 0.7-2.2 ppm as Li Boric acid Variable from 0-4000 ppm as B Silica (58 0.2 ppm, maximum AluminumiS8 0.05 ppm, maximum Calc'ium888 0.05 ppm, maximum Magnesium 5 0.05 ppm, maximum NOTES:

1.

Oxygen concentration must be controlled to less than 0.1 ppm in the reactor coolant at temperatures above 180'F by scavenging with hydrazine or by maintaining the proper hydrogen concentration.

During power operation with the specified hydrogen concentration l maintained in the coolant, the residual oxygen concentration must not exceed 0.005 ppm.

, 2. Halogen goncentrations must be maintained below the specified values at all times regardless of system temperature.

4-11 l

l l

(From BVPS-2 FSAR)

TABLE 4.3 (Continuec')

NOTES: (Cont)

3. Hydrogen must be maintained in the reactor coolant for all plant operations with nuclear power above 1 MWt. The normal operating range should be 30-40 cc/kg H2 0,

4. Solids concentration determined by filtration through filter having 0.45 micron pore size.

5. The specified limits for lithium hydroxide must be established for prestart-up testing prior to hearup beyond 150*E. During cold hydrostatic testing and hot functional testing, in the absence of boric acid, the reactor coolant limits for lithium hydroxide must be maintained to provide inhibition of halogen stress corrosion cracking. Upon plant restart, the lithium hydroxide limits should be established at 180*F.

6. These limits are included as recommended standards for monitoring coolant purity. Establishing coolant purity within the limits shown for the species is judged desirable with the current data base to minimize fuel clad crud deposition which affects the corrosion resistance and heat transfer of the clad.

1 e

e

. e e

d*

4-12

(From BVPS-2 FSAR)

TABLE 4.4 STEAM GENERATQR STEAM SIDE AND FEEDWATER CHEMISTRY SPECIFICATIONS Cotd Hydro /

Cold llot Functional /

Wet Hot Shutdown /

1.avup Hot Standby Start-up From Hot Standby Normal Power Oper.a._t_t .o.n_. --

Chemtstry Blowdown Blowdown Feedwa.ter Parameter Blowdown Feedwater Otowdown Control Control Expected Expected Control Control Expected Expected Control Control E xp_ecle_d .

pH e 25*C 10.0 - 8.8 - 8.8 - 8.8 - NA 8.5 - 8.5 - 8.8 - NA 10.5 9.2 9.2 10.0 8.5 - R.5 -

10.0 10.0 9.2 9.0 9.0 free hydrouide ND O.15 <O.15 NA NA O.15 <0.15 NA NA as ppm CACO O.15 <O.15 Catton conduc- NA 2.0 <2.0 NA NA ttvtty 7 <7 NA NA 2.0 <2.0 whos/ cia

• 25'C I

N Total conduc- NA NA NA NA NA NA NA ttvity 54 NA NA NA nihos/cm 9 25'C Sodium. ppm NA NA <O.1 NA NA NA <O.5 NA NA NA <0.1 Chloride. ppm <0.5 NA <0.15 NA NA NA <0.5 NA NA NA <0.15 NHs, ppm As pH NA 50.5 NA NA NA <10.0 requires 50.5 NA NA 50.25 HyJeartne. ppm 75 - NA NA [Dej + [Dej + NA NA [Dej 4 [Dej +

150 NA NA 0.005 0.005 0.005 0.005 0essolved <100 NA <5 oxygen, ppb

<100 <100 NA <S c5 <5 NA <5 SIOe, ppm NA NA <t.O NA NA NA <5 NA NA NA <t.O Fe. ppb NA NA NA <100 NA NA NA <10 NA NA NA Cu. ppb NA NA NA <50 NA NA NA <5 NA NA NA Suspended NA NA <1 NA NA NA NA NA NA solids, ppm NA <t.O Blowdown NA NA As NA rate. gom/SG NA Maximum Maximum NA NA As <1.0 required required to main-tain cortt ro l parameters

YlBRATION .

REYlEW I

PASSED 1

FERRITIC STEEL AUSTENITIC STEEL e 8DG e RCS e SIS e RHS HYDROXlDES HYDROXIDES ABSENT ABSENT OXYGEN OXYGEN ABSENT ABSENT SULPHER NITRATES reduced forms ABSENT ABSENT CAUSTICS HALOGENS e Chlorides (high pH) e Fluorides ABSENT ABSENT CARBON ^

Threshold -

DIOXIDE

< 180'F ?

ABSENT YES l

t ..

PASS Figure 4.1 Corrosion Review 4-14

SECTION 5 CRACK PROPAGATION ANALYSIS A fatique crack growth analysis is performed at all postulated rupture locations which could result in pipe whip and jet impingement targeting of enmential safe shutdown equipment. Not all break locations are explicitly analyzed: instead, at a minimum, at least one break location for each pipe size is analyzed, and these particular breaks correspond to the highest fatigue usage in that size. This analysis is done by assuming that a material, flaw of a size permitted by the acceptance criteria of ASME Section XI is present at this piping location.

This postulated flaw is then analytically subjected to the internal piping loads which occur at each specified rupture location and analytically examined in order to determine the maximum potential for growth of the flaw into a through-wall crack of a size which would, at some further point. present risk of unstable growth resulting in a complete pipe severence. The process for performing the crack growth analysis involves determination of the through-wall stress distributions. stress intensity factor calculation. a crack growth law, and a postulated defect size. These items will each be discussed separately.

5.1 STRESS DETERMINATION 5.1.1 Thermal Transient Development The NHIPJET Program utilizes ASME Class 1 piping analyses.

Thermal transient analysis of ASME Class 1 piping is performed in accordance with the ASME Section III requirements for pipe stress analysis where the system design specification defines all thermal transients associated with anticipated plant operating conditions. The resulting piping loads are used for the NHIPJET review of crack growth. The ASME Class 1 piping systems associated with the NHIPJET crack propagation analysis are the reactor coolant system, the safety injection system, and the l

residual heat removal system.

The ANSI B31.1 piping analyzed in the WHIPJET program are the steam generator blowdown lines 2BDG-003-222-4. 2BDG-003-223-4 l

and 2BDG-003-224-4 The thermal transients which could potentially result in the development of fatigue cracks growing l

in this piping are determined using a method similar to a Class 1 analysis but making many conservative assumptions.

To facilitate the analysis, the piping system is subdivided into appropriate thermal zones. The temperature conditions experienced by the piping in each zone is detailed in the Stress Analysis Data Pabkage which provides input to the pipe stress analysis. Where appropriate, the transients for each piping zone are reviewed and combined into more generalized transient 5-1

envelopes. The ratas of temperature change are conservatively estimated as stepwise changes for this evaluation. Only significant temperature changes which exceed 25 degrees are considered.

5.1.2 Pipe Stress Evaluation The state of stress at the piping locations selected f or NHIPJET analysis is a combination of piping loads (i.e. forces and moments) and internal thermal induced stresses. The method of analysis used to develop these loads is the same analytical approach wnich satisfies the ASME III Section NB 3600 Code requirements for Class 1 piping analysis. Therefore, for those piping segments which are Class 1 the pipe stress data provided as input to the NHIPJET piping review come directly from the final SNEC pipe stress calculation results.

For ANSI B31.1 piping a separate calculation is performed to assemble the necessary stress terms since this level of analysis in not normally required to qualify the piping. The calculation method employs the same analytical techniques used in the Class 1 analymim.

The piping loads are evaluated using a SNEC computer progran NUPIPE. which uses a finite element stiffness method of analysis.

In the analysis. the effects of individual static and dynamic loading conditions are input and a set of forces and moments are developed for each piping location. Thermal expansion (including anchor displacement). piping deadweight, and the seismic inertia (including seismic induced anchor movements) are the loading conditions analyzed.

All of the piping thermal load cases are examined in order to

! determine the dominant thermal case which represents the most l likely thermal state of the piping. The forces and moments of this representative thermal expansion load case are then added algebraically to the corresponding forces and moments for the deadweight load case.

For ASME Section III NB 3600 thermal transient analysis a SNEC computer program "HTLOAD" which uses a one-dimensional finite difference method of calculating piping through-wall temperature variations is used. The resulting linear ( AT1) and nonlinear

( A T2 ) portions of the temperature differential predicted to occur in the pipe wall resulting from a given change in internal fluid temperature due to a thermal transient is computed. The A T1 term produces the linear portion of the through wall stress compressive on one side, tensile on the other, depending on the direction of the temperature change. The nonlinear A T2 term I

produces the inside skin stress which is seen to become negligible a short distance from the inside piping surface and would contribute minimally to the overall through-wall stress in the cracP growth evaluation. This is discussed in greater detail next.

5-2

l. -, . - _ -

In addition to these terms, the gross discontinuity stress (related to TA and TB) would be considered at any structural and material discontinuities such as branch connections and dissimilar metal interfaces. HTLOAD addresses these discontinuities using a conservative one dimensional analysis which predicts a peak value for the temperature differential existing between the two dissimilar materials or geometries.

5.1.3 Through-wall Stress Distribution Approximation The pipe through-wall stress distribution needed to grow a defect by fatigue is characterized by a tensile stress at the inner radius and a compressive stress at the outer radium. The through-wall stresses used in this program are approximated from the NUPIPE piping analysis results for the appropriate transients resulting in fatigue usage. A linear stress distribution is assumed for simplicity which is broken down into membrane and bending stress components following the approach used in Appendix A of Section XI of the ASME Boiler and Pressure Vessel Code.

The determination of the through-wall stress distribution involves two Section III stress analysis approximations. The first item is the use of "K" indices for peak stress concentration, and the second is the breakdown of the true thermal distribution into a linear through-wall bending moment and a nonlinear portion which has a zero average value of stress and bending moment but account for an increased inner wall stress addition (which falls off very rapidly).

5.1.3.1 "K" Stress Indices The "K" indices are essentially stress concentrations of the inner pipe wall used for the internal pressure, moment loading, l

and thermal loading components of the Equation 11 (NB 3600) peak stresses. The breakdown and definition of terms are as follows:

Po Do

1) Pressure K1 C1 (PO) 2 t Do i 2) Moment: K2 C2 ----

Mi I

(MI) 2 I 1

3) Thermal: K3 E a AT1 (DELT1) 2(1 - V) 5-3 l___ - - - - - - - - - - - - - - - -

4) Thermal: K3 C3 Emb aaTA - a bTB (TA-TB) 1

5) Thermal: E a A T2 (DELT2) (1 -v) where:

Po = internal pressure Do = outside pipe diameter t = pipe thickness Mi = bending moment I = moment of inertia v = Poisson's ratio E = Young's modulus a = thermal coefficient of expansion aa = thermal expansion coefficient for material A ab = thermal expansion coefficient for material B A T1 = linear temperature distribution ,

A T2 = nonlinear temperature distribution TA = range of average temperature on aide A of discontinuity TB = range of average temperature on side B of discontinuity The "C" indices are secondary stress indices which account f or more global stress discontinuities. The "K" indices are stress scaling factors which account for local geometric discontinuities at welds or transition welds in Class 1 piping. The factors which make up the indices are a function of reducer / transition length, pipe diameter and weld condition as shown in Table 5.1 which in Table NB-3683.2-1 from the ASME Code. Minter 1972 Addenda. Mhen the weld is flush with the base metal, the "K" indices are less than the "as-welded" case. In each case analyzed in the NHIPJET program, the break location was modeled as a girth butt weld or a transition weld. The resulting "K" values ranged f rom 1.2 to 1.8 No brea% location was ammociated with either curved pipe, branch connections or butt welding tees.

Therefore. no " global" geometric effects to the "K" indices are experienced beyond the inside pipe walls.

When the geometry of the weld or base metal transition zone is changed (e.g. a postulated crack 10% through the wall) the underlying assumptions which validate the "K" stress indices are I altered radically. Foremost of the assumptions pertain to the I

fundamental characteristic of a peak stress that is very local in nature and affects only a small volume of material. In addition, the crack propagation (fracture mechanics) methodology converts the global piping stresses to local. inside-wall stress intensity factor values at the crack tip, thereby accomplishing the intent of the surface-effect "K" stress indices. Therefore, the "K" indices do not, provide a meaningful measure of actual stress conditions for the case involving postulated part-thro, ugh-wall cracks. The "C" indices are maintained, but argument could be given for their deletion also.

5-4

In summary, the reasons for minimizing the effect that "K" indices have on Equation 11 peak stresses are:

1 "K" indices are " surface effect" stress intensifications.

MHIPJET postulates fatigue cracks which begin at greater than 10% through the wall.

2. The local geometric discontinuity resulting in any of the prescribed "K" values is not significant once a crack has formed and penetrated an appreciable distance through the weld or base metal.

3 The fatigue crack propagation (fracture mechanics) method used in the HHIPJET program intensifies and localizes stresses at the postulated fatigue crack tip.

5.1.3.2 A T2 Nonlinear Stress Effect As indicated earlier, the actual through wall non-linear temperature gradient is modeled in Equation 11 calculations by two separate thermal parameters. One is a temperature difference between the outside surface and the inside surface based on an equivalent linear distribution and is termed DELT1 The second is a temperature difference between the surface and the equivalent linear distribution. Figures 5.1 and 5.2 give a graphical representation of this discussion.

The AT2 ( DELT2 ) stress component is a non-linear portion which has a zero average value and a zero first moment with respect to the mid-thickness. If integrated over the entire wall thickness, the net effect to stress would be zero. Its effect is'seen most dramatically on the inside wall surface and falls off sharply as the outer wall is approached.

The value of DELT2 is computed as follows: it is the maximum of

1. To - T -

0.5 DELT1l 2

lTi-Tl -

0.5 lDELT1l 3 O where T = Average temperature distribution ~

Ti = Inside surface temperature To = Outside surface temperature 9

5-5

On the inside wall. DELT2 is a fraction of the A T1 (DELT1) value and is an additive stress. Near the mid-thickness region, the DELT2 stress, while remaining a small fraction of DELT1 become.s a st'ress which counteracts the DELT1 stress. Near the outside wall, the DELT1 stress is compressive and DELT2 makes it less compressive. Neither stress greatly adds to crack growth at the outside wall.

For example, assume the inside peak DELT1 stress is 10.000 pai.

The corresponding DELT2 stress is typically 2.000 pai (see Figure 5.3). The total inside thermal stress is 12.000 psi. At the midpoint. the DELT1 stress is 0 while the DELT2 stress is compressive by a value of approximately 2.000 psi. At the 0.60t (t = wall thickness) point (the maximum crack growth allowed in the NHIPJET program), the total stress is -1000 pai.

The analysis used in the NHIPJET program assumes that the DELT1 stress contributes to the total through-wall stress as indicated in the ASME Code and shown in Figure 5.5. However. the DELT2 stress is an additive tensile stress in the first 0.25t only and is, in fact. a stress which tends to close fatigue cracks beyond t94t point and up to about 0.75t. A value equal to 50% of the DELT8 inside wall peak stress (see Figure 5.4), taken up to 0.25t au fa addi tive tensile stress, is conservatively assumed.

BeynpQ this point (i.e. from 0.25t to 0.60t), a value of (10% x DELT2) la used for the stress distribution. Again, because the fatigue cracks are postulated at a part-through-wall depth greater than 0.10t. the crack tip does not experience the skin effect of the DELT2 stress value. In other words. WHIPJET assumes that the DELT2 stress adds a tensile stress, which wculd tend to grow the fatique crack, across the entire pipe' wall.

These approximations are very conservative, but were used for all break locations analyzed. Only the pressurizer surge line (RCS) could not meet the fatigue growth acceptance criteria presented later in this section. In this one case, the through-wall distribution for DELT2 was chosen to be 50% of maximum at the beginning decreasing linearly to zero at 0.22t and remaining equal to zero from there on to 0.60t (which is the maximum amount of crack growth allowed).

5.2 CRACK GRONTH LAN The fatigue crack growth analysis of an austenitic weld or base metal defect is based on compiled experimental crack growth data.

The Battelle/EPRI data base management system. EDEAC '(5.1 )

provides this information. For the present study, the crack growth correlation of James and Jones [5.2) was used. This work was based on the EDEAC data and was recently presented at the 5-6 i

ASME BAFVC Section XI Morking Group on Flaw Evaluation. The crack growth relation for 304 316 and weld metal is given by:

3.3 da/dn = (F)(C)(S) ( A K) wheres F is a frequency factor C is a temperature factors and S is an R-ratio factor determined mas S= 1.0 for R<0 or S = 1.0 + 1.8R for 0 < R < 0.79 or i

S = -43.35 + 57.97R for 0.79 < R < 1.0 This da/dn relationship describes incremental crack growth, where. 6 K is the difference between the maximum and minimum values of the stress intensity factor (Kmax - Kmin) and R in the ratio Kmin / Kmax. The temperature factor. C. is met as a constant corresponding to 600 F -- 2.OE-20 units for da/dn are in./ cycle and for AK units are psi-sar(in.). The da/dn least squares fit is dependent on many factors, and environment is not directly included. To account for a water environment with a low

(<50 ppb) oxygen content, the value for F was taken as 2.00 at 550 F. wheream in air at temperatures less than 600 F it would be 1.00 For comparison purposes PHR environment data were compared with this crack growth law as shown in Figure 5.5 [5.3). The comparison shows good agreement when F = 2.0 as used in the NHIPJET calculations.

For ferritic steels, the ASME Section XI. Appendix A curves for crack growth are used. These curves are shown in Figure 5.6 5.3 STRESS INTENSITY FACTOR CALCULATION j A two pronged approach is taken in calculating the linear elastic stress intensity factors. Values of Kmax and Kmin are computed to determine A K f or use in the crack growth law described in Section 5.2 These values are approximated using cylindrical pipe methods (using both tensile and bending stresses) which promote the growth and extension of an interior semi-elliptic surface crack. The through-wall stresa distributions for L given set of transients are used to determine KI using the methods of Raju and Newman [5 mi). The Raju and Newman analyses determines the stress-intensity factor using influence coefficients derived from three-dimensional finite element elastic stress analysis for a wide range of semi-elliptical surface cracks on the inside of cylindrical vencela. Most importantly, these solutions can be superimposed to obtain stress-intensity factor solutions for pressure and thermal-related loads. For the NHIPJET program, the Raju and Newman stress intensity factor solutions were used to l

1 5-7

i correlate and predict fatigue crack propagation rates under various loading conditions.

5.4 INITIAL CRACK SIZE Section XI of the ASME Code Tablem INB 3514-1 and -2 are used to obtain the initial (allowable) surface crack size. The conservative assumption is that the crack aspect ratio (2c/a) = 6 remains constant as the crack grows radially through the wall is used. For example, for a defect in an austenitic stainless steel pipe with an aspect ratio of 2c/a = 6 the normalized crack depth (a/t) is 0.110 Here the radial depth is a. the circumferential half length is c. and the wall thickness is t.

5.5 ACCEPTANCE CRITERIA FOR FATIGUE CRACK GRONTH For each fatigue crack propagation analysis performed on a part-through-wall defect. certain allowable acceptance criteria must be matisfied as established by the NRC staff. The first criteria are for the maximum allowable defect depth and the second criteria are concerned with the maximum defect length.

' After each transient cycle. the crack grows and the remaining pipe wall thickness (ligament) is compared to the crack-induced plastic zone size. The plastic zone size (r )is determined from the following relationships c 1 AK r = 2 c ( )

g n 2 o yC where:

8 = 6 for plain strain.

AK =

Kmax - Kmin.

o =

l cyclic yield stress (approximated for stainless steel l YC as twice the monotonic yield strength at operating l temperature).

If the plastic zone size is greater than the remaining wall ligament. the fatigue crack propagation computer calculation terminates and a warning statement is issued. As previously stated. this calculation is performed after each cycle. In i addition to the plastic zone size criterion, the fatigue crack depth must not grow beyond 60% through the wall. To verify compliance with this requirement. the 40-year life cycle crack '

growth for each part-through-wall crack was checked to ensure it has not grown more than 60% through the pipe wall. This I

requirement was always met. as well as the plastic zone size check. ,

5-8

There are 2 criteria to be satisfied with regard to the crack length. The length of the fatigue crack must be less than both of the following:

1. Critical size through-wall crack [1.414x(Normal +SSE)]

2 Critical size through-wall crack (Normul+SSE) 2 Again, each defect length is compared to these two crack criteria after calculating the 40-year crack growth. For calculations in this program, all of these criteria were satisfied.

5.6 EXAMPLE In order to show the methodology used to determine the linear through-wall stress distribution in a BVPS-2 Clars i line. node 301 (Break 016) in line number 2 SIS-012-288-1 will be developed as an example. This is a 12-in, austenitic safety injection line with a cumulative usage factor greater than 0. 7 5.6.1 Stress Analysis Requirements First, a node for a given system is identified within the HHIPJET program as requiring a fatigue crack growth analysis based upon a high fatigue usage factor. The following documentation is gathered to identify the applicable stress information:

1 The postulated break location table which includes the applicable node, pipeline, stress calculation number, and usage factor (see Table 5.2),

2. The internal loads at the break location resultir.g from deadweight, thermal, and SSE loads (see Table 5.3).

3 The contributions to the nodal usage factor. the NB 3600 stresses at the break locations, and the stress indices (see Tables 5.4 5.5 and 5.6).

5.6.2 Load Ranges Node 301 has four load cases which contribute to cumulative fatigue usage during the 40-year plant life. These Icad cases can be seen in Table 5.4 along with their contributions to the usage factor. For example, the load range which consists of the l transients 33 and 35 has a usage factor of 0.6452 The total I usage factor is 0.7739 The transient pairs come from the five separate transients listed below: '

1 Load of Flow in Idle Loop (576 F to 498 F - Load Case 32) 5-9 l

2a, Initiation of Cooldown (Part 1) with RHR flow through SIS Nozzle (Load Case 33),

2b. Initiation of Cooldown (Part 2) with RHR flow through SIS Nozzle (350 F to 70F to 350F at 66F/ minute

- Load Came 34),

3 Inadvertent RC System Depressurization (Load Case 35),

4 Loss of Load (Load Case 37),

5.6,3 Equation 11 Stresses and Through-wall Stress Distribution The stresa distribution is a maximum at the inside wall and decreases linearly to a minimum at the outside wall. The inside wall peak stream is unique for each load range and is calculated using the actual NUPIPE Equation 11 peak stress values (see Table 5.5). In each case, the Class 1 stresses were determined without "K" stream indices as described earlier, because the fatigue crack propagation analysin localizes and intensifies nodal stresses at the crack tip, Also, as discussed earlier, because the DELT2 stress is a skin effect, a fractional value of the DELT2 stress was used in calculating the inside wall peak stress, The steady-state condition stress was determined from the internal loads shown in Table 5,3 and the pressure loading,

, For this example, the total inside wall stress (Si) for node 301 (see Table 5,5) for the load came 33-35 is:

Si = PO + MI + DELT1 + (TA-TB) + DELT2 Si = 1,700 + 6,955 + 173,756 + 4,383 + 46,069 Si = 232,863 pai (tension)

These values are calculated as described in Section 5.1.3.1.

This extremely high stress is a summation of the individual peak i

stream components for load case 33-35 All "C" and "K" stress indices (see Table 5.6) are included and the full DELT2 value is used. By contrast, the NHIPJET program total inside wall stresa for node 301 and load came 33-35 is:

7 Si = PO/K1 + MI/K2 + DELT1/K3 + (TA-TB)/K3 + 0, 5 ( DELT2 )

Si = 1,417 + 3,864 + 102.209 + 2,578 + 23.035 i Si = 133,103 pai (tension) i This stress, although still very high, is a more realistic estimate of the actual nodal stresses for this particular transient condition as required for a fracture mechanics crack growth analysis. A similar technique was used for each load Lair,

(

3 5-10 l

' The stress distribution through the pipe is modeled to decrease linearly with a discontinuity at 0.25t (t = wall thickeness ) . The 0.25t point corresponds to the location where DELT2 changes from a 50% to a 10% maximum values recall that this approximation for DELT2 is very conservative (see Figure 5.4).

from 133.103 psi to about 82.000 psi is due toThe thelinear lineardecrease decrease in the DELT1 stress component. These stresses and the I

value at the 0.60t point (which is the maximum allowable crack depth) can be seen in Figure 5.7 1

The final stress manipulation concerns the bending and membrane (tensile) components as suggested in Section XI of the ASME Code.

Appendix A.

l This approach has been shown to be conservative

[5.5). To find the bending component, the stress at the 0.50t point is needed.

This stress is then subtracted from the inside wall stress determined above. For the SIS line. 2 separate bending stresses are required. One for the 50% DELT2 stress and one for the 10% DELT2 stress. When the 50% DELT2 stress is used.

the stress at 0.50t is 30.894 pai (tension): i.e.. the 12.466 psi shown in Figure 5.7 at 0.50t plus 18.428 pai (the difference between 10% and 50% DELT2). Therefore, the bending stress is: '

Sb = 133.103 - 30.894 Sb = 102.209 pai The remaining tensile stress is taken to be the membrane stress (Sm). In this case, the membrane stress is the stress at 0.50t:

30.894 pai. A similar argument is used for the case for a 10%

DELT2 stress value (i.e. beyond 0.25t); the corresponding Sb and Sm values are 102.209 pai and 12.466 pai, respectively.

1 5.6.4 Crack Growth The next step is to take the linear stress distributions for each load range, apply the change in stress to the postulated part-through crack, grow the crack for that load case. and analyze the next transient. After all cycles during the 40-year plant life are analyzed. this process determines the total fatigue crack growth.

In the case of the SIS line. the four load cases were analyzed in a random combination manner (using a random number

! generation procedure) for 280 cycles during the 40-year plant life (see Table 5.4). The change in stress is the difference between the peak inside wall stress described above and the steady-state stress conditions existing in the pipe wall prior to the transient. The steady-state stress is found by combining the loads from Table 5.3 in accordance with the ASME Code. Section i III, thereby maintaining conservatism.

Results for the SIS sample calculation are shown in Table 5.7 For this sample case with a relatively high usage factor. no

! significant cr,ack growth occurs during the 40-year plant life, i

None of the acceptance criteria described in Section 5,5 are violated. -

i

(

5-11 i

1 l

5.7 FATIGUE CRACK GRONTH RESULTS The results for all lines analyzed for fatique crack growth are listed in Table 5,8 All lines pass the crack growth acceptance criteria, The RCS surge line required a slight modification of the DELT2 stresses as mentioned previously. Also, results by Nestinghouse for a generic surge line (not BVPS-2 specific) indicate that fatigue crack growth is limited to less than 40% of the pipe wall (5.6]; the Nestinghouse analysis was performed in much more detail than here and is more realistic (i.e., less conservative).

The results for the BDG lines (ferritic steel) required special  ;

analyses since these lines are non-nuclear safety lines and were '

analyzed using ANSI B31,1 piping design. Figure 5.8 lists tne thermal transients for the BDG line. These thermal transients were then evaluated in a similar manner as if a Class 1 analysis was to be performed. In same cases, transients were reviewed and combined into more generalized transient envelopes which were determined to contribute to fatigue usage. Temperature rate changes were conservatively estimated as stepwise changes, and only significant changes (> 25 F) were considered. The results are listed in Table 5.8 and show that fatigue crack growth is not an issue.

REFERENCES 5,1 Data Base for Environmental Crack Model Development, maintained for EPRI by Battelle Columbus Laboratorten.

5.2 L.A. James and D.P. Jones. " Fatigue-Crack Growth Correlations for Austenitic Stainless Steels in Azr".

Predictive Capabilities in Environmentally Assisted Cracking, edited by Ravi Rungta, PVP-Vol 99, ASME Hinter Annual Meeting Special Publication. Novembe-1985 5,3 G.T. Yahr, A.K. Richardson. R.C. Gwalthey, and H.L.

Server. " Case Study of the Propagation of a Small Flaw Under PHR Loading Conditions and Comparison with the ASME Code Design Life " NUREG/CR-3982, prepared by Idaho National Engineering Laboratory for the U.S.

Nuclear Regulatory Commission. November 1984.

5,4 Raju. I.S. and Newman. J.C.. Stress Intensity. Factor Influence Coefficients for Internal and External Surface Cracks in Cylindrical Vessels". Aspects of Fracture Mechanics On Pressures Vessels and Piping.

PVP-58 American Society of Mechanical Engineers, New York, New York, July 1982, 5-12

5. 5 J. M. Bloom and M. A. Van Der Sluys. ** Determination of Stress Intensity Factors for Gradient Stress Fields."

Journal of Pressure. Vessel Technology. Vol. 99. No. 3 August 1977, pp.477 - 484.

5,6 S. A. Swamy et al . " Leak-Before-Break'of PHR Auxiliary Piping." EPRI Research Project 1757-55 in press.

O 5-13

TABC 5.1 SECTION III SIRESS INDICES TABLE 7ABLE NB.3&S3.21 STRESS INDICES FOR USE WITH EQUATIONS 9.10. AND 11 of NB.3650 Internal FAoment Thermal Pressure Loadingt 'I Loading Corrsonent 8, C, K, 8, C, K, C, K, C,'

Strairt cic+. remote from welds or other dos:entendties 0.5 1.0 1.0(' 3 1.0 1.0 1.0 1.0 Girth burt v. tid benveen straig'it cice or between 1.0 ..

pipe and butt welding components (8 3 (a) flush 0.5 1.0 1.1('I 1.0 1.0 1.1 1.0 1.1 (b) as

• elded r>3/16" 0.5 1.1 1.2('I 0.5 1.0 1.0 1.8 1.0 1.7 (c) as welded t<3/16* 0.5 0.5 1.1 1.2('I 1.0 1.4 2.5 1.0 1.7 Girth fillet eld to socket meld fittings 0.5 slipon flasges, or socket 4velding flar.ges('8 -

0.5 2.0 3.0 1.0 1.5 2.0 1.8 3.0 1.0

'Long tudinal butt we!ds in straight pipe ('I (a) flush 0.5 1.0 1.1('I 1.0 1.0 (b) as we;ded t>3/16* 1.1 1.0 1.1 0.5 1.1 1.2(' I 1.0 1.2 1.3 (c) as welded t<3/16* 1.0 1.2 0.5 1.4 2.5(' I 1.0 1.2 1.3 1.0 1.2 Tase nd raisition io;nts per N3 25St.2 ..

anc Fi;. h a-42*.G.1(' ?

0.5 1.4 1.5 1.0 1.2 1.8 1.0 1.5 1.0 Bran:3 cor.*ectio s cer.NS 3643'83 1.0 2.0 1.7 ('I ('I ('I 1.8 1.7 1.0 Curved c:se or butt.r.e:cing elao..s per 24.r(* 3 ANSI B16.9. AfiSt 816.38 1.0 1.0('I (*I (*I or f.tSS SP48(3 1.0 1.0 1.0 0.5 2(R-r)

Butteeldirq. tees per ANSI B16.9 or IASS SP48(' 'I 1.0 1.5 4.0 I'I ('I 1.0 1.0 1.0 But.v.elding reducers per ANSI B16.9 0.5 or 1.tSS SP48ti ei 1.0 1.5 2.0 1.0 1.3 1.0 1.0 1.0 0.5 l

I 5-14

• i TABLE 5.2  !

1 1

POSTULATED BREAKS IN THE SIS SYSTEM

= - - _ _ .-_ __

BREAK PIPELINE NODE TYPE OF REASON FOR NUMBER NO. BREAK BREAK 2 SIS-016-C-C 2 SIS-012-288-1 301 Circum- Terminal ferential End

==-

STRESS USAGE FACTOR MAX NOR. OPER. CONDITION CALCULATION T(F)

P (PSI) 12241-NP(B)-X70B 0.7739 636 2537 TABLE 5.3 BREAK LOCATION INTERNAL LOADS REFERENCE NUPIPE RUN NUMBER: 4140 DATE: 2/10/83

_=----==_ - =--__=__- _ _ _

NUPIPE INTERNAL LOADS AT BREAK POINT (1)

LOAD CASE NO. Fx Fy F2 Mx My Mz

__= _ _ _ _ - - -

=- .=-

DEADHEIGHT 2 -4207 -292 518 -628 -i218 11159 c.

THERMAL 27 -813 -2951 1868 -18172 10896 -10328 SSE 10 4356 5251 783 3d57 6547 30596 (1) Forces are in Ibf, Moments are in in-lbf 5-15

TABLE 5.4 USAGE FACTOR CONTRIBUTORS LOAD CYCLES USAGE RANGE ACTUAL ALLON FACTOR (1)

(33-35) 20 31 0.6452 (34-37) 80 2723 0.0578 (33-34) 120 1711 0.0701 (32-33) 60 72198 0.0008

_ _ - - = _ _ _ _ - - -

TOTAL 280 (2) 76663 0.7739 (1) USAGE FACTOR = ACTUAL /ALLON (2) 40-YEAR PLANT LIFE t

TABLE 5.5 PIPING STRESSES Equation 11 (Peak Stress, psi)

___- _ _ _ - - - - - - - - - = - - - - - - _________-

PO MI DELT1 (TA-TB) DELT2 TOTAL

= - - . -

-=-

(33-35) 1700 6955 173756 4383 46069 232863 (34-37) 13706 12424 77661 2184 20793 126768 (33-34) 0 3314 105342 2700 24130 135486 (32-33) 12103 5004 46974 410 11660 76151 5-16

TABLE 5.6 STRESS INDICES NODE C1 C2 C3 K1 K2 K3 301 1.10 1.00 1.00 1.20 1.80 1.70 TABLE 5.7 SIS LINE CRACK GRONTH RESULTS INITIAL FINAL INITIAL FINAL a/t RATIO a/t RATIO CRACK DEPTH CRACK DEPTH

---

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

0.11 0.123 0.143" 0.171" 5-17

TABLE 5.8 BVPS-2 FATIGUE CRACK GRONTH RESULTS

==

SYSTEM LINE SIZE LIMITING a/t FINAL SURFACE (INCHES) BREAK INITIAL FINAL CRACK LENGTH LOCATION (INCHES)

SIS 6 091 0.117 0.356 1.53 SIS 6 147 0.117 0.365 1.57 SIS 6 158 0.117 0.343 1.48 SIS 12 016 0.110 0.118 0.93 SIS 12 003 0.110 0.111 0.87 SIS 12 032 0.110 0.117 0.92 RHS 10 002 0.111 0.112 0.76 RHS 12 003 0.110 0.110 0.86 RCS 8 009 0.114 0.115 0.62 RCS 14 241 0.109 0.678

• 5.72 BDG 3 034 0.124 0.132 0.24

==_ -------------

• The acceptance criterion of 0.60 was slightly violated, but detailed results by Hestinghouse (5.6] on essentially the same line indicate crack growth less than an a/t = 0.4C. The highly conservative approach used her e is judged to meet the intent of the NRC requifements.

5-18

i i

1

Io Outside "

kd** \ y = T(vi +- - T m VI2 L ~

7

.ss. ., g, ,-, , 7 inside *'

_.l T;

e,

+;

V12 A

- +

,+-

f A Surfxe .

% y  %

Figure 5.1 DECOMPOSITION OF TEMPERATURF. DISTRIBUTION RANGE

> k.aTz I

~

\

l 1

Tmnerature Dis'tribution e a r, w arc j

Figure 5.2 TCTERATURE DISTulBtTTION i

5-19

12.50 CONTRIBilIION OF DELII AND DELI 210 IHERNAL SIRESS

,x.

! 10.00 A '\S combined AT, and Ar2 stresses

~

g .

7,50 "

-h.

s 5,00" AT2 2.50c stress 0,00

~~ ~.____ _ -

Contribution__$__

s

-2,50-

2 -

~

-5,00 o ib.

-7.50-

-10.00- -

-12.0 t,00 20.00 40.00 60.00 80.00 100.00 WALLRADIUS(%)

Figure 5.3 Conparisons of AT1 , AT,, and 'Iheir Combination 5-20

WHIPJETDELIAT2STRESSAPPR0XIMAIl0H 100.00 80.00 "

60.00 f

ff 40.00"

• MIIPJET Approximation 20.001 l

0.00

-20.00-

-40.00- Actual AT #

2

-60.00"

~

'f 00 20.00 40.00 60.00 80.00 PIPE RADIllS (X)

Figure 5.4 Approximation Used for AT 2 in the Mi1PJET Program 5-21

Aeeuc0 cycuc STRESS INTENSITY RANGC,68Po8 6 78 9 IO 20 30 40 SO 60 80 10 0 i

a a i 1 . i e i i . . i 4

- m<

eo -

w

~

w h ,

C** u y -

o

z. s

• E F = 2.0 E (e (PWR) p 2

E,,o.s  :

x o _

F = 1.0 *

(AIR) E 5 -

so a o e - _

o _

g -

w e -

4 a

=

p e4 _

A<

Actual PWR Emrironment

~

Data Points

~

80-8 g.y 1 1 1 1 I e e f a 1 e t t~

S 6 7 8 9 to 20 30 40 SO 60 80 100 APPLICO CYCLIC STRESS INTENSITY MANGE.ksi8 Figure 5.5 Conparison of Actual PWR Enviroment Data with the Austenitic Stainless Steel Crack.

Growth Law 5-22

l Fig. A-4300-1 SECITON XI - DIVISION I 1986 Edition 1000

/

~

• Linear interpolation is recom-  !

O -

mended to account for ratio dependence of water environment

[

[

500 - curves, for 0.25 <R< 0.65 for 8%

~

shallow slope: p [

i j

8 = (1.01 X 10'll 0 a g1.95 2

+ ,f j l dN & p /

~

02 = 3.75 # + 0.06 4

4 [

R=Kmin /Kmaz /[ w 200 _ [

/

/

Subsurface flaws [

{ 100 -

(air environment)

~ /

i

~ 11 = (0.0267 X 104) A K g3.726 dN  ?

Ip y N ~

- - / ,2 ,a, 50 -

Determire the AKat which the

[ gg i law changes by calculation of g4 g ~ theintersectior of the two

• 4 curws.

8/ [

y Surface flaws c (water reactor

[

)

f 6

20 - environment) applicable for

1. < 0.25

• 0.25 < R < 0.65 f R 3 0.65 10 -

g=gm, /g,,,

/

7 - y l N t I  %

5 - g

• Linear interpolation is recommended p

g f {]

to account for A ratio dependence of water environment curws,for

~

U g

l f (

O 0.25 < R < 0.65 for steep slope:

• l d #.8 = (1.02 X 104) Og AK5.95 U[k j . du

- S/k Oj = 26.94 5.725 I

1. =Kmin /Kream I

/

i i l l I I i 111 l l l l 1Ii1 I 2 5 7 10 20 50 70 100 Stress Intensity Factor Range (AKg ksi [.)

Figure 5.6 REFERENCE FATIGUE CRACK GROWTH CURVES FOR CARBO *1 AND LOW ALLOY FERRITIC STEELS 5-23

IHR0i!GH-WALL SIRESS DISIRIB!!IION FOR LOAD CASE 33-35 168.00 f140.00( .

3 y 120.001 .

5 100.002

.~.,'.

s

' ,, f .25t 0

80.00-K 60.004 \s 4g,gg. \ .,,

s

~

20.00- .

%s 0.00

-. .,)

-20.00t

-40'0

^

s 00 20.00 40.00 60'00 80.00 100.00 WALL RADIl!S (x) l .

l Figure 5.7 I$ rough-wallStressDistributionfor load Case 33-35 5-24

Figure 5.8 STERM GENERATOR BLOWDOWN SYSTEM THERMRL TRANSIENTS FROM FEEDVATER SYSTEM

. s  ::s 2ONE 3

+ t g + +

TO BLOVDOVN T ANK ZONE 1 + + ZONE 2 PIPING TR ANSIENT INITIAL FINAL R ATE OF NUMBER 20NE DESCRIPTION TEMP TEMP CHANGE OF TIMES i  !  !  !

! HEATUP 1 AMBENT 547*F . 100* PER HR i 200 1&2 COOLDOVN 547'T AMBIENT 100' PER HR 200 I&2 # 15% => 100% @ 5% / MIN ' ,

541*F ,_ 517'F , _500' PER HR , 18300 3 100% => 15% @ 5% / MIN , 275'F 390'F ' 100' PER HR 18300

_ t &2 , 15% => 100% @ 5% / MIN 517'F ,

541*F , 100* PER HR , 18300 3 l 15% => 100% @ 5% / MIN ,_ 390'F 275'F ' 500* PER HR 18300 1&2 ! STEP LOAD w/ STEAM DUMP 517'F ,

557'F 3 STEPVISE ,

200 1&2 STEP LO AD w/STE AM DUMP 557*F ' 522'F

• STEPVISE 200 1 &2 , STEP LO AD v/ STEAM DUMP 522*F _, 548*F , _160' PER HR , 200 3 STEP LO AD w/STE AM DUMP 390'F 250'F ' STEPVISE 200

_I.&.2.. l.0_SS OF LO AD/ NO UNIT TRIP' 517'F ,__ 562'F , _ 3' PER SEC , 80

__3 LOSS OF LO AD/ NO UNIT TRIP, 390'F 70*F STEPVISE 80 1&2 , LOSS OF POWER 517'F ,

567'F , 3' PER SEC 40 3 LOSS OF POWER 390*F 70*F

' 100*/SEC 40 1 &2' ,

FULL POVER TRP_ 517'F , 548'F ,,,, 3' PER_ SEC 400 3 FU.LL POVER TRP 441*F 70'F

, ' 100* PER HR 400 I&2 . TEST CONDITIONS  ! 547*F 475*F 154' PER HR 10 1&2 j STEP LO AD w/STE AM OUMP 522*F f 548'F _ f 160' PER HR f200

... 3.,, L ST.EP_ LO AD w /STE AM DUMP,,4,,, .,390'F 250'F STEPVISE 200..

1&2 , LOSS OF LOAD / NO UNIT TRIP'- 517'F 562*F 3' PER SEC 80 3 _ iLOSS OF LO AD/ NO UNIT TRIP' 390*F 70*F STEPVISE 80 1&2 j,~ ' _ LOSS OF POVER Sl ?'F 567'T ,.

' ,,_ 3_* PER SEC,,,,_{__ 40 3 390'F 70*F LOSS _ OF POVER

,__ 100'/SEC i 40 1 &2 ,, FUL.L POVER TRIP 517'F 400 3 . 548'F . _ 3* PER SEC !.

FULL POVER TRIP ^

441*F i 70*F ^ 100* PER HR i 400 1 &2 _ , _ TEST CONDITIONS 547'F _, 475'F , 55d' PER HR { 10 3 j TES(COND,lTIONS  ; 547'F  !

475'F ' 15.4' PER HR 10 1&2 I_ _ _ . BLOVDOVN AMBENT I 547'T i STEPVISE .

200

,3_ BLOVDOVN ,lNTE_RUPTION 390'F AMBENT I

100' PER HR I~100

. 3 ......... BLO,VDOWN ACTIV ATION __ '_AMBENT_, 441*F _L_STE.PVISE,4,,60,O_

3 FEEDy.MER TEMP DECRE ASE , _ _ 390'F 275'F__ ' ,STEPVISE i_ ___2000_

3 FEEDV ATER TEMP INCRE ASE ' 275'F__ 390'F  !

STEPVISE 2000 5-25

SECTION g.

MATERIAL PROPERTY DATA 4

6,1 AUSTENITIC STAINLESS STEEL LINES

' Tensile true stress-strain and J-resistance curve data for the applicable BVPS-2 stainless steel material (Type 316, 304, and welds) have been collected. The stainless steel material data for both base (Type 316 and 304) and welds (both shielded metal are and submerged are) were gathered from numerous engineering sources (see Appendix B).

The stainless steel base metal information, derived from previously published data, was required for analytical purposes.

The NHIPJET program required stress-strain and J-R curve

)'

properties. These data, along with the piping certified material '

test reports (CMTR). show that the BVPS-2 BOP high energy linen j are made of high grade material, extremely resistant to unstable j tearing. The characteristics of the true stress-strain diagram and the J-R curves were used for leak rate calculations and crack stability analyses.

Appendix B also shows how the Andustry data were evaluated to derive appropriate stream-strain and J-R curve properties.

The stainless steel shielded metal are weld (SMAM) and submerged are weld (SAM) metal information was also derived from existing data, Again, the properties were required to perform leak rate

, and crack stability analyses for BOT high energy piping wolds, j Appendix D discusses the weld procedures and conditioning of i materials used for actual BVPS-2 welding, Referring to Table 3.4 the only SAM shop welds were in the RCS 14-in, line (line I.

number 084. break number 242) and in the SIS 12-in, lines (line number 287, break numbers 041 and 043: line number 288 break number 017 and line number 289 break number 010), All other shop and field welds were SMAM: the shop and field welding parameters were essentially identical with the same range of heat inputs and the same filler materials.

6.2 FERRITIC STEEL LINES Properties for the applicable BVPS-2 BOP ferritic material (i.e.,

carbon steel. SA106-B and SMAM field welds) have been gathered in Appendix C. An evaluation leading to appropriate stream-strain and J-R curve properties is presented in Appendix C for application to BVPS-2 piping.

e 6-1 L__~,.,___-.- , . ~ . . . . . _ _ ~ ~ - - _ _ - ~ - - - - - - - - - -

i The lowest temperature seen by the ferritic high energy lines is greater than 200 F during operation when any significant pressure exists. which is on the upper shelf as measured using Charpy V-notch specimens [6.1]. Originally a testing program for the applicable ferritic materials was developed due to the limited published data. However. because of the advanced state of construction of the BVPS-2 pipe rupture mitigation hardware. the testing program was determined to be not cost-effective.

Instead, the ferritic material properties employed in the analyses were obtained from recent test results provided by others as described in Appendix C.

l 6.3 HIGHEST STRESSES AND LONEST MATERIAL PROPERTIES NUREG 1651. Vol. 3 [6.2) indicates that within a particular line only .he location of highest stress corresponding to the lowest material properties needs to be analyzed for LBB. This name philosophy was expressed by the NRC staff during an August 27, 1986 meeting on this program. Therefore, the break locations in Table 3.4 have been evaluated in terms of highest stresses and lowest material propertie:t as shown in Table 6.1, Note that in some cases, more than one location in a line was chosen since the trade-off between highest stresses and lowest material properties is difficult to discern apriori without performing all of the NHIPJET frncture mechanics analyses. For the longitudinal breaks. in a pipe size in which the pressure is a constant, all results will be identical. Therefore, only three cases for longitudinal breaks are recaired: 14-in. RCS, 6-in. SIS. and 12-in. SIS.

REFERENCES 6.1 A. Zahoor et al.. Evaluation of Flaws in Ferritic Piping " fW nal Report on EPRI Research Project RP1757-l 51 Novetich Corporation and General Electric Company.

Octobe r 1985 6.2 Report of the U m S m Nuclear Regulatory Commission Piping Review Committee Evaluation of Potential for Pipe Breaks. NUREG 1061 Vol. 3 November 1984.

6-2

TABLE 6.1 HIGHEST STRESS LOCATIONS MITH LONEST MATERIAL PROPERTIES FOR CIRCUMFERENTIAL BREAKS LINE NUMBER BREAK NUMBER MATERIAL 2BDG-003-222-4 2BDG-005-O-C SMAN. SA 106B 2BDG-003-223-4 2BDG-022-O-C SMAN, SA 106B 2BDG-003-224-4 2BDG-034-O-C SMAN. SA 106B 2RCS-008-020-1 2RCS-004-C-C SMAN, TYPE 304 SS 2RCS-008-021-1 2RCS-001-C-C SMAN. TYPE 304 SS 2RCS-008-040-1 2RCS-008-C-C. SMAN. TYPE 304 SS 2RCS-008-041-1 2RCS-005-C-C SMAN. TYPE 304 SS 2RCS-008-060-1 2RCS-012-C-C SMAN. TYPE 304 SS 2RCS-008-061-1 2RCS-009-C-C SMAN. TYPE 304 SS 2RCS-014-084-1 2RCS-242-C-C SAM. TYPE 304 SS 2RCS-014-084-1 2RCS-250-C-C BASE. TYPE 304 SS 2RCS-014-084-1 2RCS-256-C-C SMAN, TYPE 304 SS 2RHS-010-024-1 2RHS-002-C-C SMAN. TYPE 316 SS 2RHS-012-001-1 2RHS-003-C-C SMAN. TYPE 316 SS 2 SIS-006-012-1 2 SIS-153-C-C SMAN, TYPE 316 SS 2 SIS-006-016-1 2 SIS-177-C-C SMAN. TYPE 316 SS 2 SIS-006-024-1 2 SIS-119-C-C SMAN. TYPE 316 SS 2 SIS-006-025-1 2 SIS-148-C-C SMAN, TYPE 316 SS 2 SIS-012-287-1 2 SIS-035-C-C SMAN, TYPE 316 SS 2 SIS-012-287-1 2 SIS-039-C-C BASE. TYPE 316 SS 2 SIS-012-287-1 2 SIS-043-C-C SAM, TYPE 316 SS l

2 SIS-012-288-1 2 SIS-016-C-C SMAN, TYPE 316 SS 2 SIS-012-288-1 2 SIS-017-C-C SAN. TYPE 316 SS 2 SIS-012-289-1 2 SIS-010-C-C SAM. TYPE 316 SS 2 SIS-012-289-1 2 SIS-012-C-C SMAN, TYPE 316 SS 0

6-3

SECTION _7 LEnK DETECTION

7.1 INTRODUCTION

One of the key parameters needed for leak-before-break (LBB) is the lowest detectable leakage for both inside and outside containment. This parameter is then multiplied by a margin for use in the LBB (NHIPJET) analyses. For breaks inside containment. the NRC staff position is that a value of 10 must be applied for the leak rate margin. This margin on leak detection is highly conservative, but meets NUREG-1061 Volume 3 (7.1 ) and the NRC broad scope rule change to GDC-4 (out f or public 3 comment). The NRC staff has indicated that a lower value for the leak rate margin may be used for outside containment lines where adequate proof for monitoring can be assured.

The following discussion focuses on BVPS-2 lear detection systems. leak detection capabilities and operator actions.

Values for the limiting detectable leakage are determined, i

A diverse number of containment parameters are monitored to detect leakage inside containment. The HHIPJET systems within the containment are RCS, RHS. and SIS. These piping systems are

all within the reactor coolant system pressure boundary as defined in Section 50.2 of 10CFR50 The BVPS-2 technical specifications define the categories of leakage to be

1. Identified Lemkage -- Identified reactor coolant system l pressure boundary leakage is leakage into closed systems, such as pump semi or valve packing leaks that are captured and conducted to a collecting tank, or leakage into the containment atmosphere from sources thnt are both specifically located and known not to interfere with the operation of leakage detection system or not to be pressure boundary leakage, or leakage of reactor coolant through a steam generator to the mecondary system.

2 Unidentified Leakage -- Unidentified reactor coolant system pressure boundary leakage is all reactor coolant system pressure boundary leakage which is not

! identified.

3 Pressure Boundary Leakage -- Pressure boundary leakage

} is leakage (except steam generator tube leakage) through a nonisolable fault in a reactor coolant system component body, pipe wall, or vessel wall.

I 7-1

The technical specifications provide limits for each of these reactor coolant system pressure boundary leakage categories.

The leak detection systems are employed to monitor leakages so that any abnormal leakage can be diagnosed well before the technical specification limits are reached. RCS inventory balances are also performed to determine the unidentified leakage from the RCS, Containment entry may be required to determine the source and the quantity of leakage.

The only NHIPJET system outside of containment is the BDG system.

Leak detection for this piping is discussed in Section 7.5 7.2 LEAK DETECTION SYSTEMS (INSIDE CONTAINMENT)

The leak detection systems employed within containment include the following:

1. Containment sump pump flow and level monitoring.

2. Containment atmosphere particulate and gaseous radioactivity monitoring, and 3 Containment pressure, temperature and humidity monitoring.

A brief description of these systems. which are provided to meet the intent of Regulatory Guide 1.45 follows:

1 Containment Sump Pump Flow and Level Monitoring -- This system employs a programmable controller which periodically measures leak rate to the containment sump i

by determining flow rate for the sump pumps and monitoring the pump operating times. This programmable controller (1) spm over normal.

alarms if there is an increase in flow of one In addition, sump flow rates are determined and logged on a shift basis by dividing the difference between the current shift and the previous shift totalizer readings, in gallons. by the elapsed time between reading. Operators in the control room also monitor sump pump running frequency and changes in the water level of the containment sump to check for any abnormal operation.

2 Containment Airborne Radiation Monitoring -- This system monitors both gas and particulate radioactivity.

Indicators and alarms are located in the main control room, i

3 Containment Pressure. Temperature and Humidity Monito[ing -- Containment pressurr temperature and humidity values are indicated and r ecorded in t'he main control room.

7-2

7.3 LEAKAGE DIAGNOSIS The leak detection systems are monitored f requently. The containment sump pump discharge is monitored on a shift basis; at least one reading every 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> is required by technical specifications. The containment atmosphere gaseous and particulate radioactivity is also monitored at least at this frequency.

Plant operator logs are reviewed for trending. If the containment sump pump flow shows a step increase of 50% between shift readings or an increase of 50% over a seven day period, an abnormal leakage procedure is followed. If the containment sump pump flow alarms indicating an increase of one (1) spm over normal. the same procedure is followed.

The abnormal leakage procedure calls for an evaluation of the containment parameters to determine as accurately as possible the source and magnitude of the leak. In addition, the procedure calls for an RCS inventory balance which is a controlled test performed to determine the magnitude of RCS unidentified leakage.

Operating experience from other plants indicates that the average long-term unidentified leakage from the RCS is between 0.1 and 0.3 gpm. Recent data for BVPS-1 indicates an RCS unidentified leakage rate of 0.32 spm inside containment.

Based upon the above discussion (plus the operator actions described in the next section), the detectable leak rate limit is 0.5 gpm. This value when multiplied by the NRC margin of 10 gives 5 gpm as the level for crack size determination under normal operating loads. Section 8 on leak rate calculations uses the 5 spm leak rate level for breaks inside containment.

7.4 ACTIONS Actions depend on the evaluation of the containment parameters.

Actions are based on Table 7.1 and the calculated value for RCS unidentified leakage.

7.4.1 Analyze / Evaluate Containment Parameters The following actions are performed whenever the conditions of Section 7.3 are exceeded:

1. Analyze information from control room instrumentation.

including a comparison of charging and letdown flow rates, to determine the nature of the leak.

2. Upon determination, notify the operating supervisor.

7-3

3. Evaluate all information available from control room instrumentation as well as locally mounted instruments to identify the leak and determit.s its location.

7.4.2 Perform An RCS Inventory Balance An RCS inventory balance is performed when the conditions described in Section 7.3 are exceeded as specified by the abnormal operating procedure. It is a controlled test to determine the magnitude of the unidentified RCS leakage. The accuracy is approximately +0.2 gpm.

It is noted that in addition to being performed when these conditions are exceeded this inventory balance is performed every 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br /> as required by the technical specifications. The following equations summarize the procedure:

Uncorrected RCS leakage rate = Leakage rate based on Volume Control Tank level change Corrected RCS leakage rate = (Uncorrected RCS leakage rate)

+ (Leakage rate correction for temperature change) i + (Leakage rate correction for pressurizer level change)

Next, the unidentified reactor coolant system leakage is determined by subtracting the leakages that are known to go to collecting tanks as follows:

Unidentified RCS leakage rate = (Corrected RCS leakage rate)

- (Leakage rate to pressurizer 4 relief tank) l' (Leakage rate to primary drains transfer tanks) 7.4.3 Containment Entry / Inspections Containment entry / inspection is necessary when unidentified RCS leakage is greater than 0.5 spm. For leakage i O.5 spm. action is decided by the Nuclear Shift Supervisor (NSS).

When containment inspection is called for. the source of leakage is determined by visual inspection and measured by a graduated cylinder to determine the leakage rate. When the source is specifically located, the leakage is then considered identified.

The following technical specification limits apply:

Pressure Boundary Leakage from a RCS Component Body. Pipe Mall, or Vessel Nall - None J

i i

7-4 4

RCS Unidentified Leakage i 1 spm RCS Identified Leakage i 10 gpm 7.4.4 Limiting Conditions for Operation Depending on the location and magnitude of the leak, appropriate repair actions will be taken or the plant will be shutdown as described below.

In accordance with technical specifications, if the calculated unidentified leakage is greater than 1 epm or identified leakage is greater than 10 gym. the leakage must be reduced to acceptable limits within 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> or be at least in HOT STANDBY within the next 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> and in COLD SHUTDOWN within the following 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />.

In accordance with technical specification. if there is any pressure boundary leakage from a RCS component body, pipe wall or vessel wall, the plant must be in at least HOT STANDBY within 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> and in COLD SHUTDOWN within the following 30 hours3.472222e-4 days <br />0.00833 hours <br />4.960317e-5 weeks <br />1.1415e-5 months <br />.

Evaluations will consider the source and magnitude of the leak, rates of change of detection variables and if shutdown is required. The evaluations will be used to determine shutdown rates and condition. Action taken will be recorded in a written log.

After shutdown, corrective action will be taken before operation is resumed.

7.5 LEAK DETECTION OUTSIDE CONTAINMENT The only system outside containment in this program is the steam generator blowdown (BDG) system. This piping can be visually inspected from the daily walkdown path. Leakage would be detected by visual inspection, and the magnitude of leakage would be determined by a graduated cylinder. EPRI has recently concluded tests to provide a better handle on what leak rates can be detected visually if a through-wall flaw develops.

The significant finding from the EPRI tests [7.2) is that visual detection is very effective for discovering leaks in high energy piping lines. BVPS-2 insulation (calcium silicate) is permeable to both steam and liquid flow, and the aluminum canister surrounding the insulation deforms as a result of the leak.

Steam and water leaks were clearly visible within minutes down to ,

the lowest flow rates tested -- approximately 0.1 gpm. Based on these tests, it is assumed that approximately 0.05 spm can be visually detected. There appears to be concensus among the NRC staff and indusfry that a factor of 10 on leak rate is  !

l 7-5

s inappropriate for outside containment lines which are readily visible during daily walkdowns. Therefore. for BVPS-2 application. a factor of 4 was chosen as the leak rate margin. A value of 0.2 spm is thus being used which includes the margin of 4 on leak rate (i.e.. 0.05 spm x 4). This level of leakage is used in Section 8 for evaluating outside containment breaks.

REFERENCES 7.1 Report of the U u S u Nuclear Regulatory Commisson Piping Review Committee: Evaluation of Potential for Pipe Breaks. NUREG-1061 Vol. 3 November 1984.

7.2 EPRI tests at Hyle Laboratories. October 9, 1986 report to be written. B. Chexal. EPRI Program Manager.

0 7-6

TABLE 7.1 EFFECT ON CONTAINMENT PARAMETERS VS. THE TYPE OF LEAK Containment RCS Secondary Component Auxiliary Parameter Leak Leak Cooling Cooling (Feedwater Water Water or Steam) Leak Leak Air Particulate Increase No No No and Radioactive Change Change Change Gas Monitors Containment Increase Increase No No Humidity / Dew- Change Change point Temper-atures Containment Increase Increase Increase Increase Sump Flow rate Containment Increase Chromates Sump Chemistry Activity in Sample Analymis RCS Hater Net No No No Inventory Loss Change Change Change

-- == --

__===_

W l

7-7 l

1

1 i.

SECTION 8 LEAK RATE CALCULATION

SUMMARY

The NHIPJET program used the EPRI PICEP (Pipe Crack Evaluation Program) [8.1) computer code to analyze pipe leakage. The PICEP code calculates the crack-opening area, the limit load critical l crack length, and flew rate for through-wall cracks in pipes.

The cracks can be oriented either axially (longitudinally) or circumferential1y. Crack opening area was calculated using elastic-plastic fracture mechanics methods similar to those used later in the NHIPJET program for determining crack stability:

leak rate calculations employ two-phase flow based on homogeneous non-equilibrium critical flow theory. The extensive validation results for FICEP are presented in Appendix E. The accuracy of PICEP is judged to be +25%.

4 Each applicable BVPS-2 break location in NHIPJET (see Table 6.1) was analyzed using tne PICEP computer program to determine leak rate as a function of crack size. The forces and moments used to calculate leak rates were combined algebraically (1): for the leak rate determinations, normal operating loads were considered to be the combination of deadweight, thermal. and pressure. All of the piping thermal load cases were examined in order to determine the dominant thermal case which represents the most likely thermal state of the piping during normal steady-state i operation. Only those loads which resulted in an opening of the crack were used (i.e., torsional moments were ignored).

The PICEP code also required a stress-strain relationship to be used for determining the crack-opening area needed to calculate leak rates. Results obtained by Combustion Engineering [8.2]

indicate that the crack-opening area in the case of circumferential welds is governed better by the stress-strain properties of the base metal. Therefore. all leak rate calculations used base metal properties for estimating crack-opening areams this was also true for the longitudinal welds since the size of a crack extends appreciably outside the small width of the weld itself. Additionally, at the direction of the l NRC staff, since lower bound stress-strain properties result in larger crack-opening areas (and subsequently higher leak rate estimates), best fit stress-strain properties for the base metal were used for conservatism (see Appendices B and C).

Results from the leak rate calculations are presented in Appendix F as a series of leak rate versus crack size plots. Based upon the limiting detectable leak rate of 0.5 gpm inside containment plus the leak rate margin of 10 (required by the NRC staf f ), a 1

8-i t

leak rate of 5 spm was used to determine a detectable leak rate (with margin) crack size for crack stability analysis.

Therefore. the leak rate results in Appendix F were analyzed to determine the crack size corresponding to 5 spm for normal operating loads, These results are tabulated in Table 8,1 for the stainless steel lines inside containment. The next sections of this report will then use these results for evaluating crack stability.

For outside containment situations (BDG lines), the same approach was taken except that a final leak rate (with margin) of O,2 spm was used rather than 5 spm. These results are tabulated in Table 8,2 REFERENCES 8,1 D. Norris et al., PICEP: Pipe Crack Evaluation Program. EPRI NP-3596-SR (Revision 1), in press, 8,2 D. M. Norris, private communication regarding unpublished Combustion Engineering work for EPRI, 8-2

TABLE 8.1 LEAK RATE RESULTS FOR STAINLESS STEEL LINES INSIDE CONTAINMENT LINE NUMBER BREAK NUMBER 5 GPM CRACK SIZE (1)

(inches)

CIRCUMFERENTIAL BREAKS 2RCS-008-020-1 2RCS-004-C-C 5.90 2RCS-008-021-1 2RCS-001-C-C 6.55 2RCS-008-040-1 2RCS-008-C-C 6.05 2RCS-008-041-1 2RCS-005-C-C 6.67 2RCS-008-060-1 2RCS-012-C-C 5.64 2RCS-008-061-1 2RCS-009-C-C 5.80 2RCS-014-084-1 2RCS-242-C-C 8.29 2RCS-014-084-1 2 RCS-2 50-C-C 6.28 2RCS-014-084-1 2RCS-256-C-C 5.30 2RHS-010-024-1 2RHS-002-C-C 8.83 2RHS-012-001-1 2RHS-003-C-C 7.82 2 SIS-006-012-1 2 SIS-153-C-C 5.17 2 SIS-006-016-1 2 SIS-177-C-C 5.33 2 SIS-006-024-1 2 SIS-119-C-C 5.33 2 SIS-006-025-1 2 SIS-148-C-C 5.63 2 SIS-012-287-1 2 SIS-035-C-C 7.15 2 SIS-012-287-1 2 SIS-039-C-C 7.45 2 SIS-012-287-1 2 SIS-043-C-C 7.10 2 SIS-012-288-1 2 SIS-016-C-C 7.85 2 SIS-012-288-1 2 SIS-017-C-C 7.75 2 SIS-012-289-1 2 SIS-010-C-C 7.02 2 SIS-012 -2 89-1 2 SIS-012-C-C 7.14 NOTE:

(1) Crack size corresponding to a 5 spm leak rate under normal j operating loada 8-3

TABLE 8.1 (continued)

LINE NUMBER BREAK NUMBER 5 GPM CRACK SIZE (1)

(inches)

LONGITUDINAL BREAKS 2RCS-014-084-1 2RCS-243-C-L 5.56 2 SIS-006-012-1 2 SIS-152-C-L 3.67 2 SIS-012-288-1 2 SIS-018-C-L 5.08

=- -_

NOTE:

(1) Crack mize ccrresponding to a 5 spm leak rate under normal operating loada l

0 4

8-4

TABLE 8,2 LEAK RATE RESULTS FOR FERRITIC STEEL LINES OUTSIDE CONTAINMENT LINE NUMBER BREAK NUMBER 0.2 GPM CRACK SIZE (1)

.. (inchen)

CIRCUMFERENTIAL BREAKS 2BDG-003-222-4 2BDG-005-O-C 1.97 2BDG-003-223-4 2BDG-022-O-C 1,76 2BDG-003-224-4 2BDG-034-O-C 2,16 NOTE:

(1) Crack size corresponding to a 0,2 spm leak rate under normal operating loada D

I 8-5

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

SECTION 1 I CRACK STABILITY CALCULATIONS IFOR NORMAL + SSE LOADS) i The methodology to evaluate the stability of through-wall cracks requires knowledge of the applied loads. a leak rate crack size.

and the material properties. The piping axial loads and bending i moments were obtained f rom the NUPIPE stream calculations. The j dead weight, thermal and pressure loads and moments were j combined algebraically to define normal operating conditions (exactly the same as for the leak rate calculations in Section 8). These loads and moments were then combined absolutely with the SSE inertia loads and moments to define the appropriate conditions for ammessing crack stability.

The appropriate leak rate crack sizes for both inside and outside containment were developed in Section 7 For inside containment

piping, a leak rate crack mize for 5 epm (which includem a l multiplicative margin of 10) was used as required by the NRC 1 staff. For piping located outside containment, in a region where 1 visual inspection is possible during routine walkdown, the leak rate crack size was chosen for 0.2 gym (including a j multiplicative margin of 4). These crack sizes were evaluated f or stability by applying a further margin on crack size of at

! least 1.8 as required by the NRC staff.

The material properties used for crack stability analyses were all lower bound as shown in Appendices B and C. The NRC staff defined two sets of calculations to be performed for circumferential weldas first. lower bound weld metal stress-strain and J-R curve properties were used, and second. Iower bound base metal stream-strain and J-R curve properties were j evaluated and compared to the weld metal results. The more j conservative of the two sets of calculations was used to assess

! stability.

j For the circumferential break locations, crack stability was

assessed using the EPRI computer code FLET (Flaw Evaluation for Tearing Instability). The FLET code is based on elastic-plastic

fracture mechanica, and the verification results for this code are discussed in Appendix G. The deformation plasticity failure assessment diagram (DPFAD) method for evaluating crack stability was chosen FLET utilizes up to four separate analytical j mpproaches including the two most common. DPFAD and J-T.

Required input into FLET includes the material stress-strain l properties (in terms of Ramberg-Osgood curve fit parameters) and j the experimental fracture resistance curve (J-R). In the DPFAD j mpproach. the quantity K = (J / J) a p1 tted versus S *U# L r r I

I a

9-1 i

{-_ - - . - . .- -

to define the DPFAD failure curve using the material stress-strain and J-R values, J' is the elastic J. J is the total J. o is the applied stress, and U is p the limit load stress, Next.

l

assessment points using applied J values are plotted on the same K r and S r curve to determine whether initiation or instability has occurred. Instability assessment points lying inside the DPFAD curve are safe, and those lying outside the curve indicate an unsafe. unstable situation for the particular size crack and applied l oa ds . Figure 9,1 illustrates the DPFAD curve and assessment points.

The circumferential crack results are listed in Tables 9,1 and i 9,2 for the inside and outside containment lines, respectively, The columns in the tables identified as " margin on crack size" correspond to the minimum value of 1.8, When the DPFAD margin (for normal + SSE loads) is greater than or equal to unity. the

crack is stable. Also shown in this table are the results from limit load calculations using the ultimate tensile stress f or net section plastic collapse [ corresponding to note (1)]

these results when compared to the DPFAD results suggest that limit '

load conditions may be more limiting than elastic-plastic fracture instability for the 6- and 8-in, lines. However, it must also be noted that the FLET DPFAD calculations for stability in the 6-in, SIS lines involve crack sizes which extend slightly beyond 180 degrees around the pipe, which is in the extrapolation i regime for FLET applicability, It is felt, due to the conservative use of lower bound material properties and the fact that there is only a small discrepancy between the limit load and i DPFAD crack stability analyses, that the two 6-in, SIS lines

! (line numbers 024 and 025) pass the intent of the margin on crack size, ,

For longitudinal cracks. only pressure loads are important in assessing crack opening area and stability. As shown in Table 9,3 the margin on crack size was evaluated by ratioins the critical crack size calculated using the empirical results of Eiber (9.1 ) to the 5 epm leak rate crack size. Note that these j resultant margin numbers are significantly greater than the limit of 1,8, If a stability analysis had been performed these numbers would only decrease slightly, especially since failure would occur in base metal for these long cracks which extend well outside any welds, REFERENCES

! " Investigation of the Initiation 9,1 R. J. Eiber et al,.

and Extent of Ductile Pipe Rupture." Battelle Columbus Laboratories Report BMI-1908 June 1971, i

b I

I i

, 9-2

)

i

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

TABLE 9,1 CIRCUMFERENTIAL CRACK STABILITY EVALUATION FOR INSIDE CONTAINMENT MARGIN ON DPFAD MARGIN FOR LINE NUMBER BREAK CRACK SIZE NORMAL + SSE NUMBER --- - - - -=- -

(1) (2) BASE HELD 2RCS-008-020-1 004 1,98 1,00 1,27 >1,44 2RCS-008-021-1 001 1,92 2.00 1,29 1,40 2RCS-008-040-1 008 1,95 1,80 1,24 >1,39 2RCS-008-041-1 005 1,89 1,80 >1,48 >1,41 2RCS-008-060-1 012 1.97 1,80 1,24 >1,44 2RCS-008-061-1 009 2,02 2,00 1,04 1,54 2RCS-014-084-1 242 2,17 2,00 1.20 1,47 2RCS-014-084-1 250 2,47 2,00 1,40 (3) 2RCS-014-084-1 256 2.38 2,00 1.20 1,47 2RHS-010-024-1 002 1,92 2,00 1,02 1,55 2RHS-012-001-1 003 2,03 2,00 1,41 2,02 j 2 SIS-006-012-1 153 1,81 1,80 1,07 1,58 2 SIS-006-016-1 177 1,80 1,80 1,10 1,58 2 SIS-006-024-1 119 1.75 1,80 1,00 1,46 2 SIS-006-025-1 148 1,70 1,80 >1.06 1,39 2 SIS-012-287-1 035 2,54 2,00 1,72 2,47

2 SIS-012-287-1 039 2,44 2,00 1,61 (3) 2 SIS-012-287-1 043 2,44 2,00 1,56 1.97 2 SIS-012-288-1 016 2,28 2,00 1,36 1,95 2 SIS-012-288-1 017 2.30 2,00 1,40 1,81 2 SIS-012-289-1 010 2,54 2,00 1,71 2,17 2 SIS-012-289-1 012 2,52 2,00 1,69 2.43 NOTES

(1) Ratio of the limit load critical crack size (for Normal + SSE loads) to the 5 gpm crack mize (Normal loads only),

(2) Ratio of DPFAD stability crack size analysis (for Normal +

SSE loads) to the 5 upm crack size (Normal loads only):

minimum value in 1.8. corresponding to DPFAD margin {1,00 (3) Base metal break location. no weld metal analysis required 9-3

TABLE 9.2 a

1

(

CIRCUMFERENTIAL CRACK STABILITY EVALUATION FOR OUTSIDE CONTAINMENT MARGIN ON DPFAD MARGIN FOR i

LINE NUMBER BREAK CRACK SIZE NORMAL + SSE l NUMBER ---------------------------- -------

j (1) (2) BASE HEkR d

2BDG-003-222-4 005 3.01 1.80 1,53 1.78 2BDG-003-223-4' 022 3.28 1.80 1.30 1.60 2BDG-003-224-4 034 2.88 1.80 1.63 1.83 d

NOTES:

2 (1) Ratio of the limit load critical crack size (for Normal + SSE loads) to the 0.2 gpm crack size (Normal loads only).

(2) Ratio of DPFAD stability crack size analysis (for Normal +

SSE loads) to the 0.2 spm crack size (Normal loads only);

i minimum value is 1.8 corresponding to DPFAD margin 11.00.

9-4 i

I

TABLE 9.3 LONGITUDINAL CRACK STABILITY EVALUATION FOR INSIDE CONTAINMENT LINE PIPE SIZE RATIO OF CRITICAL SIZE (INCHES) TO 5 GPM LEAK RATE SIZE (1) i RCS 14 4,o4 SIS 6 3.57 SIS 12 4.34 I

i NOTES:

(1) Crit.ical size is based upon empirical failure results (M) for pressure loads only

.?

1 9-5

. - - - - ---,w , ..w ,n_w..,-- n,g---, ,- - ,nw- w,,nny-,,-,-, ,, -,,---ym-, ,-,-,---.-.------------v_ , -n---

10 -

P

,i Unstable

/. e Region 0.8 -

! e N '.

0.6 -

• table Region e b

0.4 -

• DPFAD Margin =

(b/a)

Assessment Points a/ .

1r d i I I I I I ? I 1 0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 s,

t Figure 9.1 Failure Assessnent Diagran in Tems of Stable Crack Growth D

9-6

SECTION 10 CRACK STABILITY UNDER EXCESSIVELY HIGH LOADS Crack stability for excessively high loads was the final fracture analysis effort performed on the high energy piping systems.

Assuming a detectable, but stable. leakage size crack has occurred in the pipe, it was shown that. if an SSE occurs prior to detection of the leak, this through-wall crack was stable under loads that are significantly greater than SSE. The excessively high loads that were applied are 1,414 x (normal +

SSE),

The stability analysis followed the procedures described in Section 9 except that no extra margin was placed upon the leakage size crack. The required margin was applied to the loads.

Results for the circumferential breaks are listed in Tables 10.1 anc. 10,2, Since all of the DPFAD margins are significantly greater than 1,00 all of the high stress break locations meet the requirements for stability under high loads.

For the longitudinal breaks. the approach used in Section 9 was again used except that stability was based upon loads equal to 1,414 x pressure. The results show a significant margin on crack size above 1,00 for these high loads, and therefore. stability is verified, a

6 4

10-1

TABLE 10.1 EXCESSIVE LOAD CIRCUMFERENTIAL CRACK STABILITY FOR INSIDE CONTAINMENT LINE NUMBER BREAK NUMBER DPFAD MARGIN FOR 1,414 (NORMAL +SSE]

BASE NELD 2RCS-008-020-1 2RCS-004-C-C 2,02 3,07 2RCS-008-021-1 2RCS-001-C-C 2.34 3,56 2RCS-008-040-1 2RCS-008-C-C 2,05 3.12 2RCS-008-041-1 2RCS-005-C-C 2.38 3,61 2RCS-008-060-1 2RCS-012-C-C 1.87 2.85 2RCS-008-061-1 2RCS-009-C-C 2,08 3,18 2RCS-014-084-1 2RCS-242-C-C 1.94 2.42 2RCS-014-084-1 2RCS-250-C-C 1,71 (1) 2RCS-014-084-1 2RCS-256-C-C 1,34 2.03 2RHS-010-024-1 2RHS-002-C-C 3,25 4.76 2RHS-012-001-1 2RHS-003-C-C 2.36 3,55 2 SIS-006-012-1 2 SIS-153-C-C 2,09 3.20 2 SIS-006-016-1 2 SIS-177-C-C 2.21 3.38 2 SIS-006-024-1 2 SIS-119-C-C 2,06 3,14 2 SIS-006-025-1 2 SIS-148-C-C 2,16 3.28 2 SIS-012-287-1 2 SIS-035-C-C 2,54 3.85 2 SIS-012-287-1 2 SIS-039-C-C 2.50 (1) 2 SIS-012-287-1 2 SIS-043-C-C 2.33 2.97 2 SIS-012-288-1 2 SIS-016-C-C 2.30 3,46 2 SIS-012-288-1 2 SIS-017-C-C 2.32 2,97 2 SIS-012-289-1 2 SIS-010-C-C 2.50 3,21 2 SIS-012-289-1 2 SIS-012-C-C 2.50 3,78 tlOTES :

(1) Base metal.. break location, no weld metal analysis required 10-2

TABLE 10.2 EXCESSIVE LOAD CIRCUMFERENTIAL CRACK STABILITY FOR OUTSIDE CONTAINMENT LINE NUMBER BREAK NUMBER DPFAD MARGIN FOR 1.414 (NORMAL +SSE)

BASE WELD t

2BDG-003-222-4 2BDG-005-0-C 1.52 1.98 2BDG-003-223-4 2BDG-022-0-C 1.25 1.68 2BDG-003-224-4 2BDG-034-0-C 1.68 2.15

/

9 8

10-3

TABLE 10,3 EXCESSIVE LOAD LONGITUDINAL CRACK STABILITY EVALUATION FOR INSIDE CONTAINMENT LINE PIPE SIZE RATIO OF CRITICAL SIZE (INCHES) TO 5 GPM LEAK RATE SIZE (1)

RCS 14 2.79 SIS 6 2.49 SIS 12 3.01 1

d NOTES:

(1) Critical size is based upon empirical failure results (M) for (1.414 X preneure) loada 9

10-4 i

APPENDIX 3, NUREG-0582 FLUID TRANSIENTS 9

APPENDIX A NUREG-0582 FLUID TRANSIENTS This appendix is a list of fluid transients compiled from a review of NUREG-0582 as applied to BVPS-2 The tabulation of transients was extracted directly from project design documentation and includes systems beyond the HHIPJET scope.

Also, when appropriate, references to operation and maintenance procedures (OM - # ) are provided as indications of operational controls intended to mitigate certain fluid transient postulations.

O A-1 l

l d

i BVPS-2 FLOW TRANSIENT APPLICABILITY Title / Description Applicability l

A. Pump Start with Inadvertently Voided Discharge

, Lines A1. RHS Not Applicable - To establish a controlled cooldown the system is vented and temperature and pressure i between RCS and RHS must be within certain criteria before RHS operation (See Operating Manual OM-2.10.4.A)

A2 HHSI NA - Discharge lines are normally filled and operating as the chemical and volume control system. System is filled and vented in accordance

} with OM-2.7.4.W.

4 j A3. LHSI NA - System is normally filled and vented in accordance with OM-2,11.4.H.RWST minimum water level is above the elevation of the cold leg

) injection nozzles, keeping lines full of water.

j A4. SWS Applicable-Enveloped by transient of restart of l pumps. (See Section F, Transient No. F2). Vacuum i

breakers were installed to reduce water hammer from restart of pumps. Voids would form as a result of pump stop and SWS water drain back to ultimate heat sink. Before normal start of a service water pump

] (See Operating Manual OH-2.30.4.A), the system is

,. filled and vented in accordance with OM-2.30.4.F.

AS. CCP NA - The CCP System is filled and vented in accordance with OM-2.15.4.L. RHS coolers and RHS pump seal water coolers are the only two coolers i

not normally in operation. The portion of the i

CCP system to the RHS coolers is kept full since only the inlet motor operated isolation valve is closed during normal operation.

i A6. FWE NA - Voids are not expected due to filling and venting procedures after ISI tests.

l A7. RSS Applicable - RSS discharge lines to spray rings are normally empty. Enveloped by system fill of empty lines (see Section B, Transient No. B6). Analyzed in EMD Fluid Transient Cale. No. 12241 - NP(B)-

316-FA and -260-FIA.

I 1

j A-2

)

i

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

Title / Description Applicability A8. QSS Applicable - QSS discharge lines to spray rings are normally empty. Enveloped by system fill of empty lines (see Se-tion B, Transient No. B7). Analyzed in EMD Fluid transient Calc. No. 12241-NP(B) -

273-FA and -319-FA.

A9.RCS Not Applicable - System is filled and vented in accordance with OM-2.6.4.E or a steam bubble is forced in the pressurizer before a Reactor Coolant Pump is started in accordance with0M.2.6.4.A.

B. Expected Flow Discharge into Initially Empty Lines Bl. RCS Safety and Applicable - Analyzed in EMD Fluid Transient Calc.

Relief Valve Nos. 12241-NP(B)-328-FA and -517-FA Discharge B2. RHS Suction Applicable - Analyzed in EMD Fluid Transient Calc.

Relief Valve No. 12241-NP(B)-312-FA and -298-FA.

Discharge (2RHS*RV721A,B)

, B3. HHSI NA - Lines normally filled and operating as CVCS.

B4. LHSI - SIS NA - Lines normally filled with water from the RWST Lines from Voids are not expected due to filling and venting RWST to RCS procedures. (See OH-2.11.4.H).

Cold and Hot Legs B5. LHSI - Cross Applicable - Analyzed in EMD Fluid Transient Calc.

over from RSS No. 12241-NP(B)-287-FA for Longterm Safety In

, jection from the Containment 4

Sep i B6. RSS - Normal Applicable - Analyzed in EMD Fluid Transient Calc.

1 Pump Start in Nos. 12241-NP(B)-316-FA and -260-FIA to Empty Lines and Full Flow Pump Test B7. QSS Normal Applicable - Analyzed in EMD Fluid Transient Calc.

Pump Start Nos. 12241-NP(B)-273-FA and 319-FA.

into Empty Lines .

A-3

- - - - . +- - r-.--- r--. -- - ~ . - _ _ . , -_ - --m.. - - - - . - - - - - - . - - -

Title / Description Applicability B8. MSS Safety Applicable - Analyzed in EMD Fluid Transient Calc.

Relief Valves Nos. 12241-NP(B)-464-FD and -278-FC.

(2 MSS *SV101-105A, B, C) and Atnospheric Dump Valves (2SVS*PCV101A,B,C)

Opening B9. MSS Steam Sup Applicable - Initially empty lines at ambient ply to Aux. temperature are filled with main steam.

Feedwater Pump Condensate pots remove any condensate formed.

Turbine Analyzed ia EMD Fluid Transient Calc. No.

12241-NP(B)-452-FD.

2 B10. CCP Flow into NA - All coolers supplied by CCP system are Empty Coolers normally filled except when down for maintenance.

Venting procedures during gradual fill eliminate the possibility of this transient occurring.

(See OM-2.15.4.L.)

Bl.l. SWS Flow into Applicable - When the RSS system starts af ter an Initially Em- accident, service water valves isolating the RSS pty RSS cooler are opened filling the empty lines and Coolers coolers. Analyzed in EMD Fluid Transient Cale No.

12241-NP(B)-381-FA.

B12. MSS Radiation Applicable - Water slug between monitor and outlet Monitor isolation valve passes through RO into the Operation discharge piping when valve opens.

B13. FWE Opening of Applicable - Relief valve opens filling empty RV101 on the line (relieves excess pressure due to overspeed Turbine Pump of the turbine). Analyzed in EMD Fluid Transient Discharge Calc No. 12241-NP(B)-276-FA B14. ASS Flow into Not Applicable - OH-2.27.4.A for the ASS startup Empty Lines includes provisions and cautions to prevent water /

steam hammer or thermal shock.

C. Valve Opening, Closing, and Instability C1. RHR - Closing / NA - Suction and discharge motor operated Opening of isolation valves have an opening / closing time Suction or of 120 sec. No water hammer will occur due to Discharge valve closing or opening. Also, since the HCVs fail Isolation open and a recire loop is provided around the pumps 4

Valves.' a flow path for the pumps would always exist after failure of any one of the control valves. No water hammer would occur.

A-4

1 Title / Description Applicability C2. HHSI - Closing / NA - Discharge motor-operated isolation valve has Opening of Dis- an opening / closing time of 10 sec. The BV-2 NHSI charge Isola-System operates similarly to the BV-1 NHSI System tion Valves which has not experienced any water hammer events.

Also, a flow path would always exist after a failure of the control valves. No water hammer should occur.

C3. LHSI - Opening / NA Discharge motor-Closing of Dis- operated isolation valves charge Isola- have an opening / closing tion Valves time of 12.2 sec. No water hammer should occur.

C4. FWS - Closure Applicable - Analyzed in EMD Fluid Transient Calc.

of Flow Con- No. 12241-NP(B)-258-FA. ,

trol Valves C5. FWS - Instability Not Applicable - BV-2 Flow Control Valves were of Flow Control modified to eliminate flow instability observed at Valves BV-1. Modifications to increase compatibility with Feedpumps included a reduction in trim size and the replacement of plug-type trim with cylindrical

trim.

C6. FWE Closure Not Applicable - Modulation of control valves of Flow Con- causing pressure waves not expected due to rol Valves reaction time for hand control valves. Control

)

valve closing not considered to produced sign-ificant dynamic loads.

C7. MSS Steam Sup- Not Applicable - The steam hammer event of 4

ply to Aux. closing the isolation valves is analyzed in EMD Feedwater Pump Fluid Transient Calc. No. 12241-NP(B)-452-FD.

Turbine Isola-tion Valve i Closure C8. MSS Closure of Applicable - Analyzed in EHD Fluid Transient Calc Turbine Stop No. 12241-NP(B)-274-FB.

Valves C9. MSS Opening of Applicable - Analyzed in EMD Fluid Transient Cale Turbine Bypass No. 12241-NP(B)-142-FIB.

Valve C10. MSS Closure of NA - MSIVs close in 4 seconds. Loads developed MSIVs would be enveloped by those from the closing of the turbine stop valves (150 milliseconds). '

A-5 1

e

.----,----en--,--r.- w,.-~ , - - - . . - - , . . . - - , -n . - - - v,-, n---. - . - - - . . , . , . . . . _ . - - , , . - - . . . ~ - . . - - . . - - - - , - - , - - - - - -

Title / Description Applicability Cll. RSS Isolation NA - Discharge motor-operated isolation valves Valve Closure have an opening / closing time of 60 sec.

(2RSS*MOV156A-D)

C12. QSS Isolation NA - Discharge motor-operated isolation valves Valve Closure have an opening / closing time of 60 sec.

(2QSS*MOV101A,B)

C13. ASS Closure of Not Applicable - Air operated isolation valves have Air-Operated an opening / closing time of 15 sec.

Isolation Valves C14. BDG Closure of Not Applicable - 3" Air-operated isolation valves

, Air-Operated have an opening / closing time of 10 sec.

Isolation Valves D. Check Valve Closing and Delayed Opening DI. RSS Check NA - If only 2 of 4 RSS pumps start, RSS water will Valve closure also fill the risers up to the spray headers in During Minimum the safety trains that are not operating. No water Safeguards slug is expected due to the displacement of air in Operation the riser. This transient would be, enveloped by the startup transient in Section B, No. B6.

D2. QSS Check NA - Same as above for RSS except that only 1 of Valve Closure 2-QSS pump starts.

During Minimum Safeguards Operation D3. FWS Main Feed Applicable - Check valve must maintain its water Check structural integrity only. Analyzed in EMD Fluid Valve Slam Fluid Transient Calc No. 12241 -NP(B)-293-FA.

D4. MSS Check Applicable - Check valve closure due to postulated Valve Closure MSS line break. This transient is enveloped by in the Steam stea,a fill of the empty line (Section B, Transient Supply Lines No. B9) to the Aux.

Feed Pump Turbine DS. RHS . Check Not Applicable - Pump discharge paths are Valve Closing separated and redundant. Check va.lves are located and Delayed -

in horizontal sections of piping.

Opening i

A-6 i

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

i I

l Title / Description Applicability D6. HHSI Not Applicable - The BV-2 HHSI System operates similarly to the BV-1 HHSI System which has not experienced any water hammer events.

D7. LHSI Not Applicable - Same as HHSI E. Water Entrainment in Steam Lines El. MSS Main Steam NA - Entrained water in main steam line is limited Lines by the collection and draining, via the Steam Drains System (SDS), of any condensate formed in the lines.

E2. MSS Water En- NA - Same as above for main steam lines trainment in Turbine Bypass Piping E3. MSS Water En- NA - Same as above for main steam lines.

trainment in Steam Lines to the Aux. Feed-water Pump Turbine E4. ASS Stemalines Not Applicable - Entrained water in auxiliary steamline is limited since condensate is collected in main steamlines (aux. steam source) and steam traps are located at various points in the system to remove condesate.

F. Transient Cavita-tion (Column Separation)

Fl. CCP Water NA - BVPS-2 has a surge tank which keeps the system Column Separa- filled (no drain back after pump stop.)

tion Effects Following Pump Stopping F2. SWS Water Applicable - Analyzed in EMD Fluid Transient Calc Column Separa- No. 12241-NP(B)-173. Vacuum breakers were sized tion Effects at and located to elimiate water hammer for LOOP.

Pump Discharge Following Pump Stop.and Restart 8

A-7

Title / Description Applicability F3. SWS Water Column Not Applicable - Check valves added to eliminate Separatica Effects water column separation.

at Main Steam Valve House Cool-ing Coils Follow-ing Pump Stop and Restart F4. RSS Water NA - Pumps have been tested with inadequate NPSH Colum Separa- for 13 minutes without failure. The RSS pump tion Effects and safety trains are separated and redundant.

Following Pump A flow transient analysis for column separation Stopping or due to drain back to cont. sump after pump stop and Inadequate subsequent restart is not required since a single NPSH failure (pump stop) in one train would not affect the others.

G. Steam Bubble Col-lapse Due to Rapid Condensation G1. HHSI Collapse NA Cold water injection into the RCS from safety of Steam Bub- injection could cause rapid condensation of steam bles Formed as with resulting water hammer back through safety a Result of injection lines. Loads are not considered large Local or enough to be of concern (pg. A-21, NUREG-0582).

System Also, the BV-2 HHSI System operates similarly to Depressurization the BV-1 HHSI System which has not experienced any water hammer events.

G2. LHSI Collapse NA Same as HHSI above.

of Steam Bub-bles due to Local or System Depressuration G3. Safety Accum- NA Same as HHSI above. Safety Injection Startup ulator Collapse Test Procedure (OM-2.11.4.A) details requirements of Steam Bubble to preclude steam pocket formation during test.

Due to Local or Also, see OST 2.11.15 " Safety Injection Accumulator System check Valve Test."

Depressuration G4. RHR Collapse NA - Steam bubble formation of Steam between check and closed Bubble isolation valves due to higher temperature in the RCS avoided by startup procedures. See OM-2.10.4.A.

l l A-8

Title / Description Applicability G5. FWS Slug In- NA - Unlikely due to inverted 'J' tube design of pact Due to feedring in steam generators on BVPS-2.

Rapid Conden-sation in Steam Generator G6. BDG - Collapse Not Applicable - Steam bubble formation is not of Steam Bubble expected. Feedwater is injected into the steam Generator blowdown water downstream of the containment isolation valves to subcool and prevent slugging or flashing in the blowdown lines.

G7. CCP Steam Bub- NA - Unlikely since system and components are ble Formation filled and vented before being placed into and Collapse operation. (OM-2.15.4.L) in Heat Exchangers G8. SWS Steam NA - Same as CCP above. (OM-2.30.4.F)

Bubble Forma-tion and Col-lapse in Heat Exchangers H. Pump Start and Postulated Seizure with Full Lines Hl. HHS1 NA - Pump start /Stop negligible due to adequate venting and priming procedures before start See OM-2.7.4.W. Pump seizure is not analyzed as a transient mode on BVPS-2 since motor overload protection is provided which should trip the motor as rotational resistance increases prior to complete seizure. Also rotational momentum of the pump and motor will prevent an instantaneous pressure spike as a result of pump seizure.

50. LHSI NA - Pump start /stop should not produce large dynamic loads due to adequate venting and priming procedures before start. (See OM-2.11.4.H)

H3. RHS NA - Same as HHSI above (See OM-2.10.4.A).

H4. RSS NA - Pump start /stop are analyzed as part of transient no. B6 (pump start and fill of empty

,. lines). Postulated pump seizure is considered a single failure on BVPS-2. Since the RSS pump

. suction and discharge lines are separated and redundant, no flow transient analysis is required.

H5. QSS NA - Same as RSS above.

i A-9

Title / Description Applicability H6. CCP NA - Same as HHSI above (See OM 2.15.4.L)

H7. SWS NA - Same as HHSI above (See OM-2.30.4.F)

H8. FWE NA - Same as HMSI above H9. RCS NA - Pump start /stop should not produce significant loads due to system design and startup procedure (OM-2.6.4.A). Reactor Coolant Pump seisure is considered an accident condition. BV-2 is designed to mitigate its effects.

i o

i a

~

A-10

1 APPENDIX _B_

STAINLESS STEEL MATERIAL PROPERTIES (This appendix is under a separate cover and treated as proprietary) a f

i i

u a

i e

l B-1 i-._ .. . -_ --- --, - - . - _ . _ - - . -

APPENDIX p_

FERRITIC STEEL MATERIAL PROPERTIES (This appendix will be under a separate cover and treated as proprietary) l l

I e

i C-1

6 APPENDIX p_

HELDING PROCEDURE i

APPENDIX D HELDING PROCEDURE D.1 INTRODUCTION The terminal end butt welds considered under the BVPS-2 HHIPJET program are welded in the field using the Gas Tungsten Arc Helding (GTAN), and Shielded Metal Arc Helding (SMAH). processes.

Helds at intermediate points in the piping systems are made in either an off-site pipe fabrication shop or in the field at BVPS-

2. The discussion that f ollows is based on approved field welding procedures.

D.2 FRACTURE TOUGHNESS Helding variables which may influence fracture toughness of weld metal and weld heat affected zones are specified in ASME Boiler and Pressure Vessel Code,Section IX. The control of these variables within ranges proven by impact testing is to ensure that the weld and heat affected zone will behave in a non-brittle manner at the lowest service temperature required for system operation. The influence of these variables on the maximum obtainable toughness, that is, the upper shelf energy levels, is not determined by ASME Section IX testing. This latter subject has been reviewed, and as reported in EPRI Publication Number NP-122 Table 5.3. for carbon steel, was concluded that the upper shelf energy level is not appreciably affected by weld heat input [D.1 ] . The heat input range of SMAN welds studied.in this report was 15 to 50 KJ/in. For the Submerged Arc Helding (SAH) process, the heat input range studied was 70 to 100 KJ/in. A small decrease in upper shelf energy level was noted for one of the SAH weld flux-filler metal combinations at high heat inputs.

The absorbed energy decreased, hqwever, by less than twenty percent of the highest level observed as shown in Figure 5.38.

The welding practices used at BVFS-2 were not required to be impact tested in accordance with'ASME Section IX. However, due to the size of the welding electrodes used and their typical electrical characteristics along with weld weave width control of three to five times the electrode diameter, heat input ranges were determined. l Based on these parameters, the heat inputs used for carbon steel welding ranged from 20 to 80 KJ/in. The acceptability of this heat input range is also supported by the impact test results shown on the CMTR(s) for each heat of carbon steel weld filler metal used in the field at BVPS-2. This heat input range has also been shown to be acceptable based on the evaluation reported in Reference D.2. The heat inputs used for stainless steel range from 15 to 50 KJ/in and were established by corrosion testing of weld test coupons to ensure that excessive

. sensitization of'the base metal would be avoided. For SA106-B l material, the maximum expected upper shelf temperature occurs at 200 F [D.?_). .

9 I

D-1

D.3 BASE METAL PREPARATION The ends of the pipe, valves, or fittings to be joined by welding are prepared for welding by machining, flame cutting (except stainless steel), grinding, sawing, plasma arc cutting, or shearing. The dimensions and configurations are in accordance with the standard drawings of Stone and Webster's Specification 2BVS-920.

The method used to prepare the base metal leaves the weld preparation with reasonably smooth surfaces. The surfaces for welding are free of scale, rust, oil, grease, and other harmful foreign material. Helding is not performed on wet surfaces. If welding is not started immediately after cleaning, the weld joint is suitably wrapped to prevent contamination.

End prep configuration for standard butt welds, as in the WHIPJET lines, uses a J-bevel. Under certain circumstances, end prepping consists of a V-bevel.

If piping component ends are counterbored, such counterboring does not result in a finished wall thickness after welding less than the minimum design thickness. Hhere necessary, weld metal of the appropriate analysis is deposited on the piping component to correct minimum wall or, in the case of attachment welds to achieve a suitable fit-up.

D.4 MATERIAL ALIGNMENT The forces applied during alignment are limited to amounts that will not deform the piping or components, weld joint, or end prep. Parts that are joined by welding are fitted, aligned, and retained in position during the welding operation by using bars, jacks, clamps, tack welds, temporary attachments, or mechanical alignment clamps.

l l

l Localized heating of carbon steel piping is permitted provided the temperature, verified with Tempilsticks, does not exceed 500 F and the process is witnessed by the Helding Supervisor.

Localized heating of stainless steel piping is permitted provided the temperature does not exceed 350 F. Hhere 500 F (for carbon steel) or 350 F (for stainless steel) is not enough heat to achieve desired results, permission is requested from Engineering to " warm" bend the pipe using approved procedures and stress relieving equipment. Heating of stainless steel above 800 F, the lower limit for sensitization, is not permitted. Also, heating of carbon steel above 1300 F. or into the lower critical temperature range, is not permitted.

j ..

e D-2 l

i _

D.5 FIT-UP Fit-up tolerances are in accordance with the standard sketches referenced on the technique sheet, and the following information as it applies for ASME, Class 1, 2, and 3 and ANSI B31.1. for Class 4 systems. The acceptable misniatch for ASME III piping is 1/32" per side and for ANSI B31.1 piping is 1/16" per side.

D.6 PREHEATING Preheating of carbon steel is performed in accordance with the requirements of ASME Section III and the field piping specification 2BVS-920, as follows:

All welds in materials greater than 3/4" thick require a minimum preheat and minimum interpass temperature of 200 F.

All welds in materials 3/4" thick and less require minimum preheat of 50 F.

Preheating of stainless steel is as follows:

Preheating is not required when the base metal temperature is above 60 F.

When the base metal temperature of the materials is below 60 F. preheating is performed by uniformly heating circumferential1y to a temperature of 60 F minimum (warm to the touch).

D.7 HELDING GASSES The shielding and purging gas for the GTAH welding process is welding grade Argon. Gas flow rates are within the range specified on the applicable technique sheet. Shielding gas is used to f orm an ionized gas (plasma) for metal arc transfer, to shield the molten base / weld metal puddle and to cool the tungsten electrode.

D.8 HELD FILLER METAL All weld material is purchased in accordance with ASME Section III requirements and the Field Piping Specification. Impact testing is required for all carbon steel weld filler metal with the exception of inserts. Delta ferrite determination is required in accordance with Regulatory Guide 1.31 for stainless steel weld filler metal.

9 D-3

D.9 HELDING TECHNIQUE The filler metal sizes shown on the technique sheets are the only sizes used. Typical bead sequence and number of layers shown on the individual welding procedure technique sheets are for illustrative purposes only. The welding technique sheets are applicable to all position welding unless otherwise stated. All vertical welds using the GTAM and SMAN process are performed in the upward direction.

Each cead deposited is thoroughly cleaned of all oxidation, slag or flux using a descaling tool and/or wire brush. Crater cracks, porosity and undercutting are removed by grinding before depositing the next bead of welding.

D.10 INTERPASS TEMPERATURE The interpass temperature, specified in the welding procedure technique sheet, is checked on the piping adjacent (1" max.) to the welding groove, using "Tempilsticks", a surface pyrometer, or an approved equal. After a bead of welding has been deposited around the complete circumference of the joint, it is cool-d to below the maximum interpass temperature before starting the next bead of welding. Generally, by the time the deposited bead is thoroughly cleaned and inspected, the temperature will be below the interpass temperature.

To control the interpass temperature of stainless steel welds, ,

water quenching with clean, lint-free rags soaked with demineralized water may be used to bring the weld below 350 F before continuing to weld.

D.11 NELD FINISHING All weld finishing must be completed before NDT tests are performed. Local grinding with appropriate wheels for the type of base materials involved are used where necessary to achieve the desired surface finish. The preparation of welds for ISI is in accordance with 2BVS-920. Held edges are to merge smoothly with the base metal.

D.12 STRESS RELIEVING Stress relieving, when required, does not have to be performed immediately after the completion of the weld unless specified on the welding procedure technique sheet. The completed weld cools slowly to ambient temperature and the stress relieving may be performed at a later time.

~

Joints preheated may be wrapped with suitable materia 1 to ensure slow cooling. Stress relieving is performed after acceptable radiographic inspection (when' required) has been performed.

D-4

The stress relieving operation consists of uniformly heating, to the temperature specified on the welding procedure technique sheet, a circumferential band around the pipe. The width of the heated band on each side of the greatest width of the finished weld is not less than two times the weld metal thickness or two inches, whichever is less. This temperature is maintained for a period of time based on one hour per inch wall thickness, but in no case less than one-half hour. Upon completion of the proper heating cycle, the weld and heated area is allowed to cool slowly to ambient temperature.

The heating and cooling rate above 600 F does not exceed 400 F per hour divided by the maximum metal thickness in inches, and in no case does the rate exceed 400 F per hour. Rates of heating and cooling are not less than 100 F per hour. Hhen the stress relieving operation is being performed adjacent to a valve, the valve is approximately one-half open.

Post weld heat treatment of stainless steel is not permitted.

All pressure boundary welds in carbon steel materials greater than 1 1/2" thick require a minimum preheat and minimum interpass temperature of 200 F and subsequent PHHT for the times and temperatures specified on the weld procedure technique sheets.

All pressure boundary welds in carbon steels greater than 3/4" and less than or equal to 1 1/2" thick require a minimum preheat and minimum interpass temperature of 200 F. Post weld heat treatment is not required. All welds in carbon steel 3/4" (inclusive) and less require a minimum preheat and minimum interpass temperature of 50 F. Post weld heat treatment is not required.

D.13 CONTROL OF STAINLESS AND FERRITIC STEEL HELDING Fabrications of welded austenitic stainless steel Classes 1, 2.

and 3 comply with the requirements of ASME Section III and Section IX as supplemented by the BVPS-2 position on NRC Reg.

Guide 1.31 and Reg. Guide 1.44 contained in 2BVS-920. A typical BVPS-2 stainless steel butt weld is performed in the following manner. For stainless steel piping, a TIG root and hot pass is put down first. This is done with an ER308 insert for the root pass and bare filler wire for the hot pass. This first segment is 1/8" to 3/16" thick with nominally two layers from three

, beads. After this, the balance of the weld is made from E308 l covered electrodes using the shielded metal arc welding process.

For each weld, a count of the number of weld wires and electrodes required to make the weld is maintained. The heat input..

directly related to the size of weld bead and an indication of sensitization of the base metal, is nominally 20,000 - 50,000 Joules / inch. End prepping usually consists of a J-bevel and no post weld heat treatment is performed.

D-5

Certain shop welds included in the HHIPJET program were made with I the above practices except that Type 316 weld filler metal was  ;

used. Also, the submerged arc welding. SAH, process was used to ,

complete some of these shop welds. In these instances. Type 316 l

weld filler metal with Cr-Ni Alloy flux was used. The heat input range for shop welding processes is nominally 12,000 to 50,000 Joules /in.

Carbon steel pipe welds included in the HHIPJET program are made using GTAH's and SMAH's. The weld filler metals used for these processes are Type E70S-2 and E7018 respectively. The heat inputs used for these processes are nominally 20 kJ/in to 80 kJ/in based on the electrode sizes specified in the approved procedures.

REFERENCES D.1 Determining Fracture Properties of Reactor Vessel Forging Materials. Heldments, and Bolting Material EPRI NP-122, July 1976.

D.2 Impact Testing of the HAZ, P.F. Ivens and A.A. van den Bergh. Metal Construction and Brittish Helding Journal, July 1974.

D.3 H.S. Mehta, S. Yukawa, S. Ranganath, " Flaw Evaluation Procedures for Ferritic Piping," EPRI Project RP-2457-2, Electric Power Research Institute, Palo Alto, California, August 1985.

9 D-6

TABLE D.1 COMPARISON OF BVPS-2 HELDING TECHNIQUE Process BVPS-2 GE

• Root weld GTAN (manual) GTAN (manual)

Filler weld SMAN (manual) SMAN (manual)

Filler metal ER308/E308 ER308/E308 Filler rod dia. 1/16" - 1/8" 1/8" Argon gas flow (CFH) 15-25 20 Amps60-150 90-120 Volts (GTAN) 8-14 14 Volts (SMAH) 20-24 19-22 Preheat temperature 60 F (min) 60 F (min)

Interpass temperature 350 F (max) 350 F (max)

Stress relieving N/A None

• Taken from a GE procedure for an as-welded butt weld e*

8

4 er APPENDIX E LEAK RATE CALCULATION METHODOLOGY O

-a- _ . - _ . _ . . _

APPENDIX E LEAK RATE CALCULATION METHODOLOGY This appendix describes the computer program PICEP (Pipe Crack Evaluation P,rogram) [Em l] that calculates the crack-opening area.

the stable crack length, and flow rate through cracks in pipes.

The program evaluates the leak-before-break (LBB) hypothesis for appropriate loads, material, crack length, and pipe geometry.

IL 1 CRACK OPENING AREA CALCULATION Crack opening area for a through-wall circumferential flaw is computed by using the crack opening displacement formulas of

! Kumar and German [E. 2 ] and assuming a rectangular. diamond. or elliptical shape.

The plastic contribution to the displacement is computed by summing the contributions assuming pure bending and pure tension, a conservative LBB procedure that underestimates the displacement from combined tension and bending. Applications cover most of the practical range of material properties. pipe geometries, and I crack lengths.

Figure E.la compares measured (E. 3 ] and computed displacement in a 6.5-inch outside diameter, t= 0.425 inch. type 304 stainless steel pipe with a 90-degree circumferential crack subjected to an increasing bending load. Figure E.1b given the measured stream-strain curve for the material and the analytical fit used for the PICEP analymim. The best results for this calculation were found by matching the data in the 0.25% strain range. Other crack opening displacement comparisons are given by German et al.

[E.4].

i Evaluation of axial flaws is given in the PICEP report (E.1 ] .

E.2 CRITICAL FLAN LENGTH I The critical flaw length in the maximum stable flaw length for a i

given load. For austenitic stainless steel and its weldments.

l the critical flaw mize may be computed using the equations in Appendix C to Section XI of the ASME Boiler and Pressure Vessel Code. These equations have been implemented in the PICEP computer program.

The fracture toughness of carbon steel piping requires consideration of a wider range of fracture mechanisms and the equations of Apppndix C may be inappropriate. For these steels.

the critical flaw length is computed by using other methode proposed by the Section XI Task Group on Piping Flaw Evaluation, 1

t l

E-1

The validity of these procedures is discussed in ASME technical

support documents [Em5 E. 6 ) .

E.3 LEAK RATE CALCULATION An analytical model was developed to predict flow rates from

blowdown of initially subcooled or saturated liquid through I cracks based on Henry *a homogeneous non-equilibrium critical flow model [E.7) with several modifications made for use in PICEP

[E.8 E. 9 ] . The model has been extended to cover the flow of

maturated water-steam mixtures and superheated steam [E. iO ). The model ammumes the flow to be isenthalpic and homogeneous, but it l

accounts for the non-equilibrium "f l ashing " mass transfer process j between the liquid and vapor phases. Fluid friction due to surface roughness of the walls and curved flow paths has been incorporated in the model. Flows through both parallel and I convergent cracks can be treated. Due to the complicated geometry within the flow path, the model uses some approximations and empirieml factors which were confirmed by comparison against ,

test data.

The general features of the discharge of initially subcooled or saturated liquid through a crack are shown in Figure E.2 and the configuration of a convergent crack is shown in Figure E.3. In the region 0 i L/D i 3. a liquid det surrounded by a vapor annulus is formed. For lengths between L/D = 3 and L/D = 12 the liquid det breaks up into droplets at the surface, and small bubbles are entrained within the det. It is assumed that no mass j or heat transfer takes place between entrance and L/D = 12 Non-equilibrium effects are introduced through a parameter, which is a function of equilibrium quality and flow path L/D.

1

, For given stagnation conditions and crack geometries. the leak rate and exit pressure are calculated using an iterative search for the exit pressure starting from the saturation pressure corresponding to the upstream temperature and allowing for friction, gravitational, acceleration and area change pressure dropa. The inertial flow calculation is performed when the critical pressure is lowered to the back pressure without finding a solution for the critical mass flux.

A conservative' methodology was developed to handle the flow of wet or dry steam through a crack. The same model was used except that the single phase flow equations were not used. Rather. the flow was assumed to be two-phase from the entrance to crack exit.

, To make the model continuous, a correction factor was applied to adjust the mass flow rate of a saturated mixture to be equal to that of a slightly subcooled liquid. Similarly, a correction j factor was developed to ensure continuity as the steam became 1

superheated. The superheated model was developed by applying thermodynamic principles to an isentropic expansion of the single phase steam.

i i

j E-2 L

.w-_.-_ _,,,, m, , _ - - , - _ . .-.-.,.-c_-.,.---_,,m . , , . . _ _ . . - . - , , . - _ _ _ , , _ , , ,,, .r__ . _ , . , . , , , . - . _ _ _ , . ,

The code can calculate flow rates through fatigue or IGSCC cracks and has been verified against data from both types. The crack surface roughness and the number of bends account for the difference in geometry of the two types of cracks. The guideline for predicting leak rates through IGSCCs when using PICEP was based on obtaining the number of turns that gave the best agreement for Battelle Phase II test data (see section E.3.1.1).

For f atigue cracks, it is assumed that the crack path has no bends.

E.3.1 Verification with Experimental Data f

The experimental data for flow through tight cracks under 2-phase

! flow conditions is scarce. PICEP has been assessed using several i

open literature and proprietary data sets. Key verification j examples follow.

j E.3.1.1 Battelle Columbum Laboratories (BCL) Data I

This program [E.11 ] funded by EPRI generated data to confirm critical flow models used to predict leak rates through cracks.

Phase I tests used simulated cracks in which test variables included L/Dh ratio, stagnation temperature. stagnation pressure, and crack face surface roughness. The variation in L/DA ratio was provided by using three different specimens. The surface roughness was varied by machining the crack faces to the l desired roughness. The crack faces were initially ground to an average surface roughness of 0.039 mila as measured by a surface profilometer. After completing the ' smooth' tests, the flanges were removed, disassembled and the crack surfaces were roughened by shot blasting to an average roughness of 0.24 mils. The flanges were reassembled and installed on the test vessels. The minimum crack width opening was measured and a new L/Dg computed. For the final set of tests, the crack faces were roughened to 0.4 mila average roughness, and L/Dh was recalculated, based on new crack width measurements. Predicted versus measured flow rates for simulated cracks are shown in Figure E.4.

l l The Phase II tests used a test specimen (Figure E.5) with a 90%

l through-wall circumferential intergranular stress corrosion (IGSCC) crack. A portion of the outer pipe surface was removed to expose the crack tip and progressively wider cuts resulted in the variation of the L/D parameter. The maximum crack length was 1.1 " .

Following the Phase II tests, the pipe containing the IGSCC crack was cut and an exit to entrance area ratio for cracks was

approximated by measuring the slope of crack convergence. A crack wall surface roughness of K = 0.0051 mm was used. The flow ,

path was assumed to contain twenty 45 degree turns.

1 1

E-3

Figure E.6 shows that PICEP results are in good agreement with measured leak rates for testa 19 to 81 with crack lengths from

0,38" to 1,1", The leak rates for very narrow cracks (testa 7 to i 18 where 6 = 0,02 mm) are overpredicted because of high likelihood of plugging in the short cracks.

E.3.1.2 Duane Arnold Safe-End Crack Plant Data From June 14, 1978 to June 17, 1978 the Duane Arnold nuclear plant had an unidentified leakage source of 3 gpm into the primary containment sump [E.12], The leakage was below the i

plant's technical specifications of 5 spm and the plant operation continued between June 14 and June 17, 1978 The reactor, however, scrammed on June 17, 1978 during performance of weekly control valve tests. Following shutdown of the plant on June 17, plant personnel found the leak to be from a crack in one of the

eight recirculation-inlet-nozzle safe ends. These safe ends are approximately 12 inches in diameter and are used to facilitate welding of the stainless steel inlet piping to the carbon steel reactor vessel nozzles. The Duane Arnold recirculation inlet configuration is shown in Figure E.7a. The safe ends were manufactured from a nickel-base alloy (Inconel 600). The thermal sleeve shown in Figure E.7a contributed to crevice corrosion in the Inconel 600 The safe ends were examined thoroughly with UT, These examinations clearly showed that all eight safe ends were cracked essentially completely around the circumference. The crack penetration from the inner surface typically ranges from 50% to 75% of the wall thickness, except in the leaking safe end where an 80 degree segment had a through-wall crack. The general characteristics of the crack geometry were later confirmed by a I destructive examination of the leaking safe end and one of the remaining safe ends with a part-through crack. Figures E.7b and E 7c illustrate the general character of the cracking in the

! leaking safe end.

I l

The PICEP model used a through-the-wall crack with inside crack length of 7.2 inches and the original pipe thickness of 0,65 inches. It was not necessary to model the compound crack. The critical crack length for the PICEP model in 19.9 inches.

The following plant input data was used for the PICEP calculations:

, Pressure: 1050 psia Temperature 550 F

! Outside Diameter: 12" I Thickness: 0,65"

) Crack Surface Roughness: 2x 10E-04 inchen I No, of 45 degram Turns: 17 Crack Shaper Elliptical Crack Orientations- Circumferential i

E-4 i

l l

i Crack Type: ICSCC Young's Modulus: 28.9 kai Yield Stress: 42.0 kai Flow Stress: 75.0 kai l Work Hardening Exponent: 10.0  !

Work Hardening Constants 3.5 {

Bending Moments 0 (Case 1)

Bending Moment: 225.000 in-lbf (Case 2)

Measured Crack Length: 7. 2 " (inside)

Measured Crack Flows 3 gpm The bending moment of 225.000 in-lbf was obtained from the stress report (personal communication from H. Mehta and S. Ranganath of General Electric. San Jose) by using the SRSS of My and Mz moments. The torsional moment Mx was not considered as this does not act to open the crack. The resultant bending moment Mr =

(My 2

, g2 2 ) u4 provides a varying stress field along the circumference of the pipe. Since the crack was not exactly at the location of maximum bending stress, the problem was bounded I

by making two runs. One used no contribution from bending moment and the other used maximum contribution from bending moment. The l results from these two calculations bound the measured leak flow l rate data as shown in Figure E.7d. These two calculations also bound uncertainties in the bending moment due to the attached thermal sleeve and crack plugging and unplugging effects due to corrosion products.

4 E.3.1.3 UC Berkeley Data

C. Amos and V. Schrock (E.13 ] funded by the USNRC obtained

experimental data for critical flashing flow of initially i

subcooled water through rectangular slits. The influences of stagnation pressure, stagnation subcooling, and slit opening dimension on both the critical mass flux and the pressure profile within the slit, were studied parametrically. Stagnation pressures ranged between 4.1 and 16.2 MPa. Subcooling was in the range of zero to 65 C. Length-to-diameter ration (L/D) of the slits were between 83 and 400, with the length in the flow direction fixed at 6.35 cm. (This is a typical value for the thickness of a nuclear reactor primary cooling system pipe).

The test matrix consisted of 60 tests. Each of three slit opening dimensions. (0.127 mm. 0.254 mm and 0.381 mm) was tested at 5 pressure levels (4.2 MPa. 7.2 MPa. 9.6 MPa. 11.6 MPa. and 15.6 MPa), and with four degrees of subcooling (60 30, 15 and 3 C). Five of the highest pressure tests were not completed because of leakage problems encountered when operating the system at high pressure. Slit openings were selected to be of a size such that leakage through cracks of that dimension. with a length corresponding to 50% of the circumference of a reactor primary system pipe, would resul't in a significant loss of cooling water.

E-5

Stagnation pressures and subcoolings tested cover the range of conditions which occur in Pressurized Hater Reactors (PHRs) and Boiling Nater Reactors (BWRs) under both normal operating conditions, and during postulated accidents. In addition to the two phase tests, each test section assembly was calibrated using cold (25 C) water, A stagnation pressure of approximately 2,5 MPa was used for these tests, Data from these runs were used in determining the single-phase friction factor and entry loss coefficients for the slits, Figure E.8 shows that PICEP results are in very good agreement with measured leak rates.

E.3.1.4 Canadian Fatigue Crack Data PICEP predictions were compared with the data taken by D. Scott and A. Cook [E.14 ) for critical flow of initially subcooled water through cracks in thick wall piping. The specimens were all made from Schedule 100 ASTM 106 Grade B pipe. The pipe diameters tested ranged from 305 to 610 mm. To make the specimens, a length of pipe approximately twice the pipe diameter was used.

An Electric Discharge Machine (EDM) notch was then placed at the center of the pipe on the inner surface to initiate the crack.

The notch was 0,127 mm wide x 5.1 mm long x 2,54 mm deep. In some cases, an additional stress riser was added by welding a flange on the inside surface of the pipe. Pressure cycling was used to grow the crack. The specimen was attached to a high temperature loop which supplied water at about 250 C and 8,4 MPa.

The water in the loop was kept at a pH of 9,7 with about 40 ppb of oxygen. Strap heaters could supply 4,9 kW of heat and could bring the temperature of the pipe up to 290 C. The temperature was controlled between 255 and 265 C.

On completion of leak testing, each crack was examined. This examination consisted of:

(a) cutting out the section of pipe containing the crack, typically 0.1 m x 0,3 m (b) measuring the crack length on the pipe inside diameter and outside diameter (c) measuring the crack opening on the inside and outside diameter using a microscope (d) cooling the section in liquid nitrogen and then cracking it open to expose the leak path (e) remeasuring the crack length on the pipe inside and outside diameter, and (f) measuring the surface roughness of the leak path, E-6

Figure E.9 shows that PICEP results are in good agreement with measured leak rate.

4 An important finding of this work concerns the plugging and i

unplugging phenomena observed for short crack lengths. For l larger crack lengths, plugging phenomena are not important '

because the crack opening is too large for this to occur in the middle section of the crack. It might still occur at the ends of the crack where the opening is tight but the impact on showing LBB is insignificant. This is also illustrated by the crack opening displacement of the near critical crack size in Figure E.10 tested at Hyle Labs for EPRI to demonstrate LBB [E.15).

l This alleviates the concern that a crack might not be detected due to plugging of the crack.

E.3.1.5 ANL Experimental Data D. Kupperman and T. Claytor [E.16] funded by USNRC took experimental data to evaluate and develop improved leak detection systems. The primary focus of the work was on acoustic emission oetection of leaks. Leaks from artificial flaws, laboratory-generated ICSCCs and thermal fatigue cracks, and field-induced IGSCCs from reactor piping were examined. Water at temperature and pressure of 500 F and 1100 psia respectively is supplied to a small pressure vessel which is welded to the inner surface of the pipe and vary the crack opening.

Calculated and experimental leak rates through an IGSCC and a thermal fatigue crack are shown as a function of applied stress in Figures E.11 and E.12. As expected f rom f racture mechanics ,

the flow rate is in most cases proportional to the applied stress. This linear dependence breaks down for low stresses.

ANL speculated that the residual stresses produced when the crack is welded into the piping system may tend to hold the crack shut until a certain threshold level of stress is reached. Since the j estimate of the initial crack opening area (zero applied stress) was not known, two curves are shown from PICEP calculation. The first assumes no initial crack opening area and the other assumes I an opening that yields a flow rate equal to the zero applied stress value. The overall comparison and the trenos are good.

E.3.1.6 CREC Experimental Data M. Cumo et al. (E.17) at CREC, Italy measured the flow of subcooled water, saturated water and naturated steam through a long (1.5 meter) small diameter (5 mm) tu be . The conditions j testa covered a range of pressures (11-24 bars) and subcooling (0 to 47 F). Their data is presented in Figure E.13 together with the PICEP prediction. The agreement is very good and the predictions are generally conservative,

~

i l

E-7

l E,4 PICEP SAMPLE CALCULATIONS Three sample calculations are provided to show the general trends predicted by PICEP, Figure E.14 shows the effect of crack length on flow rate with subcooling as a parameter. For a fixed geometry surface roughness and inlet pressure conditions. the predicted critical mass flux increases as subcooling increases,

'The influence of crack depth-over-hydraulic diameter (L/D) ratio is shown in Figure E.15. The critical mass flux decreases uniformly with increasing fL/D ratios, This is consistent with the assumption that the mixture quality approaches the long tube value in large L/D cracks, The effect of stagnation pressure is shown in Figure E,16 For a fixed geometry, surface roughness and subcooling, the predicted mass flux increases with stagnation pressure.

l

( .

i E-8 l _

REFERENCES E.1 Norris. D.. Okamoto. A.. Chexa1. B., Griesbach. T..

"PICEP: Pipe Crack Evaluation Program." EPRI Report NP-3596-SR. Special Report. August 1984 E.2 Kumar. V.. et al.. " Advances in Elastic-Plastic Fracture Mechanics." EPRI Report NP-3607 Palo Alto.

California. August 1984 E3 Personal communication from Kuzuo Kishida of IHI to D.M. Norris of EPRI. April 1 1986,

. E.4 German. M.D., et al.. " Elastic-Plastic Fracture Analysis of Flawed Stainless Steel Pipes." EPRI Report NP-2608-LD. September 1982 E.5 " Evaluation of Flaws in Austenitic Steel Piping." ASME Code.Section XI. Task Group on Piping Flaw Evaluation. D.M. Norris. Editor. April 1986 To be published in the Journal of Pressure Vessel Technology.

E.6 " Evaluation of F1mus in Ferritic Piping." A. Zahoor.

et al.. EPRI Research Project RP1757-51 Oct'ober 1985 For consideration by the ASME Code.Section XI. Task Group on Piping Flaw Evaluation, i

E.7 Henry. R.E., "The Two-Phase Critical Discharge of Initially Saturated or Subcooled Liquid." Nuclear Science and Engineering. Volume 41 1970

E.8 Abdo11ahian. D. and Chexa1 B., " Calculation of Leak Rates Through Cracks in Pipes and Tubes " EPRI Report l NP-3395 January 1984 E.9 Chexa1. B., Abdo11ahian. D., and Norris. D.,

" Analytical Prediction of Single Phase and Two-Phase Flow Through Cracks in Pipes and Tubes. " 22nd ASME-AIchE National Heat Transfer Conference. Niagara Falls, j August 1984, i

E 10 Chexa1 B, and Horowitz. J., "A Critical Flow Model for

Leak-Bef ore-Break Applications. " submitted f or l presentation at the SMIRT-8. Conference. Lussanne.

Switzerland. August 1987 -

i E.11 Collier. R.. et'a1.. "Two-Phase Flow Through i Intergranular Stress Corrosion Cracks and Resulting l i Acoustic Emission." EPRI Report NP-3590-LD. Final Report, April 1984, t

i E-9 l . _ _ .

References (continued)

E.12 i

" Investigation and Evaluation of Stress-Corrosion '

Cracking in Piping of Light Mater Reactor Plants."

NUREG-0531 Pipe Crack Study Group. February 1979. )

)

1 E.13 Amos. C.N. and Schrock. V.. " Critical Discharge of Initially Subcooled Hater Through Slits. " NUREG/CR-3475 LBL-16363 September 1983 l

E.14 Scott. D.A.. " Leak Rates Through Cracks in Thick Halled Piping." AECL Report. June 1983 E.15 Bausch. H.P.. et al.. " Pipe Rupture and Depressurization Experiments." EPRI Research Project 2176-1 January 1986 Draft Report.

E.16 Kupperman, D., Shack. H.J., and Claytor. T., " Leak Rate Measurements and Detection Systems." Argonne National Laboratory. CSNI LBB Conference. Monterey. California.

September 1-2, 1983 E.17 Cumo. M. et al.. " Experimental Results on Critical Flow Internal Technical Report. NDAG 2TS4B020 Centro Ricerche Energia Casaccia. October 1982.

I E-10 i

l

0.05 . . . .

_ g g. . . . . , .

0.04 -

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~

0.03 -

y _

E

_~  :

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EXPERIMENTAL DATA -

g 0.02 u .

0.01 - PICEP CALCULATION _._

0.00 ' ' ' I I ' ' ' ' 04 .I '

0 0.2 0.6 ' ' ' 08 .l ' ' ' ' ~ 1 ob/cy 50 .. . .

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? -

h 30 -

,$ 5 - ;  :

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g  ; [HI DATA -

g 10 - -

0

' ' ' ' 02 .I ' ' ' I ' ' ' ' 06 .I ' ' ' ' 08.I ' ' ' '

O 0.4 1 ELONGATION (7.)

Figure E.1 (a) Comparison of measured versus computed crack-opening displacement versus bending stress. The stress is normalized with the 0.2%-offset yield stress 2

of24.8kg/mm;(bIstress-straindataandtheRamberg-Osgoodfitusedforthe 2

PICEP calculation: c o=24.8 kg/mm , a=1.0, n=5.0, co =0.125%. Data is at 20*C.  ;

1

rm. ein.

er.ca ChShing hbwassswm-sw__m%%wJ .. . ..

/

• e e = -

ggw%%wemmwwwwm%NNNNNNww%%%%g N Sebcooles Jet 9 ,,wie jet w e.6.ae w 7"* * "**

• d " **

Figure F.2. Two-Phase Flow Through a Long, Harrow Crack Length 9 Width 6

, 2e ,

c , l ta , mene 4 .i_

e. r - /-

,/ 4, tarea e vo = m

/ /

L' ,/ 4 ,zue, m.a.

cerectea .f Mew Figure F.3. Geometry of a Convergent Crack i

e I

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f . . - - . -

• _ _ - _ - .

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SECTION A- A Figure E.5. Sketch of Crack Details e

' " " - '"---w --- w _

3AT~E.. GSCC CX'AE > >>,-

[

5

- 100 .

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Cl-X Tests 1-6

Tests 19-81  !

Q Tests 7-18

' ' '"l ' ' ' ' ' ' " ' '

10-4 ' ' ' ' " ' ' ' ' ' ' " ' ' ' ' ' "

10-4 10-3 10-2 10-1 .

100 !O I MEASURE 0 lE AK RATE (GPMI Figure E.6. BCL Phase II Results

l 1

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! Figure S.7a Duane Arnold Recirculation Inlet j (Old Configuration)

From Drawing No. VPF 2655-96

-_ _ _ _ _ _ - _ _ . _ _m

L.

$ 3 c:: 4 arm ac

. =.m. cua

--- m.m. c . . .

Figure E.7b Representation of IGSCC in Duane Arnold Leaking Recirculation-Inlet-Nozzle Safe End l

W 0

- -- _ _ - - - - - , - - - .-.-,,,,_----,-.-_y----- -

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f s 32 t-  : s' NA* M 2 M d2 FRACTOGRAPMIC $P(C:Mus M Sveract Deposit 5,tcimens 1

Figure E.7c. Safe End N2A i

l l

I DUANE ARNOLD ENERGY CENTER - NO BENDING 6 _

.g....,,...;,..., ,,

~

1 CRITICAL CRACK SIZE - 19.934 IN

, 5 _

7 .

CASE 2:

r 4 --

PRESSURE + BENDING 7

a -

e -

.: tJ  :  :

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e _

x  : CASE 1:  :

)

PRESSURE ONLY

.; a 2 .

1 _

• ~

' ' ' ' ' ' ' ' l ' ' ' ' I ' ' ' ' i ' '-

0 O 2 4 6 8 10 CRACK LENGTH,1NCHES i

i i ,

I e

)

Figure F.7d. -

4 i

4

-_..,.~,-w,--- , . . ,. . , _ < . , . , - . . .,.,r,,,....,--3y,._.,,.,__,,-_,,w_ _,,,,,,,,..,.,_,,,..~,,mw,..m-, ,,~.,,-,,.-,,.,,.,%--,

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S -

X M Slit Width =0.381 m U -

X }

+

' 0 -

O Siit Width =0.2s4 m

y X Slit Width =0.127 m 1 W O_

4 0.1 ' ' ' ' ' ' ' !I O.1 0 .'5 1 5 1O i

MEASURED l.EAK RATE (GPM)

Figure E.8. UC Berkeley Slit Data vs PICEP Prediction

t C ANA JI AN = A~~::GU E C R AC <S 1.00 _

1 j

e-x _

0.50 -

x n^xinun tEAx roaa -

- 0 AVERAGE LEAK FLOW r -

i 1 ,

e w - g m

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F--

< c x x -

2

< 0.10 _

w _

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l w 0.05 -

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l 1 _

1 1

~ O.01 ' ' ' ' ' ' I ' ' ' ' ' '

1 o.01 0.05- 0.1 0.5 1 1 MEASURED LEAK RATE (GPM)

Figure E.9.

Canadian Fatigue Crack Data vx. PICEP Prediction l

~

lY

.r .. ..

- mmx.w.; y v +

.... f5M' MAO;t't<

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'k Figure E.10 Photograph of Pipe with a Near-Critical Ciretmferential Crack l

-- - - - - - - - - - - - , - - - - , ~ - - - - - - - - - - - - - - - - - - - - - - - - * * - -

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ANL DATA ,/ f

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O S I i.I 1b 20 API'l 11' l'1 ' . I Rl!SS Iksil Figure C Il ANL IGSCC Data vs PICEP Prediction i

i

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/

INITIAL OPENING i . - <.- -

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I Figure F.12. ANL Fatigue Crack Data vs PICEP Prediction l

CUP 0 et e. -

ROUN J 5mm TUBE C_/J=300)

I 4 i i i I I 4 i i 4 I 5.0 -

I -

X SUBCDCLED VATER -

k M SATURATED WATER

+

SATURATED STEAM M _

L1J M F-

< . M cr 1.0- -

y _

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F~ _

+

b -

o _

W 8 Cf _

O.1 ' ' ' ' ' ' ' ' I ' ' ' '

o.1 0.5 1 5 MEASURED LE AK RATE (GPM)

Figure E.13 CREC Data vs. PICEP Prediction

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

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

I

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10 1: R Al :lf I I Hn iII.lNr.'lll! S Figure E.14 Parametric Effect of Subcooling on. the Predicted Leak Rate Flow

I i

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i figure E.15 Parametric Effect of FL/D on Predicted Leak Rate flow i

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!^

APPENDIX g BVPS-2 LEAK RATE CURVES 4

0

i APPENDIX _F_

BVPS-2 LEAK RATE CURVES l

The following leak rate curves give the leak rate (in gallons per minute) from a crack in the BOP piping as a function of crack length (in inches). The leak rate curves correspond to the lines and breaks tabulated in previous sections of this report.

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9 APPENDIX Q FLET VERIFICATION CALCULATIONS

APPENDIX G FLET VERIFICATION CALCULATICNS G.1 INTRODUCTION FLET (Flaw Evaluation by Tearing Instability) [G.1) is software that computes the load for growth of an existing crack or the load for crack-growth instability. The program considers plasticity in the material near the flaw, a feature that makes it applicable to the ductile materials used in the pressure boundary of nuclear power plants.

G.2 THE FRACTURE MECHANICS OF FLET Prior to EPRI-supported research in elastic-plastic fracture starting in 1979 no simple general estimation formulas existed for flawed ductile materials. Each crack geometry and material was considered on a separate basis and generally required expensive elastic-plastic finite element analysis to determine a flawed component's fitness-for-service.

The methodology contained in FLET is the simplified elastic-plastic fracture mechanics of Kumar, German and Shih [G.2]

developed with EPRI support. The accuracy of the estimation formulas and procedures has been extensively checked by comparison with specimen tests and elastic-plastic finite element calculations. In later work, directed to cylinders, German et al. [G. 3 ) compared estimation scheme results with finite element calculations and pipe experiments and concluded that "... pipes can be modeled with sufficient accuracy by (these) 2-dimensional engineering approach solutions." Additional piping solutions were added by Kumar et al. [G.4, G.5] that further confirmed the accuracy of the technology by comparison with finite element calculations.

More recently, Zahoor and Gamble [G.6] have extended the verification of the methodology for piping by comparing crack initiation and instability data from tests of ;20 pipes containing both circumferential through-the-wall and surface cracks. In both cases they conclude that the estimation procedures provide sufficient accuracy for engineering applications.  ;

G.3 DESCRIPTION OF THE FLET SOFTWARE EPRI developed the program in early 1984 for use with round robin piping calculations held in San Antonio, Texas on June 22-23 l [G.7]. Development has continued and the EPRI results reported here use the current version. The code contains the estimation l fracture mechanics methods described above.

I G-1

The program includes all of the piping solutions given in References G.2 through G.S.

For the leak-before-break piping application considered here, the solutions for the circumferential through-the-wall cylinder crack in combined tension and bending are the most important. The software contains the tables of h functions with appropriate interpolation algorithms and perf orms computer simulation of the graphic procedures described in detail in Reference G.2.

Three solution options are available to the user. These are the J-integral tearing instability analysis (JT), the ductile fracture failure analysis diagram analysis (DPFAD), and the J-integral load analysis (JL). Except for small differences, usually each of these methods leads to the same crack initiation and instability loads [G. 7 ] . For the leak-before-break analyses in this report, the DPFAD option is used to account for the crack growth between initiation and the instability load.

G.4 FLET VERIFICATION CALCULATIONS The following discussion concerns the accuracy of the fracture mechanics methodology discussed above and implemented in the EPRI FLET computer program. Three different problems are compared for load at initiation and at instability.

G 4.1 J at Initiation - NUREG 1061 Volume 3. Figure A-7 This figure compared experimental initiation J values for 8-inch diameter, SA 106 Grade B steel pipes containing various length through-the-wall circumferential cracks with estimation solutions by the NRC Pipe Crack Study Group [G.8]. The load is pure bending.

A recalculation of time data using the FLET code with identical input and the corresponding plot of EPRI results is shown in Figure G.1. The plot also compares calculations by Zahoor and Gamble using a modified estimation formula [G 6] appropriate for this problem with a similar FLET calculation.

This modification corrects the plastic part of the J-integral formulas used for the circumferential through-the-wall pipe flaw.

The effect is to reduce the conservatism in calculation of crack-driving force as shown in Figure G.1.

The initiation J values are also given in Table G.1.

The FLET results agree well with Reference G.6 and reasonably model the experimental results.

G-2

3 G.4.2 Crack-Growth Instability Loads for Flawed Heldments EPRI compares FLET results for the load-carrying ability of flawed austenitic stainless steel flux weldments with calculation made in support of the evaluation procedures found in the ASME Boiler and Pressure Vessel Code,Section XI, INB-3640 [G.9]. The EPRI calculations use identical material properties and the same through-the-wall crack for comparison with the ASME calculation.

The analyses here are for a 28-inch diameter pipe in pure bending.

Results are shown in Figure G.2 for a submerged arc weldment and in Figure G.3 for a shielded metal arc weldment. FLET agrees well with the ASME results for the SAN weld and is somewhat more conservative than the ASME SMAN calculation. Both analyses use the unmodified estimation formulas.

G.4.3 Initiation and Instability Loads in Combined Tension and Bending These calculations compare loads computed with the three-dimensional finite element program ABAQUS [G.10]. the NO-BREAK program [G.11) (which uses an estimation procedure similar to FLET), and the FLET program. The comparison calculations are given in Reference G.12.

EPRI presents the results in terms of multiples of faulted load that are added to the Service level A loads. EPRI assumed the level A load is 564,670 lbs tension, and the faulted load is 36,000 lbs tension and 3,800,000 in-lb bending moment.

These service level A and faulted loads correspond to the normal loads and the SSE load respectively of this report for the specific line considered.

Table G.2 compares calculations for J/Jic, tearing modulus, and load at instability, Table G.3 compares calculations for J/Jic in greater detail, and Table G.4 compares load (J) and the tearing modulus (T) at instability.

The results for this set of calculations show fair agreement with the ABAQUS and NO-BREAK calculations.

G.5 CONCLUSIONS These comparisons show reasonable agreement between the FLET program, other calculational procedures, and experimental data.

EPRI concludes that the estimation fracture mechanics as implemented in FLET is acceptable for leak-before-break engineering calculations when used with the recommended factors of safety.

G-3

REFERENCES l G.1 Okamoto, A., and Norris, D., FLET: PIPE CRACK INSTABILITY PROGRAM, EPRI Draft Report, March 1986.

G.2 Kumar, V., German, M.D., Shih, C.F., An Engineering Approach of Elastic-Plastic Fracture Analysis, EPRI Report NP1931. July 1981.

G.3 German, M.D. et al., Elastic-Plastic Fracture Analysis of Flawed Stainless Steel Pipes, EPRI Report 2608-LD, September 1982.

G.4 Kumar, V., German, M.D., Hilkening, H.H., Andrews, H.R., de Lorenzi. H.G., and Mowbray, D.F., Advances in Elastic- Plastic Fracture Analysis, EPRI Report NP3607, August 1984.

G.5 "New Estimation Scheme Solutions " personal communication from V. Kumar to D.M. Norris, EPRI, February 1986.

G.6 Zahoor, A., and Gamble. R.M., Evaluation of Flawed Pipe Experiments. Final Report, EPRI Project RP 2457-8 July 1986 (in press).

G.7 Kanninen, M.F., editor, Proc. CSNI/NRC Horkshop on Ductile Piping Fracture Mechanics, San Antonio, Texas, 21-22 June 1984 (draft report to be published as USNRC NUREG report).

G.8 Report of the U.S. Nuclear Regulatory Commission Piping Review Committee, Evaluation of Potential for Pipe Breaks, NUREG-1061 Volume 3, USNRC, Hashington D.C.,

November 1984.

G.9 Evaluation of Flaws in Austenitic Steel Piping, prepared by Section XI Task Group for Piping Flaw Evaluation ASME boiler and Pressure Vessel Code Committee EPRI Report NP- 4690-SR, Electric Power Research Institute, Palo Alto, California, July 1986.

G.10 ABAQUS-ND: A Finite Element Code for Nonlinear Dynamic Analysis, EPRI NP-1552 CCM, Vols. 1-4, August.1980.

G.11 NO-BREAK, Fracture Mechanics Software for Piping Flaw Evaluation and Leak-Before-Break Analysis. User's Manual. Revision I, Novetech Corporation. Rockville, MD, January 1986.

G.12 Personal communication from B. Kee of Ontario Hydro to D.M. Norris of*EPRI June 1986.

G-4

l TABLE G.1 COMPARISON OF INITIATION J FOR EIGHT-INCH DIAMETER FERRITIC STEEL PIPES l

1 l

J-T METHOD DPFAD METHOD Ref. G.6 FLET Experiment Experiment Original Improved Original Improved Number EPRI EPRI EPRI EPRI 3 3,680, 13.261. 6,100. 13,230. 6.058.

7 5,400. 12,474. 5,800. 12.420. 5,802.

8 4,420. 7,579. 3,750. 7,512. 3,727.

11 2,340. 7,565. 3,650. 7,518. 3,571.

12 3.110. 11,802. 5,300. 10,700. 4,826.

14 4,300. 8,013. 3,650. 7,960.

3,589.

15 2,850. 10.510, 4,720. 10,400, 4,621.

S

TABLE G.2 COMPARISON OF J/JIc TEARING MODULUS, AND LOAD AT INSTABILITY INSTABILITY ABAQUS NO-BREAK DPFAD METHOD (J-T) (FLET)

IMPROVED ORIGINAL IMPROVED EPRI EPRI EPRI J/JIc 4.857 4.667 4.005 3.992 T 14.750 15.950 na na Faulted Load 5.194 5.374 4.734 5.220 Multiple

TABLE G.3 COMPARISON OF J/JIc Faulted Load Multiple ABAQUS NO-BREAK DFFAD METHOD (J-T) (FLET)

IMPROVED ORIGINAL IMPROVED EPRI EPRI EPRI 1 0.141 0.130 0.151 0.149 2 0.415 0.358 0.436 0.414 3 0.967 0.814 0.989 0.872 4 2.033 1.712 2.192 1.746 5 4.092 3.450 4.676 3.349 6 d.041 6.704 9.633 6.272 0

0

TABLE G.4 COMPARISON OF TEARING MODULUS Faulted Load Multiple ABAQUS NO-BREAK FLET *

(J-T) (J-T)

IMPROVED ORIGINAL IMPROVED EPRI EPRI EPRI 1 0.37 0.30 0.33 0.32 2 1.10 1.00 0.95 0.89 3 2.62 2.44 2.40 2.05 4 5.75 5.44 5.61 4.27 5 12.16 11.58 12.61 16.69 6 25.07 23.35 26.74 .28.62

• not applicable to the DPFAD method O

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.,