ML20151H624
| ML20151H624 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 04/30/1988 |
| From: | ROBERT L. CLOUD ASSOCIATES, INC. |
| To: | |
| Shared Package | |
| ML20151H605 | List: |
| References | |
| PROC-880430, NUDOCS 8804200452 | |
| Download: ML20151H624 (155) | |
Text
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COMANCHE PEAK-STEAM ELECTRIC STATION UNIT NUMBER 1 f
i I
i CPSES-1 WHIPJET PROGRAM REPORT' APRIL, 1988 1
I Prepared for Texas Utilities Electric by:
Robert L. Cloud & Associates, Inco'.porated 1
8804200452 000415 PDR ADOCK 05000445 A
PDR L
i TABLE OF CONTENTS SECTION PAGE 1.
Executive Summary 1-1 2.
Introduction 2-1 3.
Program Scope 3-1 4.
Screening Considerations 4-1 5.
Material Property Data 5-1 6.
Leak Detection 6-1 7.
Leak Rate Calculation Summary 7-1 8.
Critical Crack Size, Crack Size 8-1 Margin, and Load Margin calculations (Circumferential Cracks) 9.
Critical Crack Size and Crack 9-1 Size Margin calculations (Longitudinal Cracks) i 10.
References 10-1 APPENDICES A.
Fluid Transients and Pipe Cracking Incidents A-1 r
B.
Stainless Steel Material Properties B-1 C.
Welding Procedures C-1 D.
High Stress Locations D-1 E.
Leak Detection Systems E-1 1
F.
CPSES-1 Leak Rate Curves F-1 l
G.
Modified Limit Load Analysis by the Master Curve Method G-1 l
t 1
i t
k' TABLES Egi Title PAGE 3-1 Qualifying Piping Systems for WHIPJET 3-5 Analysis 3-2 CPSES-1 Piping Data for WHIPJET LBB 3-6 Analysis 3-3 CPSES-1 Pipe Breaks and Hardware 3-7 Associated with WHIPJET LBB Analysis 4-1 PWR Water Hammer 4-11 4-2 Reactor Coolant Chemistry Specification 4-12 7-1 Leak Rate Results for Stainless Steel 7-3 Lines Inside Containment 8-1 Circumferential Crack Stability Evalua-8-2 tion Flaw Margin Evaluations (Normal + SSE) 8-2 Circumferential Crack Stability Evalua-8-3 tion Load Margin Evaluations (1.4 x (Norcil+SSE) 9-1 Longitudinal Crack Stability Evaluation 9-2 (Normal + SSE) 9-2 Longitudinal Crack Stability Evaluation 9-3 (1.4 x (Normal + SSE))
A-1 CPSES-1 Flow Transient Applicability A-9 A-2 Potential Water Hammer Events A-10 B-1-1 Experimental Data for Lower Bound Tensile B-7 Curve for Type 304 Base Metal B-1-2 Typical Pipe Fitting Steels in B-8 CPSES-1 SIS, RCS and RHR Lines B-2-1 Comparison of CPSES-1 Weld Techniques B-25 for GTAWs B-2-2 Comparison of CPSES-1 Weld Techniques B-26 for SMAWs B-2-3 Comparison of CPSES-1 Weld Techniques B-27 for SAWS B-2-4 CPSES-1 (ITT) SAW Techniques B-28 B-2-6 GE SAW Techniques (Test Specimens)
B-29 11 L
TABLES (continued)
B;L.
Title PAGE 5
B-2-7 Experimental Data for Lower Bound B-30 Tensile Curve for Type 308 SMAW Welds AL
' 4. l B-2-8 Experimental Data for Lower Bound B '3 ;.
' q' e."
ti
~W Tensile curve for Type 308/316 SAW Welds 4
C-1 CPSES-1 Weld Summary for SIS, RHR, C-6 and RCS Lines D-1 CPSES-1 High Stress Locations For D-1 0
WHIPJET LBB Analysis E-1 Leak Detection Systems Sensitivity E-5 and Response Time Summary j
j E-2 Effect on Containment Parameterc' E-6 j ;
vs. The Type of Leak
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9 FIGURES
,h Title
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PAGE q
2-1 UHIPJET Program Step-by-Step Methodology 2-4 3
'.1 10" SIS to RCS told Leg Loop #1 3-9 l
i
(
3-4 10" SIS to RCS. Cote Leg Loop #4 3-12 3-5 12" RHR to RCS Hot Leg Loop #1 3-13 3-6 12" RHP to RCS riot Leg Loop #4 3-14 3-7 14" Pressurizer Surge Line 3-15 4-1 Corrosion Review 4-13 B-1-1 Yield Strength Comparisons for B-9 CPSES-1 and' (Jype 304 Stainless Steel Ba(e hetal l
B-1-2 Ultimate Tensile Strength B-10 I
comparisons for CPSES-1 and Industry Type 304lStuinless
. Steel Base Metal i
g, B-1-3 Yield Strengt( Couparisons for B-11 CPSES-1 and 1*ndustry Type 316 Stainless Steel Base Metal i
B-1-4 Ultimate Tensile Strength Comparisons B-12 for CPSES-1 and Industry Type 316 Stainless Steel Base Metal B-1-5 Stress-Strain Curves for SA376/312 B-13 Type 304 Stainless Steel Base efetal B-1-6 Stress-Strain Curves for SA376/312 B-14 Type 316L and 316 Stainles6 Steel Base Metal B-1-7 Lower Bound for Type 304 and Type 316 B-15 l
Stainless Steel Base Metal ja j/
fy i
N
[
4 l
PIGURES (continued)
Hot Title PAGE
)
B-1-8 Comparison of Experimental True B-16 Stress-versus-True Strain Curve s
/'
Against the Fitting Function (Lower Bound Fit Type 304)
.,b B-17 B-1-9 Comparison of Experimental Trud, Stress-vs.-True Strain Curve Versus
/
the Fitting Function (Best Fit
(
Type 316L) j
' B-2-2 GTAW Filler Metal B-32 B-2-3 Ultimate Tensile Strength comparisons B-33 for CPSES-1 and Industry GTAWs with
// I 308 and 316 Filler
,/
i B-2-4 Yield Strength Comparisons for CPSES-1 B-34 and Industry SMAWs with 308 Filler B-2-5 Ultimate Tensile Strength Comparisons B-35 for CPSES-1 and Industry SMAWs with 308 Filler B-2-6 Yield Strength Comparisons for CPSES-1 B-36 and Industry SMAWs with 316 Filler i
B-2-7 Ultimate Tensile Strength Comparisons B-37 for CPSES-1 and Industry SMAWs with l
316 Filler B-2-8 Ultimate' Tensile Strength Comparisons B-38 for CPSES-1 Uqd Industry SAWS with
(
Types 3,08'and,'316 Filler k
d-2-9 Stress-Strain Cdrves for Stainless B-39 Steel Weld Metal (GTAW) with j
l Type 308 Filler 1
l B-2-10 Strecs-Strain Curve Comparison of B-40 Sta:lnless Steel GTAWs and Lower Bound Type 304 and Type 316 Base Metal
,5-2-1 Stress-Strain Curver for Stainlessj B-41 Steel Weld Metal (SMAW)
B-2-12 Stress-Strain Curves for Stainless B-42 i
Steel Weld Metal (SAW)
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ME i
3:
sj' FIGURES (continued)
L No.
Title PAGE B-2-13 Stress-Strain Curve. Fit for Stainless B-43 Steel Weld Metal (SMAW)
'B-2-14 Stress-Strain Curve Fit for Stainless B-44 Steel Weld Metal (SAW)
F-1 CPSES-l' Leak. Rate Curve for 10-inch F-1 SIS Pipe, Circumferential Crack under Normal Operating Loads F-2 CPSES-1 Leak Rate Curve for 12-inch F-2
'/'
<y RHR Pipe, Circumferential Crack under Normal Operating Loads F-3 CPSES-1 Leak Rate Curve for 14-inch F-3 l
RCS Surge Line, Circumferential Crack under Normal Operating Loads F-4 CPSES-1 Leak Rate Curve for 10-inch F-4 SIS Pipe, Axial Crack under Normal l
Operating Loads-l F-5 CPSES-1 Leak Rate Curve for 12-inch F-5 i
RHR Pipe, Axial Crack under Normal Operating Loads i
F-6 CPSES-1 Leak Rate Curve for 14-inch F-6 l
RCS Surge Line, Axial Crack under Normal Operating Loads G-1 CPSES-1 RCS Surge Line Master Curve G-5 for Base /GTAW Material Properties
\\
G-2' CPSES-1 RCS Surge Line Master Curve G-6 l
for SAW/SMAW Material Properties l
G-3 CPSES-1 10-inch SIS Line Master Curve G-7 for Base /GTAW Material Properties G-4 CPSES-1 10-inch SIS Line Master Curve G-8 for SAW/SMAW Material Properties G-5 CPSES-1 12-inch RHR Line Master Curve G-9 for Base /GTAW Material Properties G-6 CPSES-1 12-inch RHR Line Master Curve G-10 for SAW/SMAW Material Properties vi L
Section 1 EXECUTIVE
SUMMARY
Fracture mechanics technology has advanced to the point that an engineering approach using the concept of leak-before-break (LBB) in lieu of postulating double-ended pipe rupture consistent with the broad scope rule change to GDC-4 is now possible.
The results for the WHIPJET program at CPSES-1 have been successfully applied on stainless steel lines inside containment greater than or equal to 10-inch nominal pipe size for the safety injection, residual heat removal, and reactor coolant (surge line) systems that have passed an initial screening process.
The screening process included evaluation for the potential for pipe cracking, excessive fatigue, water hammer, or other conditions which could result in the failure of high energy piping.
The overall results indicate that consideration of dynamic effects of postulated breaks (within the WHIPJET scope) is unnecessary.
I i
The evaluation for leak-before-break involved three key processes:
leak rate calculations for normal operating loads, crack stability analysis for normal plus seismic loads, and an excessive load case where stability is assessed for loads much higher than normal plus seismic.
Three margins were applied and satisfied for the lines analyzed at CPSES-1 as follows:
(1) a l
margin on leak detection of 10 inside containment on the minimum detectability limit of 1.0 gpm; this results in a 10 gpm value for assessing stability crack size; (2) a margin on crack size for assessing stability of at least 2.0; and (3) a margin of 1.4 on loads for the excessively high load stability check.
l The application of the WHIPJET LBB program at CPSES-1 has elimi-nated the concern over dynamic effects resulting from postulated pipe rupture, plus the need for undesirable and unnecessary pipe rupture hardware devices.
Additionally, much more is now known about the piping and its capabilities than was the case before this analysis was applied.
1-1
Section 2 INTRODUCTION WHIPJET is an engineering program, utilizing a leak-before-break (LBB) philosophy, which assures that essential structures, systems, and components are given protection at least equivalent to that provided by conventional design or hardware to protect against the dynamic effects of postulated pipe rupture.
Successful application of WHIPJET demonstrates that the fluid leakage from a postulated through-wall defect at the most controlling location (highest stress location concurrent with minimum material properties in terms of normal plus Safe Shutdown Earthquake loads) in a high energy piping line can be detected well before the possible rupture of the pipe.
The program has been previously applied to the Beaver Valley l
Power Station - Unit 2 for auxiliary [1] piping (i.e., other than the Primary Reactor Coolant Loop) as a lead plant l
application; this approach has subsequently been applied to other pressurized water reactor auxiliary piping.
This same approach is applied to the Texas Utilities plant Comanche Peak Steam Electric Station - Unit 1 (CPSES-1) and the results are presented herein.
Since WhIPJET is based on the LBB methodology, using either a conservative and generic or a localized specific elastic-plastic fracture mechanics approach for assessing the potential for pipe rupture, WHIPJET is consistent with accepted procedural recommendations and analytical criteria.
The basis for assurance of piping design and construction quality is presented in the CPSES-1 Final Safety Analysis Report (FSAR).
The WHIPJET program for CPSES-1 was limited in scope to only some of the stainless steel lines which had been previously determined to be affected by dynamic pipe rupture postulation.
l 2-1 i
Selected CPSES-1 system lines were screened for abnormal crack growth mechanisms (such as corrosion or gross global fatigue) and other potential failure mechanisms (e.g., water hammer).
Additional screening considerations included unexpected flow stratification conditions, erosion / corrosion, and other adverse industry experience.
Applicable stress calculation analyses were examined anchor to anchor on a system line basis to determine which locations (not limited to normal postulated pipe break locations) were most critical in terms of highest stress.
These locations were then coupled with all possible material properties (base metal and
'telds) in a deterministic evaluation that demonstrated sufficient margin against failure.
The pertinent system locations were analyzed to determine the stability of postulated through-wall flaws.
In flux welds, the loading conditions in the pipe (termed normal + SSE), were the combination of:
1 Safe Seismic deadweight
+
thermal
+
pressure
+
Shutdown
+
Anchor l
Earthquake Motion or:
l l
For the consideration of base metal and non-flux welds, thermal loading (expansion) was not included in the normal plus SSE.
Leak rate versus crack size was determined considering normal l
operating loads (DW + TH + P) and fluid conditions {2).
A limiting leak rate based upon a conservative detectable leakage limit of 1.0 gpm multiplied by a margin of 10 was used to determine the crack length corresponding to 10 gpm.
This leakage size crack length was analyzed for two margins against failure:
first, that a margin of 2 on crack size exists when compared to the (normal + SSE) loading condition critical crack 2-2
size; and, second, that a margin of 1.4 on the (normal + SSE) loading condition results in a critical crack size more than the leakage size crack.
Once these engineering criteria are satisfied, rupture protection hardware and consideration of postulated pipe break dynamic effects (including localized overpressurization) are not required for the lines analyzed.
In summary, WHIPJET addresses an alternative engineering approach for the provision of protection from the mechanistic effects of postulated pipe rupture.
Figure 2-1 shows the steps required in the WHIPJET program to satisfy the LBB approach.
i 2-3
Stainless Steel Lines Under Review Screening:
Industry Experience, Abnormal Crack Growth, Other Failure Mechanisms, Leak Detectability Margin Material Highest Stress Proper:i's Locations (Normal + SSE)
With Minimum Material Properties l
l 1Imiting Detectable Leak Rate Leak Rate x Calculations Normal Loads i
Margin of 10 Margin of 2 on Crack Size Margin Normal + SSE Crack Size Evaluation Loads Stability Check 1.4 (Normal + SSE)
For Excessively Loads i
High Loads FIGURE 21 WHIPJET PROGRAM STEP-BY STEP METHODOLOGY 2-4 L
Section 3 PROGRAM SCOPE 3.1 HIGH ENERGY SYSTEMS WITHIN WHIPJET The lines under consideration at CPSES-1 for the WHIPJET leak-before-break evaluation are those that have large pipe rupture hardware requirements and those in which the other dynamic effects of postulated pipe breaks are of concern.
These systems are:
RCS Reactor Coolant System RHR Residual Heat Removal System SIS Safety Injection System All of these systems are fabricated using austenitic stainless steel which have high toughness for inclusion in the CPSES-1 WHIPJET program.
The piping systems in the CPSES-1 WHIPJET program are shown in Table 3-1 as well as the piping materials and sizes.
As shown, the pipe diameters range from nominal 10 to 14-inch, and the j
pipes are ASME Class 1 and 2 entirely inside containment.
There are a total of 28 postulated break locations included in this LBB review and a total of 46 pieces of rupture mitigation hardware.
All on-site (field) welds were made using the gas j
tungsten arc welding (GTAW) process.
Approximately 40% of all welds utilized shielded metal arc (SMAW) and/or submerged arc welding (SAW).
Of course, the stainless steel welds in these
{
systems are used in the as-welded condition, except for a few solution annealed spool pieces fabricated in the shop.
3-1 i
f
The Comanche Peak Project uses the following nomenclature for piping line designation:
Line Number 10-RC-1-021-2501R-1
-- AGME Pipe Class Nominal Piping Size (e.g.,
10-inch diameter)
-- Pipe Category
- Line Sequence Number Pipe System Abbreviation Designates CPSES-1 Piping (e.g., SI(S), RH(R), RC(S)]
Table 3-2 presents the WHIPJET Program piping by line number and lists the loop number, piping dimensions, material type, operating conditions, insulation type, and insulation thickness.
The Comanche Peak Project uses the following nomenclature for pipe whip restraint designation:
Restraint Number I
RC-1-021-903-C47W LJ Pipe System Abbreviation -
[e.g., SI(S), RH(R), RC(S)]
Location & Elevation Restraint Number Designates CPSES-1 Restraint -
I Line Sequence Number Table 3-3 presents the WHIPJET program restraints which will be affected by this LBB analysis.
The table includes the l.
appropriate loop number, line number, restraint number, and type of whip restraint.
3-2 l
Figures a-1 through 3-7 show the piping isometric sketches.
Certain pieces of hardware (hard, bumper, honeycomb, and U-bar restraints) are identified in the sketches.
Portions of the lines (anchor to anchor) are class 2, but the LBB controlling stress locations are within the Class 1 portions.
3.2 SYSTEM DESCRIPTIONS The following is a description of each WHIPJET system as to its primary function and its location within safety related areas.
3.2.1 Reactor Coolant Fvstem (RCS)
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 PWR main reactor coolant loop has been demonstrated by Westinghouse and accepted by the NRC through the limited scope rule change to GDC-4 [2].
The RCS piping evaluated by the CPSES-1 WHIPJET program is the pressurizer surge line.
The piping is directly connected to the 3
primary reactor coolant loops and is part of the reactor coolant system pressure boundary.
3.2.2 Residual Heat Removal System (RHR)
The RHR sytem transfers heat from the reactor coolant system to the primary plant component cooling water system to reduce the reactor coolant system fluid temperature to cold shutdown conditions.
The high energy portion of the piping is located directly adjacent to the connection to the reactor coolant loop and is part of the reactor coolant system pressure boundary.
3-3
3.2.3 Safety Iniection System _(SIS)
The SIS is part of the emergency core cooling system (ECCS).
The short term function of the SIS is the prompt delivery of borated water to the reactor core following a loss-of-coolant accident.
The SIS consists of two parts:
the low pres.sure safety injection system piping between the accumulator tanks and the RCS cold leg, and the high pressure injection system piping between the high head safety injection pumps and the primary loop.
The high energy portion of the piping is located directly adjacent to the reactor coolant loop and is part of the reactor coolant system pressure boundary.
l l
l 3-4
T TABLE 3-1 QUALIFYING PIPING SYSTEMS FOR CPSESc1 WHIPJET ANALYSIS PIPING PIPE BASE METAL BREAI4S - (1)
SYSTEM (S)
SIZE MATERIAL (IN)
RCS/ SIS 10 SA 376, 22 TYPE 316 (2)
RCS/RHR 12 SA 376, 5
TYPE 316 (2)
TYPE 316 TOTAL 29 NOTES:
(1)
Includes both circumferential and longitudinal breaks.
Affected hardware (Table 3-3) includes hard restraints, bumper restraints, U-bar restraints, and honeycomb restraints (2)
Small sections contain Type 304 stainless steel l
l 3-5 L
TABLE 3-2 CPSES-1 PIPING DATA FOR WHIRTET IBB ANALYSIS IOOP NO.
LINE NUMBER OUTSIDE WALL PASE OPERATING INSUIATION #
DIAMETER
'IHICENESS MATEPdAL TEMP PRESS TYPE
'IHICKNESS (IN)
(IN)
(F)
(PSI)
(IN) 1 10-RC-1-021-2501R-1 10.750 1.000 SA376 TYPE 316 557 2299 REFIECTIVE 2.0 1
10-SI-1-179-2501R-1 10.750 1.000 SA376 TYPE 316 120 2299 BARE 0.0 1
10-SI-1-103-2501R-2 10.750 1.000 SA376 TYPE 316 120 650 BARE 0.0 1
10-SI-1-119-0601R-2 10.750 1.000*
SA376 TYPE 316**
120 650 BARE 0.0 1
12-RC-1-007-2501R-1 12.750 1.125 SA376 TYPE 316 617 2235 REFIECTIVE 3.0 1
12-RH-1-001-2501R-1 12.750 1.125 SA376 TYPE 316 350 2235 REFIECTIVE 0.7 1
12-RH-1-901-2501R-1 12.750 1.125 SA376 TYPE 316 350 400 REFIECTIVE 0.7 1
12-RH-1-003-0601R-2 12.750 1.000*
SA376 TYPE 316**
350 400 REFLECTIVE 0.7 2
10-RC-1-037-2501R-1 10.750 1.000 SA376 TYPE 316 557 2299 REFIECTIVE 2.0 Y
2 10-SI-1-180-2501R-1 10.750 1.000 SA376 TYPE 316 120 2299 BARE 0.0 m
2 10-SI-1-104-2501R-2 10.750 1.000 SA376 TYPE 316 120 650 BARE O.0 2
10-SI-1-120-0601R-2 10.750 1.000*
SA376 TYPE 316**
120 650 BARE 0.0 3
10-RC-1-055-2501R-1 10.750 1.000 SA376 TYPE 316 557 2299 REFIECTIVE 2.0 3
10-SI-1-181-2501R-1 10.750 1.000 SA376 TYPE 316 120 2299 BARE 0.0 3
10-SI-1-105-2501R-2 10.750 1.000 SA376 TYPE 316 120 650 BARE 0.0 3
10-SI-1-121-0601R-2 10.750 1.000*
SA376 TYPE 316 120 650 BARE 0.0 4
10-RC-1-078-2501R-1 10.750 1.000 SA376 TYPE 316 557 2299 REFIECIT/E 2.0 4
10-SI-1-182-2501R-1 10.750 1.000 SA376 TYPE 316 120 2299 BARE 0.0 4
10-SI-1-106-2501R-2 10.750 1.000 SA376 TYPE 316 120 650 BARE 0.0 4
10-SI-1-122-0601R-2 10.750 1.000*
SA376 TYPE 316 120 650 BARE 0.0 4
12-RC-1-069-2501R-1 12.750 1.125 SA376 TYPE 316 617 2235 REEIECTIVE 3.0 4
12-RH-1-002-2501R-1 12.750 1.125 SA376 TYPE 316 350 2235 REFIECTIVE 0.7 4
12-RH-1-900-2501R-1 12.750 1.125 SA376 TYPE 316 350 400 REFIECTIVE 0.7 4
12-RH-1-004-0601R-2 12.750 1.000 SA376 TYPE 316**
350 400 REFIECITVE 0.7 4
14-RC-1-135-2501R-1 14.000 1.406 SA376 TYPE 316 653 2235 REFLECTIVE 3.0 NOTE: # Reflective insulation has both an inner and outer stainless steel layer
- Portions of line are Schedule 40S; does not affect IBB calculations
- Small portions of line contains SA312 Type 304 stainless steel
TABLE 3-3 CPSES-1 PIPE BREAKS AND HARDWARE ASSOCIATED WITH WHIPJET LBB ANALYSIS TYPE LOOP NO.
LINE NUMBER HARDWARE 1
10-RC-1-021-2501R-1 RC-1-021-901-C47W HARD 1
10-RC-1-021-2501R-1 RC-1-021-902-C47W HARD 1
10-RC-1-021-2501R-1 RC-1-021-903-C47W U-BAR 1
10-SI-1-179-2501R-1 SI-1-179-901-C47W U-BAR 1
10-SI-1-179-2501R-1 SI-1-179-902-C47W U-BAR 1
10-SI-1-179-2501R-1 SI-1-179-903-Cd7W HARD 1
10-SI-1-179-2501R-1 SI-1-179-904-C47W HARD 1
10-SI-1-103-2501R-2 SI-1-103-901-C47W BUMPER 1
10-SI-1-119-0601R-2 NO ASSOCIATED HARDWARE 1
12-RC-1-007-2501R-1 RC-1-007-902-C47W BUMPER 1
12-RH-1-001-2501R-1 NO ASSOCIATED HARDWARE 1
12-RH-1-901-2501R-1 NO ASSOCIATED HARDWARE 1
12-RH-1-003-0601R-2 NO ASSOCIATED HARDWARE 2
10-RC-1-037-2501R-1 RC-1-037-901-C47W HARD 2
10-RC-1-037-2501R-1 RC-1-037-902-C47W HARD 2
10-RC-1-037-2501R-1 RC-1-037-903-C47W U-BAR 2
10-SI-1-180-2501R-1 SI-1-180-901-C47W U-BAR 2
10-SI-1-180-2501R-1 SI-1-180-903-C47W HARD 2
10-SI-1-180-2501R-1 SI-1-180-904-C47W HARD 2
10-SI-1-104-2501R-2 NO ASSOCIATED HARDWARE 2
10-SI-1-120-0601R-2 NO ASSOCIATED HARDWARE l
3 10-RC-1-055-2501R-1 RC-1-055-901-C47W HARD 3
10-RC-1-055-2501R-1 RC-1-055-902-C47W HARD 3
10-RC-1-055-2501R-1 RC-1-055-903-C47W U-BAR 3
10-SI-1-181-2501R-1 SI-1-181-901-C47W U-BAR 3
10-SI-1-181-2501R-1 SI-1-181-903-C47W HARD 3
10-SI-1-181-2501R-1 SI-1-181-904-C47W HARD 3
10-SI-1-105-2501R-2 NO ASSOCIATED HARDWARE 3
10-SI-1-121-0601R-2 NO ASSOCIATED HARDWARE l
4 10-RC-1-078-2501R-1 RC-1-078-901-C47W HARD 4
10-RC-1-078-2501R-1 RC-1-078-902-C47W HARD 4
10-RC-1-078-2501R-1 RC-1-078-903-C47W U-BAR l
4 10-SI-1-182-2501R-1 SI-1-182-901-C47W U-BAR 4
10-SI-1-182-2501R-1 SI-1-182-903-C47W HARD 4
10-SI-1-182-2501R-1 SI-1-182-904-C47W HARD 4
10-SI-1-106-2501R-2 SI-1-106-901-C47W BUMPER 4
10-SI-1-122-0601R-2 NO ASSOCIATED HARDWARE 3-7 i
TABLE 3-3 (CONT)
LOOP NO.
LINE NUMBER HARDWARE TYPE 4
12-RC-1-069-2501R-1 RC-1-069-901-C47W BUMPER 4
12-RH-1-002-2501R-1 NO ASSOCIATED HARDWARE 4
12-RH-1-900-2501R-1 NO ASSOCIATED HARDWARE 4
12-RH-1-004-0601R-2 NO ASSOCIATED HARDWARE HONEYCOMB 4
14-RC-1-135-2501R-1 RC-1-135-901-C47W 4
14-RC-1-135-2501R-1 RC-1-135-902-C47W HONEYCOMB /
U-BAR 4
14-RC-1-135-2501R-1 RC-1-135-903-C47W BUMPER 4
14-RC-1-135-2501R-1 RC-1-135-904-C47W BUMPER 4
14-RC-1-135-2501R-1 RC-1-135-905-C47W HONEYCOMB /
U-BAR 4
14-RC-1-135-2501R-1 RC-1-135-906-C47W HONEYCOMB /
U-BAR 4
14-RC-1-135-2501R-1 RC-1-135-907-C47W HONEYCOMB 4
14-RC-1-135-2501R-1 RC-1-135-908-C47W HONEYCOMB 4
14-RC-1-135-2501R-1 RC-1-135-909-C47W HONEYCOMB 4
14-RC-1-135-2501R-1 RC-1-135-910-C47W HONEYCOMB l
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Section 4 SCREENING CONSIDERATIONS Portions of the three systems (Reactor Coolant, Residual Heat Removal, and Safety Injection) within the CPSES-1 WHIPJET program were processed through a screening evaluation based on the broad scope rule change to GDC-4 and designed to consider the potential for:
o Anomalous conditions:
excessive vibration, unexpected flow stratification, failure of pipe fittings, failure of equipment supports, and uneven mixing; o
Fluid transients (water hammer) ;
o Stress corrosion cracking; and o
Other considerations:
creep, thermal aging, and erosion / corrosion.
The following sub-sections discuss each of these items in detail for the three systems (10-inch SIS, 12-inch RHR, and 14-inch RCS) analyzed in the WHIPJET program.
4.1 ANOMALOUS CONDITIONS 4.1.1 Flow Stratification Flow stratification is a thermal fluid phenomenon identified in IE Notice 84-87 (1).
It potentially occurs in large diameter fluid systems where the following conditions exist:
(1) the piping system configuration 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 scope of WHIPJET systems.
4-1
Each system within the WHIPJET program was reviewed for its susceptibility to fluid flow stratification.
It is concluded that potential for flow stratification is unlikely in the 10-inch SIS and 12-inch RHR lines analyzed in this program.
Previous industry experience has shown acceptably high flow rates at primary loop branch connections thereby minimizing the potential for stratification in the RHR piping.
The SIS is predominantly lines which experience cooler, high volume fluid flow rates.
While some fluid stratification is expected to occur in the 14-inch surge line, due to the difference in temperature between the reactor coolant loop and the pressurizer, it is estimated that this stratification has no significant effect on the surge line.
In late 1987, the Alabama Power and Light plant Farley - Unit 2, noted an increase in moisture and radioactivity levels inside containment during a restart following a refueling outage.
As indicated in IEN 88-01 [5], inspection found a crack at a weld connecting an elbow and a horizontal spool in a 6-inch SIS line attached to the cold leg of RCS loop B.
Farley-2 is a 3-loop Westinghouse PWR and the cause of the thermal fatigue crack was valve leakage in the bypass pipe around the boron injection tank.
At CPSES-1, a 4-loop Westinghouse PWR, the existing system configuration does not permit a leakage path of charging flow to the SIS or RHR lines analyzed in the WHIpJET program.
Thermal stratification is not expected to be a consideration in the horizontal sections of the WHIPJET program lines, and all lines pass the flow stratification screening criteria.
4.1.2 Induced System Vibration Positive displacement pumps are potentially a source of vibration and fatigue failure.
CPSES-1 systems in the WHIPJET program have no positive displacement pumps in service during normal operation.
General system vibration is addressed in the CPSES-1 start-up testing program.
4-2 l
4.1.3 Pine Fittinos Consideration for LBB Analysis A review was conducted to determine the presence of any cast materials for systems evaluated in WHIPJET.
All fittings (elbows, tees, and branch connections) are wrought stainless steel, SA403 WP316.
Ho cast material is used for fittings within the WHIPJET program high energy systems.
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.
Welded 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 wolds.
Service experience with elbows has been excellent except for a few instances of cracks caused by a defective longitudinal elbow weld.
All of the CPSES-1 lines in the WHIPJET 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 1 and l
NC-2550 for Class 2 seamless materials, thus assuring material soundness.
Satisfactory experience combined with the type of cibows and NDE checks, reduce the probability of any longitudinal cracking of the elbows.
Also, the material properties for the fittings are comparable to those in the wrought pipe itself.
Additionally, each selected line of piping was evaluated to determine limiting stress locations.
The limiting locations where WHIPJET is being applied correspond to those determined to be the highest stress points as determined in Section 5 of this report.
4-3 l
4.1.4 Eauipment Supports Evaluation The potential for the failure of equipment supports resulting in the subsequent failure of attached high energy piping exists if the supports have not been seismically qualified.
Since ASME Section III piping systems terminate at ASME Section II,I Seismic Category I components, seismic qualification is assured.
Furthermore, compliance with the snubber surveillance requirements assures that snubber failure rates are acceptably low.
4.1.5 Uneven Mixina Thermal fatigue due to water temperacure fluctuations caused by slow, continuous mixing of water flows with different temperaturos, has been reported to occur for certain PWR plants.
CFSES-1 WHIPJET systems were reviewed and concluded to be free of causative factors leading to this phenomenon.
4.2 FLUID TRANSIENTS 4.2.1 Introduction The potential for flow transient events was evaluated for CPSES-1 in accordance with applicable site and NEC guidelines.
A review of tne three WHIPJET systems was conducted.
Applicable flow transients, the transient description, and Casign information were identified.
Potential water hammer sources considered for the design of CPSES-1 piping systems were based on industry experience and the observations contained in NUREG-0582 (1), NUREG-0918 (2), NUREG-0927 (R), NUREG-0993 [R),
NUREG/CR-2059 [10), and NUREG/CR-2781 (11).
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Water hammer occurrences caused by,6cean voids, condensation, v
flashing and thermal mixing are contreJ1ed through system design, operating procedures, and oper$ tor training.
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The NUREGs that pertain to fluid transient detection ar#
prevention recognize that water (steam) hammer ovents gre inevitable and provide guidance to elimin'tr or'roduce the f'
af frequency and effects of these everf ts,.;
These NUREds also acknowledge that water (steam) hammer 3s not'as significant a safety issue as previour.ly anticipattqd sincoithe resulting damage from water hammer events has h an limited to piping aiid
,1
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equipment supports.
Table 4-1 is a partial reproduction of >
^<,)
Table 3.2 from NUREG-0927 [8] providing some PWR sydten water hammer causes, design considerations, and preventive measures.
The three CPSES-1 systems under review are listed in Table 4-1.
The discussion below addresses the CPSES-1 provisions for minimizing water (steam) hammer effects and, wheEe applicable, addresses specific details associated with that guidance for minimizing water (steam) hammer.
1 A comprehensive review of NUREG-0582 [6] and plant operating experience provided guidance in identifying CPSES-1 flow transients that could-cause piping syst. ens under, evalbation to a
3 be susceptible to water (steam) hammer eventsfand their associated effects.
Appendix A contains a working hummary of and comments on fluid transients and, as appropriate, included in the CPSES-1 design basis for pipe stress analysis.
Also, a brief summary of the flow transient-related NUREGs is provided in Appendix A.
I i
4.2.2 Provisions for Minimizing Water Haramer Effects t
i Systems under consideration for 1BU at CPSES-1 are not, in general, susceptible to water hammer.
The reactor coolant and c,
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i 4-5
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residual heat removal systems have been specifically designed to preclude water hammer.
Operating, experience at other plants with Westinghouse Systems have verified this design approach.
Westinghouse has conducted a number of investigations into the causes and consequences of water hammer events.
The results of these (dvestigations have been reported to operating plant
/e customero ynd have been reflected in design interface t.
requirements to the balance-of-plant designer for plants under
,ry construction.
This process assures that water hammer events I'
initiated in the secondary systems do not compromise the v
i performance of the Westinghouse-supplied safety-related systems and components.
t' In general, the approaches taken, individually or in I
combination, to address water hammer concerns were to prevent or minimize water hammer effects through system cesign features and operating procedures.
Fotential water hammer sources to be considered were based on industry experience and the concerns presented in various NUREG's.
The following subsections discuss I
in. mere detail the potential water hammer sources that were considered in the design of t'he subject systems and the actions taken to minimize and prevent water hammer effects.
4.2.3 Reactor Coolant System (RO11 There is a very low potential for water hammer in the subcooled whter solid portioPs of RCS fince these areas are designed to preclude void formation.
Relief valve discharge loads ass'ociated with the pressurizer have been specifically identified and analyzed for CPSES-1 (following NUREG-0927 (RJ).
I l
However, relief valve discharge is not included in the scope or l
.the CPSES-1 WHIPJET program.
t i
s t t
4-6
i JV
\\ li 4.2.4 Safety Iniection System (SIS) 4 It is considered tjnlikAly 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 filling and venting assures that the low and high head safety injection system piping remains filled.
In addition, the sad cf water provided by the refueling water storage tank provides a continuous mechanism for assuring 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 RHR 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 RCS discussion.
4.2.5 Residual Heat Removal System (RHR)
A portion of the RHR piping is high energy because it is normally pressurized by the RCS or SIS during normal plant operating conditions.
When RHR is operating, valve closure times and operating procedures minimize the potential for water hammer.
Proper filling and venting will initially assure that air does not become trapped in any part of the RHR during start-up.
Additionally, just prior to RHR initiation, the RHR will be cross-connected with the Chemical and volume control System (CHS).
This action utilizes the pressure head in the CHS to collapse any romaining voids prior to opening the RHR suction valves from the RCS.
When the RHR system is not operating, the normalP; pressurized portions of the system are water solid and are either at a low 4-7 1
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I temperature or subcooled.
RHR 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 (R]).
4.3 STRESS CORROSION CRACKING The pertinent lines in the CPSES-1 WHIPJET scope were reviewed t'[ determine if they were susceptible to stress corrosion c
cracking (SCC).
The review has concluded that based on
'specified chemistry, cleanliness, fabrication, and operating controls, and successful industry PWR operating experience, SCC is not expected to occur.
Figure 4-1 presents an overview of the corrosion-related review conducted for CPSES-1.
The WHIPJET program lines (SIS, RHR, and RCS) are predominantly SA376 Type 316 austenitic stainless steel; small portions of four RHR and SIS lines contain Type 304 stainless steel, as indicated in Table 3-2.
These materials when used in other than the solution annealed condition are susceptible to SCC when exposed to three conditions simultaneously:
high tensile stress, high temperature, and a corrosive environment.
.d Controls of pipe manufacture and fabrication minimized the
/
susceptibility to SCC by limiting both cold work (i.e., strain hardening) and sensitization effects (i.e., chromium-carbide precipitation in the grain boundaries).
Any cold working of the material was controlled by requiring solution annealing during the original manufacturing process and also by requiring forming operations to have bend radii greater than or equal to five pipe a
diameters.
Hot bent piping was always re-solution annealed.
1 All piping in the CPSES-1 WHIPJET program was hot bent.
- Hence, the subject stainless steel piping was furnished in the solution annealed condition prior to welding.
Welding-induced sensitization was controlled by limiting the weld interpass t
4-8 i
temperature (limited to 350 F maximum) and weld heat input.
These controls were evaluated to the criteria of ASTM A708 and were judged to be effective in minimizing weld-induced sensitization.
In summary, the measures taken to minimize material susceptibility to SCC by controlling the material condition follow the recommendations of Regulatory Guide 1.44.
In addition, further controls are used to prevent the simultan-cous occurrence of conditions required for SCC.
Specifically, CPSES-1 maintains a non-corrosive environment both inside and outside the piping r.ystems as the primary method for SCC elimination.
The primary controls are provided through adoption cf proven water quality standards and piping inner (ID) and outer diameter (OD) cleanliness ?equirements.
The reactor coolant chemistry requirements are described in the FSAR and included in this report as Table 4-2.
CPSES-1 water chemistry controls are proven effective in preventing SCC by control of oxygen content to less than 0.1 ppm when operating at temperatures above 250 F and limiting chlorides to less than 0.15 ppm at all times.
The control of oxygen content is accomplished by using hydrazine to limit the initial oxygen content to less than 0.1 ppm during heatup and by a hydrogen overpressure during normal plant operation.
Chloride content is controlled by using strict purity requirements for the procurement of reactor coolant chemical additives.
The CPSES-1 program to monitor the reactor coolant water chemistry assures strict adherance to water chemistry requirements, thereby assuring a corrosive-free environment.
Pipe inner and outer wall cleanliness is controlled in accordance with the recommendations of Regulatory Guides 1.37, 1 38, and 1.39.
Swipe testing of the pipe's outer surface for contaminants provides assurance against externally initiated SCC.
In addition, use of thermal insulation is in accordance i
4-9 1
I
with the recommendations of Regulatory Guide 1.36 in assuring that no contaminants are applied to piping that could potentially initiate SCC.
Industry experience with these materials has been reviewed in NUREG-0679 (12), NUREG-0691 [12), and NUREG-1061 [2].
SCC has neither been reported nor is it expected under the spec,1fied operating conditions.
Based on the strict fluid chemistry, cleanliness, fabrication, and operating controls, and successful plant operating experience, the subject lines are judged to be not susceptible to stress corrosion cracking.
Therefore, all lines pass this screening criteria.
4.4 OTHER CONSIDERATIONS 4.4.1 Creep All operating temperatures are well below one-half of the melting point of stainless steel.
Creep only becomes important at temperatures approaching 800 F.
Therefore, creep is not a concern for any of the CPSES-1 lines.
4.4.2 Thermal Aging Cast stainless steel material is known to lose some of its initial fracture toughness properties when exposed to operating temperatures of approximately 550 F for extended periods of time.
However, there is no cast stainless steel in the CPSES-1 WHIPJET scope, and no concern for thermal aging exists.
4.4.3 Erosion / Corrosion Erosion / corrosion problems as evidenced at Surry-2 and elsewhere are not a problem at CPSES-1 due to the use of high quality stainless steel materials.
4 4-10
]
TABLE 4-1 PWR WATER HAMMER SYSTEM PRIMARY CAUSES PREVENTIVE MEASURES OF WATER HAMMER DESIGN PLANT, OPERATION Reactor Relief Valve Include Relief Coolant Discharge Valve Discharge (PZR)
Loads in Pipe Support and Components Design Basis (3.10)
RHR Voiding Venting (3.3)
Operating Procedures (3.12), Operator Training (3.11)
ECCS/ SIS Voiding Venting (3.3),
Operating Procedures Void Detec-(3.12), Operator tion (3.1)
Training (3.11)
Numbers in parentheses refer to sections of NUREG-0927 (R) providing details of preventive measures.
4-11
TABLE 4-2 (From CPSES-1 FSAR)
REACTOR COOLANT CHEMISTRY SPECIFICATION Electrical conductivity Determined by the concentration of boric acid and alkali present, expected range is < 1 to 40 0
pMhos/cm at 25 0.
Solution pH Determined by the concentration of boric acid and alkali present, expected values range between 4.2 (high boric acid concentration) to 10.5 (low boric acid concentra-0 tion) at 25 C; values will be 5.0 or greater at normal operating temperatures.
Oxygen, maximum (npm)
Oxygen concentration of the reactor coolant is maintained below 0.1 ppm for plant operation 0
above 250 F.
Hydrazine may be used to chemically scavenge oxygen during heatup.
Chloride, maximum (ppm) 0.15 Fluoride, maximum (ppm) 0.15 Hydrogen, cc (STP)/kg H O 25 to 35, Reactor power level 2
above 1 MWt, excluding decay heat during subcritical operation.
Total suspended solids, 1.0 maximum (ppm) pH control agent (Li 0H) 0.3 x 10-4 to 3.2 x 10-4 molal 7
lithiu9).(equivalent to 0.22 to 2.2 ppm Li Boric acid, (ppm B)
Variable from 0 to approximately 4000.
4-12 I
AUSTENIT10 STEEL e RCS e SIS e RHR HYDROXIDES ACCEPTABLE OXYGEN ACCEPTABLE SULFUR (reduced forms)
ACCEPTABLE HALOGENS e alorides e Fluorides ACCEPTABLE TEMPERATURE ACCEPTABLE PASS Figure 4-1 CORROSION REVIEW 4-13
Section 5 MATERIAL PROPERTY DATA 5.1 AUSTENITIC STAINLESS STEEL LINES Tensile true stress-strain data for the applicable CPSES-1 stainless steel material (Type 316, 304, and welds) have been collected.
The stainless steel material data for both base (primarily Type 316, but some isolated Type 304) and welds (gas tungsten arc, shielded metal arc, and submerged arc) were gathered from numerous engineering sources (see Appendix B).
The stainless steel base metal information, derived from previously published data (Appendix B), was required for analytical purposes.
These data, along with the piping certified material test reports (CMTR), show that the CPSES-1 high energy lines in the WHIPJET scope arc made of high grade material, extremely resistant to unstable tearing.
The
?
characteristics of the true stress-strain diagram were used for leak rate calculations; Appendix B shows how the industry data were evaluated to derive appropriate stress-strain properties.
The stainless steel gas tungsten arc weld (GTAW), shielded metal arc weld (SMAW), and submerged arc weld (SAW) metal information i
was also derived from existing industry data.
The weld metal properties were utilized in performing crack stability analyses for the WHIPJET program high energy piping welds.
Appendix C discusses the welding procedures and conditioning of materials used for actual CPSES-1 welding.
Some SMAW and SAW off-site (shop) welds were employed in the RCS 14-inch line, the RHR 12-inch lines, and the SIS 10-inch lines.
All other welds (shop and field) were GTAW; the shop and field welding parameters were similar with the same range of heat inputs and the same filler materials.
l 5-1
5.2 HIGHEST STRESSES AND MINIMUM MATERIAL PROPERTIES For each pipe size in a functional system, LBB evaluation procedures identify the most limiting location (s) which have the least favorable combination of stress and material properties for base metal, weldments, and safe ends.
Therefore, the highest stress locations in the 10-inch SIS, 12-inch RHR, and 14-inch RCS surge lines are identified after reviewing forces and moments for normal plus SSE loads.
Normal moments are determined from the algebraic combination of deadweight and thermal components.
The normal moments are then combined absolutely with SSE loads, with the resulting moment a square-root-sum-of-the-squares (SRSS) of the directional components.
The effect of including and excluding torsional moments was considered and evaluated; the highest stress location remained constant whether or not torsion was included except for the RCS pressurizer surge line.
There, the highest stress node switched from the loop nozzle to the pressurizer nozzle when torsion was excluded.
After considering that the normal load moment (without SSE) at the surge line's loop nozzle safe end is approximately one-half of the normal load at the pressurizer nozzle safe end, the loop nozzle safe end was conservatively selected (because of lower leakage at the loop nozzle) as the j
high stress location.
Both locations meet all flaw size and load margins, but only the loop nozzle results are presented in this report.
In considering the highest stress location coupled with the minimum material properties, the piping node points were not separated by weld types or base metal.
Only the highest stress location throughout the line (anchor to anchor) was determined, and this location for each pertinent pipe size was analyzed for l
all material conditions which exist in that line (base metal, 1
5-2
The forces and moments corresponding to the highest stress locations are listed in Appendix D.
In the 10-inch SIS and 14-inch RCS lines, the high stress locations are in the class 1 portion of the lines.
In the Class 2 section of the RHR lines (Schedule 40S), two node points experience stresses which are as large or larger than the highest stress in the Class 1 portion of the lines due to the reduced thickness.
However, because this Class 2 piping sees low pressure and temperature and is isolable, the high stress location in the Class 1 portion of the RHR line is analyzed in this report.
Both locations in the Class 2 portion of the lines meet all flaw size and load requirements (with appreciably higher margins).
For consideration of longitudinal pipe breaks in which diameter, schedule, and pressure are constant, results will be very similar because axial and bending stresses have little effect on longitudinal cracking.
In accordance with accepted fracture mechanics and LBB methodology, each pipe size and system was checked for longitudinal break leakage and stability.
The localized stress effects from welded attachments were also considered.
In addition to all of the previous WHIPJET analyses, the guidelines developed in NUREG-1061, Vol. 3 (2) do not require localized stresses to be included in the LBB calculations.
Based on this philosophy, local welded attachment stresses were calculated but were not included in the analysis of the CPSES-1 lines.
There are only 12 pipe supports with welded attachments in the lines contained in the CPSES-1 WHIPJET scope, and except for 2 cases the stresses at the welded attachments are low compared to the general (nominal) stress.
Since the CPSES-1 LBB program is concerned with through-wall 5-3
t cracks and large scale local plasticity, any localized stresses from welded attachments would be redistributed resulting in little, if any, effect at the through-wall crack tips.
All stresses and data input will be reconciled when the final stress reconciliation program at CPSES-1 is completed.
No significant changes in the stresses used for the calculation results contained herein are anticipated.
The as-built results will be reconciled to the final analysis.
5-4
Section 6 LEAK DETECTION
6.1 INTRODUCTION
One of the key parameters necessary for substantiating leak-before-break (LBB) is the determination of the lowest detectable leakage inside containment.
This parameter is then multiplied by a margin of 10 for use in the WHIPJET LBB analyses.
The following discussion focuses on CPSES-1 leak detection systems, leak detection capabilities, and operator actions.
The discussion presents an assessment of these areas and substantiates the value selected for the limiting detectable leakage.
More detailed discussions of the systems and specifications can be found in Appendix E.
A diverse number of parameters are monitored to detect leakage of systems inside containment.
The WHIPJET systems are entirely inside containment and are part of the reactor coolant pressure boundary (RCPB) as defined in Section 50.2 of 10CFR50.
6.2 RCPB LEAKAGE CATEGORIES The CPSES-1 technical specifications define the categories of reactor coolant pressure boundary leakage to be either identified, unidentified, or controlled leakage.
6.2.1 Identified Leakace Identified leakage is:
o Leakage (except Controlled Leakage) into closed systems, such as pump seal or valve packing leaks that are captured and conducted to a sump or collecting tank, or 6-1
o Leakaga into the containment atmosphere from sources that are both specifically located and known either not to interface with the operation of leakage detection systems or not to be RCPB leakage, or o
RCS leakage through a steam generator to the Secondary Coolant System.
6.2.2 Unidentified Leakace Unidentified leakage is all leakage which is not identified leakage or controlled leakage.
It is impractical to completely eliminate unidentified leakage, but efforts are made to reduce this leakage to a small background flow rate permitting the leakage detection systems to detect positively and rapidly any small increase in unidentified leakage flow rate.
6.2.3 controlled Leakace Controlled leakage is the seal water flow supplied to the reactor coolant pump seals.
I 6.3 RCPB LEAKAGE LIMITS l
The technical specifications provide limits for identified and l
l unidentified reactor coolant system pressure boundary leakage i
categories.
Leakage through the RCS pressure boundary is limited to the following:
Identified leakage -- A total of 1 gpm primary-to-o secondary leakage through all steam generators not isolated from the RCS and 500 gpd through any one steam generator not isolated from the RCS.
A total of 10 gpm leakage from other sources described above, and o
Unidentified leakage -- A 1 gpm limit.
6-2 1
6.4 LEAK DETECTION SYSTEMS (INSIDE CONTAINMENT)
The leak detection systems employqd within containment are designed to find identified, unidentified, and intersystem leakage.
The leak detection systems are employed to monitor leakages from the reactor coolant and auxiliary Class 1 and 2 component systems into the containment and to provide the means to locate such leakage.
In this way, any abnormal leakage can be diagnosed well before the technical specification limits are reached.
Leak detection systems provide information which permits the plant operators to take immediate corrective action should a leak be evaluated as detrimental te the safety of the plant.
In addition to leak detection systems, RCS inventory balances are also performed to determine the unidentified leakage from the RCS.
CPSES-1 procedures allow for containment entry to determine the source and the quantity of the leakage.
The leak detection system design objectives are developed in accordance with the requirements of 10 CFR Part 50, GDC 30, and NRC Regulatory Guide 1.45.
6.4.1 Identified Leakace Identified leakage that is conducted to the Reactor Coolant Drain Tank (RCDT), to the pressurizer relief tank, and to the containment sumpc is, in general, measured with a flow totalizer in the containment sump pumps discharge header.
6-3
6.4.2 Unidentified Leakace Primary indications of unidentified leakage to the Containment atmosphere are provided by air particulate monitors, radioactive gas monitors, containment sump flow and level monitors, and specific humidity monitors.
In addition, there are other detection methods available to the operator for determination of unidentified leakage such as indication of gross leakage and reactor coolant liquid inventory.
During normal operation, the primary monitors show background levels which are indicative of normal unidentified leakage levels inside the containment.
Typical unidentified leakage rates in PWR plants range from 0.1 to 0.4 gpm.
Variations in airborne reactor coolant corrosion products or specific humidity of the containment atmosphere above the normal level signifies an increase in unidentified leakage rates and signal to the plant operators that corrective action may be required.
Similarly, increases in Containment sump flow signifies an increase in uni
.ntified leakage.
Systems and/or means used for detection of unidentified leakage include:
o Containment Air Particulate Monitor, o
Radioactive Gas Monitor, o
Containment Sump Flow Monitoring, o
Containment Sump Level Monitoring, Specific Humidity Monitors, o
--Condensate Flow Rate Measurement,
--Containment Dewpoint Monitors, o
Containment Temperature Monitors, o
Containment Pressure Monitors, o
Gross Leakage Indications, and o
Liquid Inventory.
6-4
Each pump in the two CPSES-1 containment sumps is provided with a running time indicator which indicates the duration of pump operation in seconds.
This indicator can be used to estimate gross leakage rates and acts as a backup to the-discharge flow monitors.
In addition, an alarm sounds in the Control Room if there is a 3-inch sump level change in 45 minutes or less.
This corresponds to a minimum 1 gpm leak rate to the sump.
Finally, reactor coolant volume is indicative of unidentified system leakages.
Het level changes in the pressurizer and volume control tank are functions of the1 system leakage because the Chemical Volume control System is a closed loop system.
Abnormal makeup requirements are indicative of unidentified system leakage.
6.4.3 Intersystem Leakace Leakage of reactor coolant into secondary and auxiliary systems can occur as a result of equipment defects.
As discussed previously, the principal leakage path for primary coolant into other systems is through the steam generator tubas into the secondary side of the steam generator.
An additional path is into the Component Cooling Water (CCW) system, which assures that sufficient cooling capacity is available for continued operation of safety-related equipment during normal and accident conditions.
A summary of the leak detection system sensitivity and response times are tabulated in Appendix E, Table E-1.
6.4.4 Indications in the Control Room Identified leakage is monitored and/or alarmed in the Control Room as follows:
6-5
o RCDT level, pressure, and flow indications, o
Excessive radiation levels in the steam generator blowdown and condensor off-gas systems, o
Reactor head flange leak-off temperature, Pressurizer safety and relief valve discharge piping o
temperature, o
CCW surge tank level, and o
CCW radiation monitors.
Unidentified leakage is monitored and/or alarmed in the Control Room as follows:
o Containment Air Particulate Monitor, o
Radioactive Gas Monitor, o
Containment Sump Flow Monitoring (recorded and totalized, but not alarmed),
o Containment Sump Level Monitoring (alarmed) o Specific Humidity Monitors, o
Condensate Flow Rate Measurement, o
Containment Dewpoint Monitors, a
o containment Temperature Monitors, and o
Containment Pressure Monitors.
During refueling shutdown periods, water is introduced to each containment sump to confirm the 1 gpm leak alarm switch settings l
and ala7cm annunciation.
6.5 LEAKAGE DIAGNOSIS The leak detection systems are monitored at specific intervals.
The containment sump inventory and discharge are monitored on a shift-by-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 RCS water inventory balance is normally performed every 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br /> to satisfy l
6-6 l
l l
l l
l technical specifications.
The containment atmosphere gaseous and particulate radioactivity is also monitored at least once per shift.
In addition, plant operator logs are reviewed for trending.
6.5.1 Examples Examples of leakage diagnosis techniques include:
Leakage occurring from the RCPB is identified by a o
simultaneous rise in condensate and radioactivity monitor indications, Dewpoint temperature recordings assist operators in o
locating leakage points because of the various locations of the containment dew cells, and The increased frequency of sump pump operation is also o
an indication of loakage as is the 1 gpm leak alarm in the control Room.
6.5.2 Adecuacy of Leak Detection Svqt;ggg The leak detection systems inside containment are capable of detecting leakage as low as 0.1 gpm using the air particulate monitor and as low as 1 gpm using the radioactive gas monitor and the sump level alarm.
Recently, a 3-loop Westinghouse PWR experienced a leak inside containment (1) which was determined to be 0.7 gpm of unidentified leakage.
The leakage was detected in a timely manner by an increase in containment radioactivity and moisture levels.
6.6 ACTIONS Actions depend on the evaluation of the containment parameters.
Actions are based on Table E-2 and the calculated value for RCS unidentified leakage.
6-7
6.6.1 Analyze / Evaluate Containment Parameters The following actions are performed whenever the conditions of the previous section are exceeded:
Analyze information from control room instrumentation o
to determine the nature of the leak, Upon det".uination, notify the operating supervisor, o
and o
Evaluate all information available from control room instrumentation as well as locally mounted instruments to identify the leak and determine its location.
6.6.2 Perform An RCS Inventory Balance An RCS inventory balance is performed when the conditions described in the leakage diagnosis example discussed in Section 6.5.1 are exceeded as specified by the abnormal operating procedure.
It is a controlled test to determina the magnitude of the unidentified RCS leakage.
It is noted that in addition to being performed when these conditions are exceeded, this inventory balance is normally 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.
6.6.3 Containment Entry /Insnections Containment entry / inspection is required by operating procedures when unidentified RCS leakage is greater than 1.0 gpm.
For leakage s 1.0 gpm, action is decided by the Nuclear Shift Supervisor (NSS).
When containment inspection is required, the source of leakage is determined by visual inspection and measured.
When the source is specifically located and quantified, the leakage is then considered identified.
6-8
L The following technical specification limits for RCS leakage apply:
o Pressure Boundary Leakage from a RCS Component Body, Pipe Wall, or Vessel Wall - None, o
RCS Unidentified Leakage $ 1 gpm, o
Reactor-to-secondary leakage through all steam generator not isolated from the RCS $ 1 gpm (total),
o Leakage through any one steam generator not isolated from the RCS $ 500 gallons per day, o
RCS Identified Leakage $ 10 gpm, and o
RCS Controlled Leakage (2235120 psig) $ 40 gpm.
Leakage of 0.5 gpm per nominal inch of valve size up to o
a maximum of 5 gpm at an RCS pressure of 2235 1 20 psig from an RCS pressure isolation valve.
6.6.4 Limitino Conditions for coeration r
The following RCS leak detection systems shall be operable:
o The Containment Atmosphere Particulate Radioactivity Monitoring System, o
The Containment Sump Level and Flow Monitoring System, and o
Either the Containment Air Cooler Condensate Flow Rate or a Containment Atmosphere Gaseous Radioactivity Monitoring System.
With only two of the above required leak detection systems operable, plant operation may continue for up to 30 days provided samples of the containment atmosphere are obtained and analyzed at least once per 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> when the required Gaseous or Particulate Radioactivity Monitoring System is inoperable.
l otherwise, the plant will be in at least 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 />.
6-9
Depending on the location and magnitude of the leak appropriate repair actions will be taken.
If the leakage is identified as a cracked pipe in the RCS pressure boundary (on the main coolant side of the second check valve; i.e.,
class 1 piping) the technical specifications dictate shutdown and repair (1A).
If the source of leakage is due to a cracked pipe in the non-pressure boundary RCS (Class 2), that system will be isolated (if possible), and appropriate action will be taken.
In accordance with technical specifications, if the calculated unidentified leakage is greater than 1 gpm or identified leakage l
is greater than 10 gpm, 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 />.
With any RCS pressure isolation valve leakage greater than the limit specified in Section 6.6.3, the high pressure portion of the affected system must be isolated from the low pressure portion within 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> by use of at least two closed manual or deactivated automatic valves, or be in at least HOT STANDBY i
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 />.
Evaluations will consider the source and magnitude of the leak, rates of change of detection variables and whether or not shutdown is required.
The evaluations will be used to determine shutdown rates and condition.
Action taken will be recorded in a written log, i
6-10
After shutdown, corrective action will be taken before operation is resumed.
6.7 CONCLUSION
S If abnormal leakage is detected, CPSES-1 operating procedure requires evaluation of the containment parameters to determine the source and magnitude of the leak.
In addition, the procedure requires an RCS inventory balance which is a controlled test performed over a 3-hour period 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.
Based upon the previous discussion, the detectable leak rate limit is conservatively selected at 1.0 gpm.
This value when multiplied by the NRC margin of 10 establishes 10 gpm as the level for determining crack size under normal operating loads.
Section 7 on leak rate calculations uses the 10 gpm leak rate level as an analytical basis for inside containment WHIPJET LBB evaluations.
l I
l i
6-11
r___________________
Section 7 LEAK RATE CALCULATION
SUMMARY
Th5 CPSES-1 WHIPJET program used the EPRI PICEP (EIpe Grack Evaluation Erogr.am) (15) computer code to analyze pipe leakage.
The PICEP code calculates the crack-opening area, the limit load critical crack length, and flow rate for through-wall cracks-in pipes.
The cracks can be oriented either axially (longitudinally) or circumferential1y.
Crack opening area is calculated using elastic-plastic estimation methods fully explained and validated in Reference 11 Leak rate calculations employ two-phase flow based on homogeneous non-equilibrium critical flow theory.
The accuracy of PICEP is judged to be 125%.
Each selected CPSES-1 system in the CPSES-1 WHIPJET program (see Table 3-1) was analyzed using the PICEP computer progran to determine leak rate as a function of crack size for the highest stress locations.
For the estimation of leakage, the normal operating loads (i.e., deudweight, thermal expansion, and pressure) were combined based on the algebraic sum of individual values.
The piping thermal load cases were examined in order to determine the case which represents the most likely thermal expansion loading of the piping during normal steady-state operation.
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 (11]
indicate that the crack-opening area in the case of circumferen-tial welds is governed better by the bulk stress-strain properties of the base metal.
Therefore, all leak rate calcula-tions used base metal properties for estimating crack leakage; this was also true for the longitudinal crack analyses since the 7-2
size of a crack extends appreciably outside the small longitudinal indication of the weld itself.
Additionally, 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 with regard to leakage calculations (see Appendix B).
Results from the leakage rate calculations are presented in Appendix F as a series of leak rate versus crack size plots.
Based upon the conservative limiting detectable leak rate of 1.0 gpm inside containment plus the leak rate margin of 10, a leak rate of 10 gpm was used to determine a detectable leak rate (with margin) crack size for crack stability analysis.
Table 7-1 presents the leakage size cracks for each system under analysis.
The subsequent sections of this report use thece results for evaluating crack stability.
t t
l i
7-2 1
TABLE 7-1 LEAK RATE RESULTS FOR STAINLESS STEEL LINES INSIDE CONTAINMENT LINE SIZE 10 GPM CRACK SIZE (1)
(inches diameter)
(inches)
CIRCUMFERENTIAL BREAKS SIS ic 4.35 RHR 12 6.66 RCS 14 5.94 LONGITUDINAL BREAKS SIS 10 5.29 i
.6.93 NOTE:
(1)
Crack size corresponding to a 10 gpm leak rate under normal operating londs i
l 7-3
Section 8 CRITICAL CRACK SIZE, CRACK SIZE MARGIN, AND LOAD MARGIN CALCULATIONS (CIRCUMFERENTIAL CRACKS)
A generic approach similar to that used as the basis for IWB-3640 in the ASME Code Section XI was employed to evaluate crack stability (12).
Although the PICEP computer code used for the leak rate calculation can provide a modified fracture mechanics / limit load critical crack length, the CPSES-1 WHIPJET program used a modified limit load analysis methodology called a master curve approach.
Appendix G provides both a detailed discussion of the methodology used (cons *ruction and use of the master curves) and the actual master curves developed in the analyses; pertinent parameters used to generate the master curves are identified in the Appendix G figures.
For each postulated through-wall crack, the master curve analysis, using normal plus seismic loads, was used to determine the critical crack size.
Table 8-1 presents crack.1ize margins and demonstrates that there is a margin of 2 between the leakage and critical crack sizes in all three systems.
The e ;a manter curve methodology demonstrates that leakage size cracka will not experience unstable crack growth if 1.4 tires the normal plus seismic loads are applied.
The critical crack sizes from the master curves for the three piping systems reviewed are all greater than the leakage size flaw, and thus the required margin on load is met.
Table 8-2 presents load margin results.
8-1
r.
.t i <+
TABLE 8-1 l
CIRCUMFERENTIAL CRACK STABILITY EVALUATION -
FLAW MARGIN EVALUATIONS (NORMAL + SSE)
REQUIRED MARGIN ON FLAW SIZE,(L /L10gpm) 2 2 C
+
LINE/ TYPE L
SI L
/ MARGIN ON 10gpm C
' FLAW SIZE (in)
(psi)
(in)
SIS / BASE 4.35 17418 12.65.
2.91 (10-inch)
SIS /SAW 4.35 35065 8.78 2.02 (10-inch) r SIS /SMAW 4.35 31608 9.66 4
2.22 (10-inch)
/
RHR/ BASE 6.66 13338 16.89 2.54 (12-inch) kHR/SAW 6.66 21994 14.68 2.20 j (12-inch)
.c
/
s RHR/SMAW 6.66 19925 15.42
'2.32 (12-inch)
RCS/ BASE 5.94 12529 18.83 3.17 t
(14-inch)
RCS/SAW 5.94 20285 16.58 2,79; l
(14-inch) f RCS/SMAW 5.94 18434 17.34
':t 92 (14-inch) l d
NOTES:
i
+
Lyg gp,is the 10 gpm leakagw size crack 1
l Calculated stress index for critical crack size calculation L, critical crack size obtained from master curve e
b 8-2 L
3 TABLE 8-2 CIRCUMFERENTIAL CRACK STABILITY EVALUATION -
LOAD HARGIN EVALUATIONS (1.4 X (NORMAL + SSE)]
'iC PASS EXCESSIVE LOAD ANALYSI ':
(L /L10gpm) 1 C
LINE/ TYPE L
SI L
ACTUAL 10gpm C
L /L10gpm (in)
(psi)
(in)
C SIS / BASE 4.35 24385 10.36 2.38 (10-inch)
IIS/SAW 4.35 49091 5.40 1.24 (10-inch)
SIS /SMAW 4.35 44251 6.54 1.50 (10-lach)
RHR/ BASE 6.66 18673 14.57 2.19 (12-inch)
RHR/SAW 6.66 30791 11.78 1.77 (12-inch)
RHR/SMAW 6.66 27896 12.69 1.91 (12-inch)
RCS/ BASE 5.94 17541 16.32 2.75 (14-inch)
RCS/SAW 5.94 28399 13.51 2.27 (14-inch)
RCS/S. MAW 5.94 25807 14.45 2.43 (14-inch' NOTES:
+
L is the 10 gpm leakage size crack 10 gpm Calculat6d stress index for critical crack size calculation L,
critical crack size obtained from master curve C
e 0-3
!I 3
. - -.. ~ -
4 Section 9 CRITICAL CRACK SIZE AND CRACK SIZE MARGIN CALCULATIONS j '
(LONGITUDINAL CRACKS)
For each of the systems analyzed at CPSES-1, a longitudinal (axial) crack evaluation was conducted to assure that circumferential cracks were the limiting cases.
For consideration of longitudinal cracks, the pressure is assumed to be constant.
Only pressure loads are important in assessing crack opening area and stability.
As shown in Table 9-1, the margins on crack size were evaluated by ratioing the critical crack size calculated using the empirical results of Eiber [18) (contained in the PICEP computer code [15])
to the 10 gpm leakago size
)
crack.
In Table 9-2, the margins for the ratio of the overload critical size crack to the 10 gpm crack are given.
Note that these margins are well above the minimum required values of 2.0 f
,for Table 9-1 and 1.0 for Table 9-2.
j t
/
9-1
TABLE 9 -l LONGITUDINAL CRACK STABILITY EVALUATION
'/.
(NORMAL + SSE) g LI?!E PIPE SIZE RATIO OF CRITICAL-SIZE (INCHES)
TO 10 GPM LEAK RATE SIZE (1)
SIS 10 3.07 e
r
'}
9 l
l l
/
NOTES:
(1)
Critical size is based upon empirical failure results (18_]
for pressure loads.
Minimum required ratio is 2.0.
l j
9-2
TABLE 9-2 LONGITUDINAL CRACK STABILITY EVALUATION
[1.4 x (NORMAL + SSE)]
LINE
-PIPE SIZE RATIO OF CRITICAL SIZE (INCHES)
TO 10 GPM LEAK RATE' SIZE (1)
SIS 10 2.12 RHR 12 2.38 RCS 14 2.50 NOTES:
(1)
Critical size is based upon empirical failure results (18]
for pressure loads.
Minimum required ratio is 1.0.
9-3
Section 10 REFERENCES 1.
WHIPJET Procram Final Report, Beaver Valley Power Station -
Unit Number 2, Prepared for Duquesne Light Company by Robert L.
Cloud & Associates, Inc., January 30, 1987.
2.
Report of the U.
S.
Nuclear Reaulatory Commission Pipina Review Committee; Evaluation of Potential Pine Breaks, NUREG-1061, Vol.
3, November 1984.
3.
Nuclear Regulatory Commission, "Modification of General Design Criterion 4 Requirements for Protection Against Dynamic Effects of Postulated Pipe Ruptures," Federal Reaister, Volume 52, Number 207, October 1987.
4.
Pipinu Thermal Deflection Induced by Stratified Flow, IE Information Notice No. 84-87, US NRC, Washington, D.C.,
December 1984.
5.
Safety Iniection Pine Failure, IE Information Notice No.
88-01, US NRC, Washington, D.C.,
January 1988.
6.
Water Hammer in Nuclear Power Plants, NUREG-0582, US NRC, Washington, D.C.,
July 1979.
7.
Prevention and Mitiaation of Steam Generator Water Hammer Events in PWR Plants, NUREG-0918, US NRC, Washington, D.C.,
l November 1982.
8.
Evaluation of Water Hammer Occurrence in Nuclear Power l
Plants; Technical Findinas Relevant to Unresolved Safety Issue A-1, NUREG-0927, Revision 1, US NRC, Washington, D.C.,
l March 1984.
l l
9.
Reaulatory Analysis for USI A-1, "Water Hammer", NUREG-0993, Revision 1, US NRC, Washington, D.C.,
March 1984.
- 10. Compilation of Data Concernina Known and Susoected Water Hammer Events in Nuclear Power Plants; CY 1969 - May 1981, NUREG/CR-2059, US NRC, Washington, D.C.,
May 1982.
l 1
- 11. Evaluation of Water Hammer Events in Licht Water Reactor Plants, NUREG/CR-2781, US NRC, Washington, D.C.,
July 1982.
- 12. Eloe Crackina Experience in Licht-Water Reactors, NUREG-0679, US NRC, Washington, D.C.,
August 1980.
- 13. Investication and Evaluation of Crackina Incidents in Pinina in Pressurized Water Reactors, NUREG-0691, US NRC, Washington, D.C.,
September 1980.
l l
10-1 1
- 14. Comanche Peak Steam Electric Station - Unit 1 Technical Specifications, Revision B.
o 15.
D.
M. Norris et al.,
PICEP:
Pine Crack Evaluation Procram, EPRI NP-3596-SR (Revision 1), Special Report, December 1987.
16.
B.
R.
Ganta et al., Analysis of Cracked Pine Weldments, EPRI NP-5057, February 1987.
17.
D. M. Norris, et al.,
Evaluation of Flaws in Austenitic Steel Ploina, EPRI NP-4690-SR, Special Report, July 1986.
- 18. Eiber, R. J.
et Investication of the Initiation and Extent of Ductile Pioe Ruoture, Battelle Columbus Laboratories Report SMI-1908, June 1971.
10-2
APPENDIX A FLUID TRANSIENTS AND PIPE CRACKING INCIDENTS A.1 FLUID TRANSIENTS On the basis of reactor operating experience, NCREG-0582 [1]
shows the most serious water hammer concerns were found to be:
o Slug impact due to rapid condensation in certain PWR steam generators, o
Pump start-up with inadvertently voided lines in ECCS and RHR systems of BWRs, and o
Main feedwater line transients caused by flow control valves in BWRs and PWRs.
However, specific systems of importance are presented in tabular form along with the applicable water hammer events which must be addressed.
Note, none of the serious water hammer concerns occur in the WHIPJET program lines.
The following classification of water hammer problems from NUREG-0582, was adapted:
1)
Pump Start-up with Inadvertently Voided Discharge Lines.
In water systems designed for operation with full discharge lines, inadvertent voiding of the lines due to air entrapment or draining may result in excessive dynamic loads following pump start-up and should be prevented.
2)
Expected Flow Discharge Into Initially Empty Lines.
Discharge lines in water systems which are normally empty and discharge lines from various pressure relief valves should be designed to withstand the expected dynamic loads.
3)
Valve Opening, closing and Instability.
Rapid valve j
opening and closing in both water and steam systems and l
instability of control valves in water systems may cause excessive dynamic loads.
l l
4)
Check Valve closure and sudden Delayed Opening.
Normal check valve closure following pump stopping is not expected to result in large dynamic loads, but should be considered in the system design.
For certain check valves that perform a safaty function, the valves and associated piping should be designed to withstand the l
large dynamic loads resulting from a postulated rupture A-1
upstream of the valve.
The sudden opening of a stuck check valve after pump start-up can also produce damaging pressure pulses.
5)
Water Entrainment in Steam Lines.
Water slugs driven by steam may cause excessive dynamic loads while being swept through bends in the lines and from impact on tees ur closed or partially closed valves.
6)
Transient Cavitation (column separation).
The subsequent collapse of voids formed in water systems by low pressure transients resulting from pump stopping or seizure and change in valve setting or check valve closure may produce excessive dynamic loads.
7)
Steam Bubble Collapse and Mixing of Subcooled Water and Steam from Interconnected Systems.
Damaging water hammer may result from the collapse of steam bubbles in water systems due to pressurization and condensation l
following pump start-up cr valve opening and from the mixing of steam and subcooled water from interconnected i
systems.
8)
Slug Impact Due to Rapid Condensation.
The impact of water slugs, formed and driven by forces resulting from rapid condensation of steam on subcooled water in the feedwater rings and adjacent piping, has been identified as the cause of damaging water hammer in certain PWR steam generators with top feed.
9)
Pump Start-up, Stopping and Seizure with Full Lines.
Dynamic loads resulting from pump start-up with full lines are expected to be relatively small, but should be considered in system design.
Pump stopping in some systems may produce column separation.
Postulated pump l
seizure can result in large dynamic loads and may cause i
column separation.
Listings of the NUREG-0582 fluid transient applicability with l
respect to CPSES-1 lines analyzed in the WHIPJET program are l
provided in Tables A-1 and A-2.
Table A-1 provides CPSES-1 operating procedures which are followed to assure these fluid transients are minimized and/or eliminated.
l l
A.2 WATER HAMMER OCCURRENCE IN NUCLEAR POWER PLANTS l
Total elimination of water hammer occurrence is not feasible, l
due to the possible coexistence of steam, water, and voids in various nuclear plant systems.
Experience found during the NUREG-0927 [2] review shows that design inadequacies and operator-or maintenance-related actions have contributed about equally to initiating water hammer occurrence 3.
A-2 L
The major conclusions reached are that the frequency and severity of water hammer occurrence can be and to some extent have been significantly reduced through design features such as keep-full systems, vacuum breakers, J-tubes, void detection systems and improved venting procedures, and increased operator awareness and training; and that the current potential for significant damage as a result of water hammer events is less than it was in the early and mid 1970's.
Total elimination of water (steam) hammers is not feasible, due to various inherent features of plant design and operation.
Therefore, currently accepted design practices for including anticipated water (steam) hammers as occasional mechanical loads in the design basis of piping and their support systems should be maintained.
No water hammer incidents have resulted in the loss of containment integrity or the release of radioactivity outside of the plant.
The frequency and severity of events in PWR systems are low, with the exception of steam generator water hammer and feedwater-control-valve-induced water hammer.
A.3 REGULATORY ANALYSIS FOR USI A-1, "WATER HAMMER" NUREG-0993 [3] is the NRC staff's regulatory analysis dealing with the resolution of the Unresolved Safety Issue (USI)
A-1, Water Hammer.
NUREG-0993 contains the value-impact analysis of l
this issue, public comments received, and staff response,.or action taken, in response to those comments.
The USI A-1 deals with safety concerns related to water hammer occurrence in nuclear power plants.
Reported water hammer damage has been principally confined to pipe hangers, snubber systems, and equipment-mounting structures.
Furthermore, l
approximately half of the water hammer events reported since 1969 occurred in the plant preoperational phase or first year of commercial operation.
I A risk assessment study was performed to assess the significance of risk from water hamm3r occurrence with respect to overall I
plant risk.
The results of these risk assessments, where both calculated public doses and core melt frequencies are shown, l
conclude that water hammer effects on PRR risk are negligible, l
I A.4 PWR WATER HAMMER As part of the NUREG/CR-2059 (4] review, five Category I*
events were identified in the RCS.
All involved flow-into-l 1
Category I denotes known and suspected water hammer events.
A-3
voided lines and occurred in the pressurizer discharge line when the pressurizer relief valve opened and a slug of water was propelled into an essentially voided line.
In four of the accidents, the relief valve was installed with an upstream water seal to prevent valve seat erosion.
In the remaining incident, the valve stuck open due to boric acid buildup.
Of these five Category I events, two occurred in Westinghouse PWRs.
Both occurred within one year after the plant's' commercial operation date and were design related.
The damaged items consisted of:
o Rupture disc, o
o o
Support, and o
Drain line.
In the RHR system compilation, only one event, identified as steam hammer, involved a pump start with an incorrect valve line-up.
However, this was not a Westinghouse PWR.
In the SIS, four events were identified.
Three events, identified as flow-into-voided-line, involved the safety injection lines not being water filled.
One event, identified as steam-bubble collapse, involved pressure reduction in a safety injection line during testing.
Of the four SIS Category I events, three occurred in Westinghouse PWRs.
All occurred after the first year of the plant's commercial operation date and were either design or procedure related.
The damaged items consisted of:
o Supports, and o
Valve bolts.
The majority of the Category I events identified occurred in feedwater, condensate, or steam systems.
This review provides support for the claim that water hammer effects have been minimized in the WHIPJET lines.
A.5 EVALUATION OF WATER HAMMER EVENTS IN LWR PLANTS The evaluations of damage from and the safety implications of water hammer events indicate that water hammer is not as severe a problem as had been originally believed.
NUREG/CR-2781 [5]
shows that the frequency and severity of water hammer events in PWRs is low and none of the reported events disabled a safety system or train, had an adverse safety effect on the plant, or placed the plant in a faulted or emergency condition.
A-4
Recommendations to mitigate or prevent water hammer events include:
o Line Void Detection, Filling and Venting, o
Operator Training, o
Anticipated Loads, o
operating and Maintenance Procedures, and o
Line sloping.
Water hammer has not been of safety significance in RCS pressurizer systems because no water hammers have occurred in the system.
The safety significance of the pressurizer and the relief valve transients that have occurred in the system are moderate.
The RHR system is generically susceptible to the types of water hammer events that occur in normally idle, pumped water systems, such as flow into voided lines in the pump discharge lines and steam-bubble collapse in the high-temperature pump suction lines during the start of the shutdown cooling.
The PWR RHR system is less prone to voiding than similar BWR systems.
The level of the water source in the reactor water storage tank is above the pump discharge line and serves as a keep-full system.
The safety significance of water hammer in the PWR RHR systems is low because the one event that occurred in the system only resulted in support damage.
In general, the ECCS Safety Injection System consists of two subsystems, the passive accumulator injection and the active safety injection.
The safety significance of water hammer in the acctunulator injection subsystem is low.
The one event in this subsystem resulted from testing that is only performed during shutdown and caused only minor support damage.
The safety significance of water hammer in the active safety injection subsystem is high.
The three events in this system caused considerable support damage.
However, no event caused any pipe leakage.
As was predicted, the supports absorbed the brunt of the increased piping stresses.
A.6 PIPE CRACKING EXPERIENCE IN LWRs NUREG-0679 [5] is an evaluation of pipe cracking experience in l
LWRs.
One area of study, Intergranular Stress Corrosion I
cracking (IGSCC) in piping, found no occurrences of IGSCC have been reported to date in piping for PWR primary coolant systems.
Further, IGSCC is not expected to occur in such piping l
because the conditions required for stress corrosion cracking l
are not present.
In particular, dissolved oxygen in the coolant is controlled to levels low enough to preclude IGSCC through the use of hydrazine additions, a hydrogen overpressure, or some other method.
i I
A-5
Other impurities that might cause stress corrosion cracking, such as halides and caustics, are also rigidly controlled.
Only for brief pc.lods when the reactor is shutdown and the coolant is exposed to air and during the subsequent start-up are conditions considered even marginally capable of producing stress corrosion cracking in the primary system of PWRs.
Although not widespread, IGSCC has occurred in PWR pipi'ng in system other than the primary coolant piping systems.
These incidents have generally occurred in the heat affected zone (HAZ) of welds in austenitic stainless steel pipe, notably Type 304, where the piping was sensitized by the welding process.
IGSCC has also been reported in base material that was otherwise sensitized.
Water with a high oxygen level as well as corrodent contaminants of chlorides, fluorides, caustics, and sulfur compounds are suspected because the cracking has occurred only in piping normally or intermittently containing stagnant water where contaminants may concentrate and become more aggressive.
All the identified PWR cracks occurred in pipes containing borated (H BO ) water with dissolved oxygen at air-saturated levels 3
4 and low measured chloride levels.
A.7 CRACKING INCIDENTS IN PIPING IN PRESSURIZED WATER REACTORS NUREG-0691 (2) summarizes an investigation of known cracking incidents in pressurized water reactor plants.
Cracking from several mechanisms, including intergranular stress corrosion cracking (IGSCC), thermal fatigue, vibrational fatigue, and pipe failures due to dynamic or water hammer loading, have been observed in various PWR piping systems.
The PWR IGSCC primarily has occurred in stainless steel piping in the ASME Class 2 and 3 systems that have low or stagnant flow conditions.
Thermal fatigue cracks have been found in ferritic feedwater piping.
Mechanical loads (vibrational) fatigue typically occurs at socket welds in relatively small-diameter draining, instrument, and low-volume flow lines.
Isolated failure incidents in a few lines have been attributed to water hammer or dynamic loadings.
In LWR plants, a primary cause of cracking in piping made of austenitic stainless steel has been IGSCC, a condition typified by grain boundary cracking of sensitized stainless steel caused by synergistic effects of high tensile stresses, a susceptible material, and a corrosive environment.
All three ingredients must be present for IGSCC to take place.
The high tensile stresses are largely due to the residual stresses from welding, coupled with the applied loads to the A-6
piping.
As such, they can vary with welding technique, piping diameter and thickness, and subsequent thermal treatment of the weldment.
However, in more cases, a significant difference in the residual stress patterns after welding between BWR and PWR piping is not expected.
Residual stresses in piping can achieve magnitudes that are at or near yield.
In stainless steels, sensitization renders the material susceptible to IGSCC.
In this condition, the materia 1' adjacent to grain boundaries has been depleted of chromium due to the precipitation of chromium carbides in the grain boundaries.
Sensitization may result from heat treatment, welding, or any other process that keeps the austenitic stainless steel in the necessary temperature range.
The primary cause of sensitization in the LWR piping has been heating during welding.
In this case, sensitization occurs primarily in an area of the pipe next to the weld commonly referred to as the heat-affected zone (HAZ) of the weld, which coincides with the zone of peak residual stresses in many cases.
The susceptibility to sensitization increases with the carbon level of the stainless steel.
For sensitization from welding to lead to IGSCC under LWR conditions, it appears that the carbon level of the austenitic stainless steel pipe must be at the higher end of the acceptable range.
The role of oxygen is to shift the corrosion potential to a value more active for grain boundary corrosion of sensitized stainless steel.
Other corrodents that can promote IGSCC in sensitized austenitic stainless steels includes fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur such as sulfides, sulfites, or thionates.
These may be present as impurities in the coolant, from welding fluxes or surface contamination during fabrication, or from carrosion of I
l the steel itself, obviously, their presence causes the greatest effects in stagnant lines, where they are not routinely removed by demineralization and can build up or become absorbed in the passivating oxide films on the stainless steel.
If the local concentration of chloride ions is high enough, transgranular stress corrosion cracking (TGSCC) could also develop in stainless steel that is not sensitized.
Since chloride ions are known to be strongly chemisorbed in metal oxide or hydroxide films, they may be present locally in l
sufficient quantity to cause TGSCC even though their concentration in the bulk coolant is less than 1 ppm.
The tendency for TGSCC is increased by low pH, increasing temperature, and the presence of dissolved oxygen in the coolant.
As with IGSCC, chloride-induced TGSCC would most likely occur in stagnant lines.
I A-7
Again, although not widespread, IGSCC and TGSCC have occurred in PWR piping in systems other than the primary coolant system from which oxygen is not rigidly excluded.
To date, stress corrosion cracks have been reported in 8 PWR units in the United States.
Of these 8, one occurred in the SIS, where borated water was discovered leaking from a crack in an 8-inch Schedule 10 Type 304 stainless steel pipe.
Oxygen was present where the piping was not adequately vented.
Low but definite levels of chloride were present in the 12% boric acid content solution.
The affected pipe sections were in a stagnant flow condition from the time of construction; thus, impurities were never diluted or washed out.
The remaining 7 instances involved other low-flow systems in the plant, nor the RCS, RHR, or SIS.
A.8 REFERENCES 1.
Water Hammer in Nuclear Power Plants, NUREG-0582, US NRC, Washington, D.C.,
July 1979.
2.
Evaluation of Water Hammer Occurrence in Nuclear Power Plants; Technical Findinas Relevant to Unresolved Safety Jssue A-1, NUREG-0927, Revision 1, US NRC, Washington, D.C.,
March 1984.
3.
Reculatory Analysis for USI A-1, "Water Hammer",,
NUREG-0933, Revision 1, US NRC, Washington, D.C., March 1984.
4.
Compilation of Data Concernina Known and Suspected Water Hammer Events in Nuclear Power Plants; CY 1969 - May 1981, NUREG/CR-2059, US NRC, Washington, D.C.,
May 1982.
5.
Evaluation of Water Hammer Events in Licht Water Reactor Plants, NUREG/CR-2781, US NRC, Washington, D.C.,
July 1982.
6.
Pipe Crackina Exoerience in Licht-Water Reac? ors, NUREG-0679, US NRC, Washington, D.C.,
August
.380.
7.
Investication and Evaluation of Crackina Incidents in Pipina in Pressurized Water Reactors, NUREG-0691, US NRC, Washington, D.C.,
September 1980.
A-8
TABLE A-1 CPSES-1 FLOW TRANSIENT APPLICABILITY Title / Description ADolicability A.
Pump Start with Inadvertently Voided Discharge Lines o
RHR SOP-102A o
SIS SOP-201A B.
Expected Flow Discharge into Initially Empty Lines o
RCS (Safety and Relief Valve Discharge)
SOP-101A o
RHR (Suction Relief Valve Discharge SOP-102A o
SIS SOP-201A C.
Valve Opening, closing and Instability o
RHR SOP-102A o
SIS SOP-201A D.
Check Valve Closing and Delayed Opening o
RHR SOP-102A o
SIS SOP-201A E.
Water Entrainment in Steam Lines N/A F.
Transient Cavitation (Column Separation)
N/A G.
Steam Bubble Collapse Due to Rapid Condensation o
RHR SOP-102A o
SIS SOP-201A H.
Slug Impact N/A I.
Pump Start and Postulated Seizure with Full Lines o
RHR SOP-102A o
SI6 SOP-201A A-9
TABLE A-2 POTENTIAL WATER HAMMER EVENTS SYSTEM EVENT RCS RHR SIS
- 1) Pump Startup X
X
- 2) Expected Flow X
X X
- 3) Valve Opening X
X
- 4) Check Valve X
X
- 5) Water Entrainment
- 6) Transient Cavitation
- 7) Steam Bubble X
X
- 8) Slug Impact
- 9) Pump Startup, Stopping X
X A-10
APPENDIX B STAINLESS STEEL MATERIAL PROPERTIES This appendix presents the material properties used in the CPSES-1 leak-before-break analysis and explains their derivation.
Section B1 discusses the stainless steel base metals and Section B2 discusses the stainless steel welds.
Because there exist no CPSES-1 archival material property tests conducted at temperatures near the upper range of normal plant operation, the WHIPJET program has established a data base of industry tensile data for similar heats of materials for temperatures that span the plant operating temperature range.
This appendix demonstrates room temperature representativeness of the CPSES-1 stainless steel data as compared to industry data.
Because of that representativeness, and because many of the industry room temperature specimens were of the same heat as the 550*F indurtry specimens, use of the industry data is justified.
For the materials of interest - austenitic SA 376/312 Types 304, 316, and 316L stainless steels and Types 308 and 316 stainless steel welds - the stress-strain curves are of primary interest.
The best fit tensile properties for the base metal are used in this report for determining leakage size cracks.
Lower bound flow stress values are used for crack stability calculations.
B-1
SECTION B1 SELECTION OF STRESS-STRAIN CURVES FOR CPSES-1 AUSTENITIC STAINLESS STEEL BASE METAL IN HIGH ENERGY LINES B.
1.1 INTRODUCTION
The purpose of this section is to show the room temperature representativeness of the CPSES-1 stainless steel base metal data as compared to industry data, and to present the stress-strain curves that can be used in a leak-bufore-break (LBB) analyses at both room and elevated temperatures.
Austenitic stainless steel high energy piping at CPSES-1 is made of SA376/312 Type 304 and 316 wrought stainless steels.
For the safety injection system (SIS) and the residual heat system (RHR) removal lines under consideration for LBB, greater than 95% of the steel is SA 376/312 Type 316.
A few sections of SA312 Type 304 pipe up to 7 feet long exist in the vicinity of some weldolets and 90-degree elbows and have been weld annealed, pickled, and water quenched - most likely for the purpose of relieving weld residual stresses.
The reactor' coolant line (surgeline) under consideration is 100 percent SA 376/312 Type 316 steel.
Nozzles attached to the main loop are near the high stress location in the 10-inch SIS line.
The high stress location in the 14-inch RCS line is located near the loop nozzle.
These nozzles are made from Type 304 stainless steel and safe ends are typically Type.316 or 316L steel.
The high stress location in the,12-inch diameter RHR line is located near a check valve.
A summary of typical pipe fitting materials'is given in Table B-1-1.
As discussed below, limited tensile property data-for the Type 316 stainless steel is available.
The one Type 316 stress-strain at 550*F shown in Figure B-1-6 was reconstructed from limited data obtained from references 2 and 16.
Even so, it is reasonably bound by the two 550'F Type 316L curves (curves 5 and 6).
Therefore, the results for Type 304 and Type 316L stainless steel are used to set the bounds for Type 316 stainless steel.
The Type 316 stainless steel data are always higher than the bounding Type 304 stainless steel results.
The ASME Boiler and Pressure Vessel Code Section III, Table I-2.2 shows the Code yield strength values of Type 304 and 316 stainless steels to be identical near room temperature but at 550'F the Type 316 yield strength value is approximately 3% higher than Type 304 stainless steel value.
Table I-3.2 in the ASME Code shows ultimate tensile strengths of Type 304 and 316 stainless steels near room temperature to be identical but at 550'F the Type 316 ultimate strength is 13% higher than B-2 t
em,.,--.e-.~,-
.-,,,-er, nn
,-,nm,--,-,
-,--,r,,
,_---n-
-, - - _ - - - - - - - ~ -, - - +. - -
r
the Type 304 stainless steel value.
Therefore, this suggests that Type 316 steel has higher a work hardening rate at 550*F than does Type 304 steel.
The fact that the ASTM /ASME specifications A376/312 (SA376/312) allow both Type 304 and 316 stainless steels to be used for seamless and welded pipe for high temperature service suggests similarity.
In terms of chemical composition, some molybdenum and a bit more nickel are added to increase the corrosion resistance of the Type -316 compared to the Type 304 steel.
In general, these trace element additions also slightly increase the high temperature toughness and strength.
B.1.2 COMPARISON OF YIELD STRENGTH DATA CPSES-1 Certified Material Test Report (CMTR) reports of yield strength and ultimate tensile strength values are used as-parameters to establish representativeness of the industry data base as compared to the CPSES-1 piping.
Yield strength and the work hardening coefficients are the dominant tensile parameters used in estimating pipe leakage.
Therefore, the comparisons of room temperature yield strength, and ultimate tensile strength are presented in this subsection.
CPSES-1 stainless steel tensile data are limited to room temperature yield, ultimate tensile strength and percentage elongation.
The equivalent industry yield and ultimate tensile strength data are compared to the CPSES-1 data in Figures B-1-1 and B-1-2 for Type 304 stainless steel and in Figares B-1-3 and j
B-1-4 for Type 316 steel.
Type 304 yield strength data for CPSES-1 (Figure B-1-la) tend to fall at the upper end of the industry yield strength data range (Figure B-1-lb) probably due to the weld solution, annealing.
The ultimate tensile strength CPSES-1 data (Figure B-1-2a) is bounded by the industry data (Figure B-1-2b).
Most of the room temperature plastic stress-strain curve lies beyond the yield stress an.d is therefore bounded by industry room temperature stress-strain data at larger plastic strains.
Note that only a small portion of the piping under consideration for rupture hardware removal is made of Type 304 steel.
Also, no high stress locations occur in the Type 304 base metal.
Type 316 yield strength data for CPSES-1 (Figure B-1-3a) are bounded by the industry yield strength data (Figure B-1-3b).
Type 316 ultimate tensile strength data for CPSES-1 (Figure B-1-4a) are bounded by the industry ultimate tensile strength data (Figure B-1-4b).
B-3
{
B.1.3 INDUSTRY STRESS-STRAIN CURVES Few stress-strain curves are available in the industry data base because usually only yield, ultimate tensile strength, and percentage elongation properties are provided.
Available industry true stress-versus-true strain curves for Type 304 and Type 316L steel are shown in Figures B-1-5 and B-1-6.
An example of the higher work hardening rate for Type 316L steel compared to Type 304 steel is shown in Figure B-1-7.
References appearing in Figures B-1-5 and B-1-6 correspond to the references at the end of this appendix.
B.1.3.1 Discussion of Stress-Strain Curve Data For Type 304 tensile tests at 550*F, the lower bound tensile properties (Figure B-1-5) correspond to specimen curves 4 and 5 (specimen codes 4BT4 and 4CB-T4).
Upper bound tensile properties would be represented by curve 9 (specimen code A34),
while best fit properties by curve 2 correspond to specimen A8-40.
The yield strengths for Type 304 specimens 4BT4, 4CB-T4, A34, and A8-40 are 23.1, 23.7, 34.2 and 26.8 ksi, respectively.
A lower bound fit for Type 304 steel is indicated in Figure B-1-8.
The fit itself is taken from Reference 2.
The actual data for specimen 4CB-T4 is given in Table B-1-2.
A best fit for Type 316L true stress-strain data is indicated in Figure B-1-9.
It is used for leakage size crack determination.
For Type 316L stainless steel at 550*F, the only two available tensile curves (Figure B-1-6) are almost identical (curves 5 and 6, specimens DP2A56 and DP2A55).
The yield strengths are 27 and 24 ksi, respectively, for specimens DP2A55 l
and DP2A56.
The ultimate tensile strengths for these same two specimens are 68.4 and 68.0 ksi, respectively.
Type 316L steel has slightly different teasile properties than the Type 316 steel used at CPSES-1.
Note that except for the safe ends, there is no Type 316L steel in the WHIPJET lines; 1
but, Type 316L data are used to estimate Type 316 tensile properties.
The ASME Code Section III Table I-2.2 shows the room temperature yield strength of Type 316L to be 5 ksi lower than Type 316 steel.
Similarly, Table I-3.2 shows the Type 316L ultimate tensile strength to be 5 ksi lower than Type 316 l
I steel.
Similar trends should hold at 550*F.
The industry Type 316 tensile data is limited, but the room temperature and l
550*F yield and ultimate tensile strength data are bounded l
by the industry Type 316L tensile data.
l l
l B-4
The type 304 steel lower-bound stress-strain curve is a lower bound for both Type 316 and 316L.
Therefore, for conservatism, the Type 304 lower bound tensile properties (specimen 4CBT4) are used as the reasonable lower bound tensile properties in the reported base metal stability analysis.
B.1.3.2 Ecuations The equation for the true stress-strain (a - c) curve fit above yield for Types 304 and 316 stainless steel to be used in the LBB analysis is:
[a \\
(L )
[g )
D a
+
'o l
- o l a,l where e and o are the yield strain and yield o
n strength, and n are curve fit coefficients to this Ramberg-Osgood tensile relationship.
Recent EPRI research suggest that a curve fit over small strain ranges (less than one percent) is preferable [11].
For this reason the above Ramberg-Osgood functional form used true stress-strain data only to one percent.
The appropriate curve fit parameters for CPSES-1 application are:
E = 25.5E6 psi
- o = 24.5E3 psi r
of = 46.2E3 psi Type 316L-best fit based on specimen DP2A56 --
u = 68.0E3 psi used for leakage calculations o
3.75 a
=
n=
4.82 0.3 v
=
E = 25.5E6 psi
- o = 23.7E3 psi og = 42.8E3 psi Type 304 and 316L-lower bound based on specimen 4CBT4 -- of
- u = 61.9E3 psi used for crack stability analyses 7.3 a
=
n=
8.9 v=
0.3 B-5
E is the elastic modulus, of is the flow stress corresponding to the average of yield (oo) and ultimate tensile strength (ou), and v is Poisson's ratio.
B.1.4
SUMMARY
After examining CPSES-1 tensile properties and industry data, the lower bound Type 304 and best fit Type 316 stainless steel stress-strain properties have been determined.
These data are considered representative of the CPSES-1 piping materials.
4 I
l I
l B-6
TABLE B-1-1 TYPICAL PIPE FITTING STEELS IN CPSES-1 SIS, RCS AND RHR LINES FITTINGS MATERIALS Elbow SA403 WP316 SA182 GRF316 or F304 SA240 Type 316 SA376 Type 304 Sockolet SA182-F304 Reducer Type 304 Type 316 Reducing Tee SA376 Type 316 Weldolet SA182-F304 SA182-F316 Special Ends SA403 WP304 Check Valve SA182 F316 Nozzles SA376 Type 316 SA312 Type 304 isozzle Safe Ends SA182 Type 316 or 316L SA376 Type 316 or 316L Coupling SA182 Type F304 or F316 B-7
TABLE B-1-2 EXPERIMENTAL DATA FOR LOWER BOUND TENSILE CURVE FOR TYPE 304 BASE METAL Specimen 4CBT4 (tensile)*
True Strain True Stress (ksi) 0.00460 23.6 0.00851 24.6 0.0163 26.2 0.0237 27.4 0.0313 28.9 0.0383 30.4 0.0468 31.9 0.0550 33.4 0.0596 34.9 0.0687 36.8 0.0859 40.1 0.101 43.1 0.116 46.8 0.129 50.3 0.162 57.2 l
l l
l I
l l
f I
l i
- Elastic data unavailable l
l l
B-8
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Room Temperature 304 Bcise 6
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sa 42 4s s
a4 NY Room Tempemture 304 Base 6
i b)!
a-l l
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l l
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l l
=
x l
l l
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2-i l
l l
l l
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l l
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l l
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o, 30 34 38 42 46 50 S4 WW $drergth (W) 12EG irdu*y r
Figure B-1-1 Yield Strength Coaparisons for CPSES-1 and Type 304 Stainless Steel Base Metal ;
B-9 1 I
Room Temperature SA312/376 Type 304 3
a);
2-t 1
'~
r; s'
i i
i
/
/
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0 60 64 68 72 76 80 84 88 92 96 100 (luQ l
Room Temperature SA312/376 Type 304 l
i li S:
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}
i l
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cl l j
j
)
1-t R
R 0,...............................
60 64 6 ~,
72 76 80 84 88 2
96 100 LR M Ten W h(@
Figure B-1-2 Ultinnte Tensile Strength Carparisons for CPSES-1 and Indust.ry Type 304 Stainless Steel Base Metal
'3..
B-10 j
Room Temperature 316 Base i6
,p 1s -
a) p/
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/
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g o-l I
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30 34 38 42 46 50 54 g sigh M Figuts B-1-3 Yield Strength Ccxrparisons for CPSES-1 and Industry Type 316 Stainless Steel Base Metal B-11
-1 l
k Room Temperature SA312/ 76 Type 316 to s
9-B) a-7-
4 f
[
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=
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o 70 74 78 a2 S6 90 94 98 Ut' m*p Tende 9rerr/h (ksJ) e iUfJ W y Figure B-1-4 Ultivate '.Mitsne,sdrmgth Conparisons for CPSES-1 and uidustry Type 316 Stainless Steel. 'Lsc 2btal P
r B-12 '
/
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+
1
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.(
s f
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-3 9
8 100 3
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y 2 BCL A8.40 330 1
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~
6 DM RILA.8 100 4
I 7 BQ, A45-1 550 8 BC, A8-48 70 5
20 \\.[' / %e' 9 EQ, A34 550 6
10 trNlRC CEF.P1 550 6
!(
11 DDERC QD-P1 550 10 Ii 12 BQ.
AD 1 550 11 i
1 0
~
O.1 """
l
- 0. 0 t
T o'. 2 0;3 tur smni i,
2 J
Figure B-1-5 Stress-Strain curves for SA376/312 j
Type 304 Stainless Steel Base Metal i
L 1
?
/
B-13 5
e,
,-g-,
n
(
l l
l l
140 220 1
2 m
7 4
5 6
80 C
MRE SECIMFN TDiP
_g g
[
1 BCL DP2A51 72F 7 316L 7
2 BCL DP2A52 72F 7 316L u"
3 BC.
DP2A33
'0OP 7 316L l
40 4 BCL DP2A54 300F 7 316L p
5 BCL DP2A55 550F 7 316L 6 BC.
DP2A56 550F 7 316L 7 W/AVI 68 T2 550F 2 316 20 I
i 0
0 0.1 0.2 0.3 0.
0.5 Tna Strain
)
Figure B-1-6 Stress-Strain Curves for SA376/312 Type 316L and 316 Stainless Steel Base Metal B-14
i I
I I
I 4
i e
i I
60 T-550 F 1
50 2
40 v
m m
30 em
- 3 g
20
- 1. DP2A56 TYPE 316 BASE METAL 10 ~ }
- 2. 4CBT4 TYPE 304 BASE METAL
~
)
I I
I I
I I
I i
0' 0.02 0.04 0.06 0.08 o,10 O.0 TRUE STRAIN l
t I
/
Figum B-1-7 Icar Bound for Type 304 and Type 316 Stainless Steel Base tietal i
B-15 1
30 I
I I
I I
I I
I I
W/AWI Specimen 4CBT4 (550*F) o curvefit of above
=
n 0
20._
i n
gc g3 u.)- u.J + (m) g 6
o,-
23.7E3 psi 0
E 25.5E6 psi
=
d 10,_
i
- 61. 9E3p si.
o Type 304 of=
42.8E3 psi 8
U Base Metal 7.3 a
=
8.9 n
=
e 0
e i
e i
I 0
0.002 0.004 0.006 0.008 0.01 True Strain i
Figure B-1-8 Coaparison of Experinental True Stress-versus-Tnie Strain Curve Against the Fitting Ftnction l
(Imer Bound Fit Type 304) l I
i l
B-16 l
BCL R2-A5-6 fME STRESS - TEE STMIN, 559F 32.96 t
ITPE 316 (BASE HETAL)
,#=
n l
24.99-S j
i m
7 20.00 I
! N.I
- i-) 9 16,96 s
i I
5 25 5r4 9t 12.99 24.5E3 psi i
o*.
46.2E3 pst j
,g,g t
68.0E3 poi I
o 3.75 e
e.
4.e2 4.99-l l
'k00,
TEE SIMIN ( *19x*-3 )
4.99 8.00 Figure B-1-9 Couparison of Experinental True Stress - versus True Strain Curve versus the Fitting Fmetion (Best Fit Type 316L)
B-17
SECTION B2 SELECTION OF STRESS-STRAIN CURVES FOR CPSES-1 AUSTENITIC STAINLESS STEEL WELDS IN HIGH ENERGY LINES B.
2.1 INTRODUCTION
The purpose of this section is to show the representativeness of the CPSES-1 stainless steel Welds as compared to industry data, and to present the stress-strain curves at both room and elevated temperature.
It is important to note, however, that the lower bound flow stresses for SAW and SMAW developed in this section are not used in the master curva approach to crack stability.
A far more conservative flow sbress - 51 ksi - is used per the procedures of Appendix G.
Altheugh the flow stress of 51 ksi is shown below to be too conservative, crack stability for the flux welds is nevertheless demonstrated.
Austenitic stainless steel high energy piping welds at CPSES-1 consist mainly of Type 316 gas tungsten arc welds (GTAW).
The balance of the welds use the submerged arc weld (SAW) process (32%) or the shielded metal arc weld (SMAW) process (7%) as shown in Figure B-2-1.
In most of the SAW cases, the welds are partly GTAW, SMAW, and SAW, but predominantly SAW.
Also, for the SMAW cases the root passes use the GTAW process, but are predominantly SMAW.
B.2.2 WELDING OVERVIEW The butt welds considered under the CPSES-1 WHIPJET program were welded on-site using the GTAW process.
On-site welds include a few GTAW welds fabricated in an on-site shop.
Welds in spool pieces fabricated off-site used the GTAW process entirely or the GTAW process iullowed by the SMAW and/or SAW process.
The GTAWs use ER-316 or 2R-316L filler metal.
The SAWS were made using Type ER-316 Ciller metal with a combined GTAW/SMAW/SAW procmss.
The SMAWs were made with a GTAW/SMAW process using Type E9-316 and E316-16 electrodes.
Some SAW and SMAW welds used 316LH consumable inserts.
One SMAW weld used E-309 electrode for the CTAW roo* pass; however, the bulk of the filler metal was ER-316/PR316L.
SMAW material properties were used for all shop welds which have SMAW intermediate and balance passes.
SAW material properties were used for all welds that had any SAW processing.
B-18
The end preparation (prep) configuration groove design is in accordance with the CPSES-1 standards or design drawings contained in the weld specification.
End prep configuration for standard butt welds, as in the WHIPJET lines, uses a J-level; under certain circumstances, end preparation consists of a J/V-bevel combination.
Further details on material alignment, preheating, welding gases, welding techniques, interpass tenperatures, weld finishing, stress relieving, heat input, and a weld summary can be found in Appendix C of this report.
A comparison between CPSES-1, Westinghouse /AWI (H/AWI) from Reference 14, and the naval laboratory DTNSRDC from Reference 16 is given in Table B-2-1.
The brown and Root, Inc. and ITT Grinnel procedures are labeled BR and ITT, respectively.
A comparison between CPSES-1 and General Electric Company (GE) field welding techniques is shown in Table B-2-2; the GE field welding procedure was used to produce fracture and tensile specimens analyzed here for lower bouwd curves to be applied to CPSES-1 LBB analyses.
H/AWI SMAW well data, which are also used for bounding purposes, are quite similar to those shown in Table B-2-2 with heat inputs of about 30 kJ/in; other details of the welding procedures used in the H/AWI specimens are not available.
Overall, the GE, H/AWI, and CPSES-1 SMAW welding procedures are equivalent.
The H/AWI test specimen welding procedure should not be confused 3
with the H weld procedure used for the CPSES-1 surge line SMAW weld.
Details of the CPSES-2 shop welding techniques are given in Table B-2-3 with further details given in Tables B-2-4 to B-2-6.
Weld procedures for Battelle Columbus Laboratories (BCL)
SAW welds for test specimens A45SW-1, A45SW-2, A45W-1, and A45W-2 are listed in Table B-2-4.
The H/AWI SAW weld procedures are similar to the BCL and CPSES-1 shop welding procedures.
B.2.1 COMPARISON OF TENSILE DATA Yield strength as determined from Certified Material Test Reports (CMTRs) was chosen as the parameter for establishing representativeness of the industry data base as compared to the CPSES-1 piping.
Yield strength and the ultimate tensile strength j
are the dominant tensile parameters ueed in estimating crack stability.
Therefore, the comparisons of room temperature yield strength and ultimate tensile strength are presented in this subsection.
CPSES-1 GTAWs have ER-316/ER-316L filler metal (85%) and EE-308/ER-308L filler Metal (15%) as shown in Figure B-2-2.
GTAW ultimate tensile strength data at room temperature are shown in Figure B-2-3.
No CPSES-1 GTAW yield strength data for Type 316 filler metal exist.
B-19
CPSES-1 SMAW yield strength data at room temperature are shown in Figure B-2-4.
The yield strengths for E308 welds range from 57 ksi to 66 ksi.
Available industry yield strength data are also
<,ompared with the CPSES-1 data in Figure B-2-4.
Since yield strength test data are not required by ASME Code Sections II, III, or IX for weld metal, ultimate tensile strength data are more available than yield strength data.
CPSES-1 and industry ultimate tensile strength data are compared in Figure B-2-5 for Type 308 filler metal and in Figure B-2-6 for Type 316 filler metal.
Similar comparisons for ultimate tensile strength are given in Figures B-2-7 and B-2-8.
CPSES-1 SAW ultimate tensile strength data at room temperature are shown in Figure B-2-8 for Type 316 and 308 filler metals.
CPSES-1 data for SAWS with Type 316 filler is unavailable.
B.2.3.1 Discussion of Tensile Data (GTAW) i Little industry tensile data exist for GTAWs with Type 308 filler.
Also, no industry tensile data exists for GTAWs with strictly Type 316L filler metal.
The ultimate tensile strength data does coincide with some of the CPSES-1 data.
Type 316 filler metal produces a somewhat higher ultimate tensile strength than does the Type 308 filler metal.
The little data which exists is representative of the CPSES-1 GTAW data.
Only one GTAW l
tensile spr -imen with a Type 308 filler had data at both room temperature and 550*F.
The yield and ultimate tensile strengths of this specimen are 68.9 and 90.5 ksi, respectively.
B.2.3.2 Diajussion of Tensile Data (SMAW)
For th2 SAWS with Types 308 ano 316 filler metal, the industry yield strengi.h and ultimate tensile strength usually bracket the CTSES-1 data as shown in Figures B-2-4 through B-2-5.
The i.1dustry data are judged to be representative of the CPSES-1 SMAW welds.
B.2.3.3 Didcussion of Tensile Data (SAW)
The SAW welds J:.-e a combination GTAW/SMAW/SAW process.
The industry data aty use either a GTAW/SAW or a GTAW/SMAW/SAW process.
Indbstry room temperature yield strength data ranges from 42.0 to to ksi for SAW welds, 52 to 70 ksi for Type 308 SMAWs, and 47.0 to 76 ksi for Type 316 SMAWs.
The industry room temperature ultimate tensile strengths range from 82 to 98 ksi for Type 308 SMAWs, from 74 to 95 ksi for Type 316 SMAWs, from 82 to 99 ksi for Type 308 SAWS, and from 82 to 92 ksi for Type 316 SAWS.
Comparing the Type 308 SAW industry data to the CPSES-1 data shows that the CPSES-1 data are at the lower end of the industry data.
This may be due to the industry tensile specimens B-20 i
J
having fewer SMAW passes than the CPSES-1 specimens.
Type 316 SAW data for CPSES-1 are unavailable, but the industry tensile data for Type 308 and 316 filler metals are similar.
- Overall, the industry tensile data are judged to be representative of the CPSES-1 SAW properties.
B.2.4 INDUSTRY STRESS-STRAIN CURVES Stress-strain curves for GTAWs are shown in Figure B-2-9 and B-2-10.
The few stress-strain curves available from the industry data base are shown in Figure B-2-11 for SMAW Type E-308 welds.
The few stress-strain curves available from the industry SAW data base are shown in Figure B-2-12.
B.2.4.1 Discussi(
of Stress-Strain Curve Data (GTAW) 0 For GTAWs at 550 F, the lower bound stress-strain curve is curve 5 (specimen code A25WO-1).
The best fit curve is given by stress-strain curve 2 (specimen code GEF-WI).
All the 550'F curves fall below the room temperature stress-strain curve.
The 550*F GTAW stress-strain curves are bounded by the 550'F base metal curves 7 and 8 as shown in Figure B-2-10.
Tensile properties for GTAWs are at least as good as the base metal properties.
For this reason the base metal LBB calculations are sufficient for the GTAW welds.
B.2.4.2 Discussion of Stress-Strain Curve Data (SMAW) 0 For Type 308 SMAW tension tests at 550 F, the lower bound tensile properties (Figure B-2-11) correspond to curve 3 (specimen code 4CSMAWT4) and the upper bound properties to curve 1 (specimen code SG1SMAWAR).
The yield stter.gths are 49.4 and 45.7 ksi for specimens 4CSMAWT4 End SG1SMAWAR, respectively.
The lower bound tensile curve that could be used in the stability analysis is curve 3 in Figure B-2-11.
This curve is representative for use in CPSES-1 stainless steel SMAW welds in high energy lines.
The actual W/AWI curve fit is shown in Figure B-2-13 along with the curve 3 data from Figure B-2-11.
This low strain range fit is good through strains of approximately 10%.
Tabulated values of the actual stress-strain curve 3 are given in Table B-2-1.
The curve fit parameters for specimen 4CSMAWT4 are from Reference 14.
B.2.4.3 Discussion of Stress-Strain Curve Data (SAW)
For the Types 316 and 308 shop welds at 550 F, the lower bound tensile properties (Figure B-2-12) correspond to curve 1 (specimen 316-3AR) and the upper bound properties to curve 2 (specimen A45SW-1).
The yield strengths are 44.8 and 47.0 ksi for specimens 316-3AR and A45SW-1, respectively.
The lower l
B-21
bound curve which fits to 1% strain, is shown in Figure B-2-14.
Tabulated values of the actual stress strain data are given in Table B-2-7.
B.2.4.4 Ecuations The equation for the stress-strain curve fit to be used in the LBB analysis is the Ramberg-Osgood expression presented in Section B.1:
f.C )
=
E \\
+
E a
'o l q ol' (Oo)
EPRI research suggests that a fit over small strain ranges (less than one percent) is best.
For this reason the above Ramberg-Osgood functional form used true stress-strain data only to one percent. The appropriate parameters for CPSES-1 analyses would therefore be:
26.0E6 psi E
=
'o =
49.4E3 psi of =
57.0E3 psi Austenitic SMAW 308 filler lower bound based on u = 64.7E3 psi specimen 4CSMAWT4.
Th!.s of o
could be used for SMAW crack a=
9.0 stability analysis n=
9.8 v=
0.3 26.0E6 psi E
=
44.8E3 psi a
=
o og =
53.8E3 psi Austenitic SAW 308/316 filler lower bound based 62.9E3 psi on specimen 316-3AR.
This a
=
u of could be used for SAW crack stability analyses.
aa 3.32 5.91 n
=
o=
0.3 B-22
B.2.5
SUMMARY
After examining CPSES-1 tensile properties and industry data, lower bound stress-strain relations have been determined.
These are the lower bound data available to the industry and are considered representative of the CPSES-1 piping wolds.
This discussion has justified the use of flow stresses of 53.8 ksi (SAW) and 57.0 ksi (SMAW) for flux weld crack stability analyses.
However, an additional - and perhaps unnecessary -
margin of conservatism is achieved by using a flow stress of 51 ksi for both SAW and SMAW stability analyses.
The selection of 51 ksi was made in order to remain consistent with the procedure of Appendix G.
B.
2.6 REFERENCES
1.
Letter to attendeos for "Leak-Before-Break:
International Policies and Supporting Research" from Wilkowski,-G. M.
at Battelle Columbus Laboratory, Columbus, Ohio, October 28-30, 1985.
2.
Landes, J.
D.
and McCabe, D.
E.
Elastic-Plastic Methodoloav to Establish R-Curves and Instability Criteria, Report NP-4768, Electric Power Research Institute, Palo Alto, California, February 1986.
3.
- Norris, D.
M.,
"Proposed Interim Evaluation Procedure for Austenitic Stainless Steel Flux Welds", ASME Boiler and Pressure Vessel Code,Section XI report prepared by the Electric Power Research Institute,'Palo Alto, California, March 1985.
4.
- Kumar, V.,
German, M.
D.,
and Shih, C.
F.,
An Enoinee.i..J Acoroach for Elastic Plastic Fracture Analysis, NP-1931 Electric Power Research Institute, Palo Alto, California, Tuly 1981 Page 6-23.
i 5.
Personal correspondence from R. J. Olson at Battelle Columbus Laboratories, Columbus, Ohio to D. F. Quinones at Robert L.
Cloud & Associates, Berkeley, California, March 25, 1986.
6.
Wilkowski, G. M.
et al, Decraded Pioina Procram-Phase II.
Semiannual ReDort. October 1984-March 1985, NUREG/CR-4082, Vol.
2, U.S. Nuclear Regulatory Commission, Washington, D.
C.,
July 1985.
Figures 2.6.2, 2.7.2, A.1, A.2 and Fig.
3.5.4.
7.
Personal correspondence with C. W. Marshall at Battelle I
Columbus Laboratory, Columbus, Ohio, to D. F. Quinones at Robert L.
Cloud & Associates, Berkeley, California, July 5, 1985.
Unpublished tensile data for Type 316 stainless steel from Miltan Vagin's (US NRC) Degraded Piping Program.
B-23
8.
Personal correspondence from D. M. Norris at the Electric Power Research Institute, Palo Alto, California to D.
F.
Quinones at Robert L. Cloud & Associates, Berkeley, California, August 1986.
9.
Egoort of the U.
S.
Nuclear Reculatory Commission Pioina Feview Committee; Evaluation of Potential for Pipe Breaks, NUREG-1061, Vol.
3, U.S.
NRC, Washington, D.C.,
November 1984.
10.
Personal correspondence from R. Hays at David Taylor Naval Ship Research and Development Center, Annapolis, Maryland, to D. M. Norris, EPRI, August 20, 1987.
11.
Personal correspondence from G.
Kramer, Battelle Columbus Laboratory, Columbus, Ohio to D.
F. Quinones, Robert L.
Cloud & Associates, Berkeley, California, February 18, 1986.
12.
Wilkowski, G.,
et. al., Analysis of Experiments on Stainless Steel Flux Welds, NUREG/CR-4878, U.S.
- NRC, Washington, D.C.,
April 1987.
13.
- Horn, R.M.,
- Mehta, H.S.,
- Andrews, W.R., Ranganath, S.,
Evaluation of the Touchness of Austenitic Stainless Steel Pipe Weldments, Report NP-4668, Electric Power Research Institute, Palo Alto, California, June 1986.
14.
Landes, J. D. and McCabe, D.E., Touchness of Austenitic Steel Picewelds, Report NP-4768, Electric Power Research Institute, Palo Alto, California, October 1986.
15.
Wilkowski, G.M. et. al, Decraded Pioina Procram Phase II, Battelle Columbus Laboratory, NUREG/CR-4082, Vol.
3, March 1986.
16.
- Hays, R.M.,
Vassilaros, M.G.,
and Gudas, J.P.,
Fracture Analysis of Welded Tvoe 304 Stainless Steel Pice, NUREG/CR-4538, U.S. Nuclear Regulatory Commission, Washington D.C.,
May 1986.
B-24
7ABIE B-2-1 EMPARISON OF CRTr:S-1 WEID 'IEGNIQUES FOR GIMs WfAWI DINSRC CPSES-1 ITP N9 (TEST SPECIMENS)
(TEST SPECIMENS)
- R Root Weld GIM GIM GIM GIM Filler Weld GIAW GDM GI%W GTAW Filler Material ER-308I/ER-316L 308L ER-316/ER-316L ER-316/ER-308 ER-316 ER-308/ER-308L ER-316L Filler rod dia. (in) 1/16 - 1/8 1/16 - 1/8 Argcn gas flow (CEH) 5 - 15 (min) 16 - 35 m
Anps 50 - 200 10 - 15C Volts 8 - 11 9 - 15 Preheat T e ature 60 F 60 F Interpass T e ature 350 F (max) 350 F (max)
Stress Felieving None None None Heat Irput (KT/in) 30 24 - 43 11 - 45
=
TMEE &-2-2 CUIPARISCM OF CPSES-1 NEID 'IECHiIQUES FOR SMHs CPSES-1 GE h
('IEST SPECIMENS)
ITT IR H
Root Weld GUN (namal)
GUN (mariual)
GDH (armaal)
GDW (namal)
Filler Weld SMAN (namal)
SMAN (manual)
SMMi (namal)
SMMi (namal)
Filler Material ER-308/E308 E308/D08L ER-308 or ER308I/E308**
ER308/E308 E316/E316L Filler rod dia. (in) 1/8 3/22 - 5/32 3/32 - 1/8 Argon gets flow (CHI) 20 16 - 35 (root) 15 Anps 90 - 120 45 - 140 40 - 150 Volts (GDM) 14 9 - 15 8 - 28 Volts (SMMf) 19 - 22 22 - 27 Preheat temperature 60 F (min) 50 F (min) 60 F (min) 60 F (min)
Interpass temperature 350 F (max) 350 F (max) 350 F (max)
'350 F (max)
Stress relieving None None None Heat Iqut (kJ/in) 22 - 30 12 - 108 10 - 124 (estimated)-
Taken fram a GE pWires for an as-welded butt wld (see Ref.1)
- or ER3161/E316
'IABIE B-2-3 CCMPARISCE OF CPSES-1 WEID 'IECINIQUES EUR SANs CPSES-1 BCL GE*
p (TEST SPECDENS)
('IEST SPBCIMENS)
ITT IR W
Hoot weld GIAW(marmal)
GIAW(manual)
GDW(manual) l Filler weld SMAN/SAW SMAW/SAW SMAN/SAW
)
Filler material ER-308/E308 ER-316/E316 ER-308 or ER-308I/E316 l
or ER-316L Filler rod dia. (in.) 3/32 - 1/8 3/32 - 1/8 3/32 - 5/32 Argon gas flow (CHI) 22 (root) 10 16 - 35 (root)
None None tn i
4 Anps 126 - 720 200 - 650 s3 Volts (GG W) 18 - 2G 9 - 15 Volts (SAW) 24 - 30 23 - 36 30 - 35 i
Preheat t=:mpeu.d^ulre 60 F (min) 50 F (min) 50 F (min)
Interpass % - d'urre 350 F (max) 350 F (max) 350 F (max)
Stress relieving None None None Heat Input (kT/in) 18-42 9.6 - 86 82 - 120 (typically 36 -38) l
- Taken frm a E pu v inre for an as-welded tutt weld (see Ref.1).
1
t TABLE B-2-4 CPSES-1 (ITT) SAW TECHNIQUES i
Root Weld Intermediate Balance Process Process Process GTAW SMAW SMAW or SAW (manual)
(manual)
(automatic) t Weld ER-308/308L ER-316 ER-308/308L Filler Metal ER-316/316L ER-316/316L Filler 3/32 - 1/8 3/32 - 1/8 3/32 - 1/8 Rod Diameter (in)
Heat Input 23 - 30 35 - 36 38 (kJ/in)
LINDE 709-5 Flux i
TABLE B-2-5 BCL SAW TECHNIQUES (TEST SPECIMENS) j Root Weld Intermediate Balance l
Process Process Process GTAW SMAW SMAW or SAW (manual)
(manual)
(automatic)
Weld ER-308 E308-16 ER-308 Filler Metal Filler 3/32 1/8 1/8 Rod Diameter (in)
Heat Input 7.5-70 5-43 12-50 (kJ/in)
Lincoln-Flux Weld t
4 I
B-28 i
f
TABLE B-2-6 GE SAW TECHNIQUES *
(TEST SPECIMENS)
Root Wald Intermediate Balance Process Process Process GTAW SMAW SMAW or SAW (manual)
(manual)
(automatic)
Weld ER-316L E316L-16 ER-316L Filler Metal Filler 1/8 1/8 3/32 Rod Diameter (in)
Heat Input 3.0-15 9-11 38-90 (kJ/in)
Cr-Ni Flux Alloy 3
1 4
B-29
TABLE B-2-7 EXPERIMENTAL DATA FOR LOWER BOUND TENSILE CURVE FOR TYPE 308 SMAW WELDS Specimen 4cSMAWT4 (tensile)*
True Strain True Stress (ksi) 0.0128 49.4 0.0231 51.6 0.0347 53.5 0.0457 54.9 0.0557 56.1 0.0666 57.4 0.0782 58.5 0.0882 59.7 0.0995 61.0 0.1120 62.3 0.1240 63.4 0.1350 64.6 0.1650 67.8 0.1900 70.9 0.2240 74.1 0.2580 77.7 0.3040 84.0 0.3650 91.3 0.4300 99.5 0.4960 108.0 0.5490 115.0
- Elastic data unavailable B-3)
TABLE B-2-8 EXPERIMENTAL DATA FOR LOWER BOUND TENSILE CURVE FOR TYPE 308/316 SAW WELDS Specimen 316-3AR (tensile)*
Engr. Strain Engr. Stress (psi) 0.00172 24804.
0.00289 34685.
0.00377 38688.
0.00469 41273.
0.00540 42017.
0.00610 43668.
0.00698 44456.
0.00776 45228.
0.00860 45559.
0.00944 46000.
0.0103 46315.
0.0110 46457.
0.0118 46836.
0.126 47009.
0.0134 46993.
0.0143 47230.
0.0150 47403.
0.0159 47387.
0.0164 47646.
0.0194 48490.
0.0231 49184.
0.0282 50178.
0.1810 62831.
- Elastic data unavailable B-31
RC N. and RHR Unee SWAW (7.3R) 14 SAW (s1.8K) 117 GTAW (80.9K)
Figure B-2-1 Weld Types in LBB Lines
{
N
~
RC. 5. ond RHR Unos ER-308/ER308L (15.45) 18 99 ER-sie/tR-stet (e44x) i Figure B-2-2 CIAW Filler Metal
Room Temporoture GTAW e
a)
Type 308 I
' 6
? 6 6 6 6 4 4 6 m
m
""T5"&.".".l"' M Room Temperature GTAW (308 Only) b)
'~
Type 308 i
1-E h
E E
E E
E
.e se es tm TM" Room Temperature GTAW C)
Type 308 & 316 l
l Figure B-2-3 Ultinnte Tenoile Strength Conparisons for CPSES-1 and Industry GTAWs with 308 and i
316 Filler i
B-33
SMAW Data - 308 Filler 4
a) 6-a-
3-t-
Y MfM- )
SMAW Data - 308 Filler b) i a-4=
3-1 1
i o
so u
a e
=
ro i
g s %g o-o Figure B-2-4 ie Stre
% ar for CPSES-1 and
(
~
B-34
Room Temperature SMAW - 308 Filler a) h p
s-
/
'~
j 5
i h
h
/
7 7
S S
S 2-
/
/
/
/
n f Ef f E3 f f f ? f f o
""tSr'M' Room Temperature SMAW - 308 Filler b) s-a-
4-E a-ll
... I. l.. l..
l o
tr=g* gw M Figure B-2-5 Ultinnte Tensile Strength cocparisons for CPSES-1 and Industry SMWs with 308 Filler B-35
SMAW Data - 316 Filler
/
/
8) 36-
/
f 3-2.5 -
/
2*
1 i.s -
/
?
7 5
/
/
/
/
/
l l
l 06-
/
/
/
/
/
/
/
/
/
/
I O
i i
i i
i i
i 46 so es so es m
m ao Tf8"'E#sM SMAW Data - 316 Filler 4
b) sa-3-
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Figure B-2-7 Ultiente Tensile Strength Comparisons for CPSES-l'and Industry SMA's with 316 Filler A
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APPENDIX C WELDING PROCEDURE C.1 INTRODUCTION The' stainless steel butt welds considered under the CPSES-1 WHIPJET program are welded using the Gas Tungsten Arc Welding (GTAW), the Shielded Metal Arc Welding (SMAW), or the Submerged Arc Welding (SAW) processes.
The SMAW and SAW (flux weld) processes start with the GTAW process.
All on-site welding, including that performed in the on-site fabrication shop, used the GTAW process.
The off-site fabrication shop welding used the GTAW process or a GTAW/ flux weld combination process.
The discussion that follows is based on the approved on-site GTAW welding procedures.
In some cases an automatic GTAW process was used.
Information is also provided for off-site flux weld procedures, as indicated.
C.2 WELD
SUMMARY
For the safety inp etion system (SIS), residual heat removal system (RHR), and reactor coolant system (RCS) lines under consideration, welding procedures from Brown and Root, Inc. (BR) and ITT Grinnell (ITT) were used.
One RCS (surge) line SMAW weld was fabricated by Westinghouse (H).
A total of 192 welds summarized in Table C-1 were examined.
All BR welds used the GTAW process with ER-3), ER-316L, and sometimas ER-308 bare wire filler metal.
Most ITT welds used the GTAW process.
Some ITT shop welds started with the GTAW process for the root pass then used the SMAW process for intermediate and final weld passes.
These welds are designated in Table C-1 as SMAW.
Some ITT welds used the GTAW process for root passes, the SMAW process for intermediate passer and the SAW process for the final passes.
These welds are designated in Table C-1 as SAW welds.
The ITT flux welds used ER-316, ER-308, and ER-316-L bare wire for the SAW process and E-316-16 covered electrodes for the SMAW process.
In one instance at: E-309 electrode was used for a GTAW root pass, but the bulk of the filler metal was ER-316 or ER-316L.
The ITT and H shop welds were performed on spool pieces shipped to CPSES-1 for attachment by ITT to the remainder of the piping system.
C.3 FRACTURE TOUGHNESS Welding 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 heat input range of GTAW welds studied in this report was 24 to 43 kJ/in.
For the C-1
SMAW process, the heat input range studied was 18 to 108 kJ/in.
For the Submerged Arc Welding process, the heat input range was 36 to 120 kJ/in.
The heat inputs used for GTAW welds were also checked by testing of weld test coupons to assure that excessive sensitization of the base netal would be avoided.
C.4 BASE METAL PREPARATION The ends of the pipe, valves, or fittings to be joined are prepared for welding by machining, grinding, sawing, plasma arc cutting, or shearing.
The pipe dimensions and configurations are in accordance with tie drawings of BR Specification 43B.
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.
Weldir.g is not performed on wet surfaces.
If welding is not started immediately after cleaning, the weld joint is suitably wrapped to prevent contamination.
The end preparation (prep) configuration for standard butt welds, as in the WHIPJET lines, uses a J-type-bevel or a J-V bevel combination.
If piping component ends are counterbored, such counterboring does not result in a finished wall thickness after welding less than the minimum design thickness.
Where 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.
C.5 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.
Localized heating of stainless steel piping is permitted provided the temperature does not exceed 350 F.
Heating of stainless steel above 800 F, the lower limit for sensitization, is not permitted.
C-2
C.6 FIT-UP Pit-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 and 2 systems.
The acceptable mismatch for ASME Section III piping is 1/32 inch per side.
C.7 PREHEATING Preheating of stainless steel is as fcllows:
Preheating is not required when the base metal temperature is above 60 F, which is often the case.
When the base metal temperature of the materials is below 60 F, preheating is performed by uniformly heating circumferentially to a temperature of 60 F minimum (warm to the touch).
C.8 WELDING GASSES The shielding and purging gas for the GTAW process is welding grade Argon.
Gas flow rates are within the range specified on the applicable technique sheet.
Shielding gas is used to form an ionized gas (plasma) for-metal arc transfer, to shield the molten base / weld metal puddle and to cool the tungsten electrode.
C.9 WELD FILLER METAL All weld material is purchased in accordance with ASME Section III requirements and the Field Piping Specificction.
Delta ferrite determination is required in accordance with Regulatory Guide 1.31 [1] for stainless steel weld filler metal.
C.10 WELDING TECHNIQUE The filler metal sizes shown on the technique sheets are the only sizes used.
The welding technique sheets are applicable to all position welding unless otherwise stated.
All vertical welds using the manual GTAW process are performed in the upward direction.
Each weld bead deposited is thoroughly cleaned of all oxidation, slag, or flux using a descaling tool and/or wire brush.
Surface defects, crater cracks, porosity, and undercutting are removed by grinding before depositing the next bead of welding.
C-3
C.11 INTERPASS TEMPERATURE The interpass temperature, specified in the welding procedure technique sheet, is checked on the piping adjacent (1 inch max.)
to the welding groove, using "Tempilsticks", a surface pyrometer, or an approved equal.
Tempilsticks are used most often.
After a weld bead has been deposited around the complete circumference of the joint, it is cooled to below the maximum interpass temperature before starting the next weld bead.
To control the interpass temperature of stainless steel welds, water quenching with clean, lint-free wipes soaked with demineralized water may be used to bring the weld below 350 F before continuing to weld.
General]y, by the time the deposited bead is thoroughly cleaned and inspected, the temperature will be below the interpass temperature.
C.12 WELD 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.
Weld edges are to merge smoothly with the base metal.
The radiographic acceptance criteria of ASME Section III, Division 1, Sub-Section NB is met.
C.13 STRESS RELIEVING i
Post weld heat treatment of stainless steel is usually not permitted.
C.14 CONTROL OF STAINLESS STEEL WELDING l
Fabrications of welded austenitic stainless steel Classes 1 and 2 comply with the requirements of ASME Section III and Section IX as supplemented by NRC Regulatory Guide 1.31 (1) and Regulatory l
Guide 1.44 [2].
A typical CPSES-1 stainless steel butt weld is performed in the following manner.
For stainless steel piping, a l
GTAW root insert usually is put down first.
This first segment is made with bare wire 1/8 inches thick maximum with nominally two layers from three beads.
After this, the balance of the weld is made from ER-316 or ER-316L electrodes using the GTAW process.
Sometimes ER-308 wire is used.
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 24 - 43 KJ/ inch.
End preparation usually consists of a J-type-bevel and no post weld heat treatment is performed (as just indicated).
l C-4 l
l
Certain off-site flux welds included in the WHIPJET program were made with similar practices except that 316 LH consumable inserts were used.
The maximum electrode diameter was 5/32 inch. Also, the submerged arc welding, SAW, process was used to complete some of these shop welds.
In these instances, Type ER-316 weld filler metal with Cr-Ni Alloy flux (Linde 7095) was used.
The heat input range for the SAW process was nominally 18 to 120 KJ/in.
C.15 REFERENCES 1.
Control of Ferrite Content in Stainless Steel Weld Metal, Reaulatory Guide 1.31, Rev.
3, U.S.
Nuclear Regulatory Commission, Washington, D.C.,
April 1978.
2.
Control of the Use of Sensitized Stainless Steel, Reaulatory Guide 1.44, U.S. Nuclear Regulatory Commission, Washington, D.C.,
May 1973.
l C-5 L
TABLE C-1 CPSES-1 WEID SuffiARY FOR SIS, MIR, AND RCS LINES 10.
LINE WELD PROCESS FILI.ER SOURCE SFOOL No.
PIECE 10.
1 BRP-Mi-1-RB-001 FW6 GTAW ER-316/ER-316L BR 2
BRP-Mi-1-RB-001 FW5 GTAW ER-316/ER-316L BR 3
BRP-MI-1-RB-001 FW4 GTAW ER-316/ER-316L BR 4
BRP-Mi-1-RB-001 FW3 GIAW ER-316/ER-316L BR S
BRP-Mi-1-RB-001 FW2 GIAW ER-316/ER-316L BR 6
BRP-Mi-1-RB-001 FW1 GTAW ER-316/ER-316L BR 7
BRP-Mi-1-RB-001 A
SAW ER-316 ITT 1
8 BRP-MI-1-RB-001 A
SMAW ER-316/E316-16 ITT 2
9 BRP-MI-1-RB-001 B
GIAW ER-316 ITT 2
10 BRP-RH-1-RB-001 C
GIAW ER-316 ITT 2
11 BRP-Mi-1-RB-001 A
GTAW ER-316 ITT 3
12 BRP-Mi-1-RB-001 B
GTAW ER-308 ITT 3
13 BRP-RC-1-RB-005 15A GIAW ER-308 BR 14 BRP-RC-1-RB-005 2-2 GIAW ER-316I/ER-316 BR 15 BRP-RC-1-RB-005 2-1 GIAW ER-316I/ER-316 BR 16 BRP-RC-1-RB-005 FW-1 GTAW ER-316I/ER-316 BR 17 BRP-RC-1-RB-005 A
SMAW ER-316/E316-16 ITT 1
16 BRP-RC-1-RB-005 B
SAW ER-316.-
ITT 1
19 BRP-RC-1-RB-005 C
SAW ER-316 ITT 1
20 BRP-RC-1-RB-005 D
GTAW ER-316 ITT 1
21 BRP-RC-1-RB-005 E
SAW ER-316 ITT 1
22 BRP-RC-1-RB-005 F
S?M ER-316 ITT 1
23 BRP-RC-1-RB-006 15 GTAW ER-308 BR 24 BRP-RC-1-R9-006 FW1 GIAW ER-316I/ER-316 BR 25 BRP-RC-1-RB-006 A
SMAW ER-316/ER-316-16 ITT 1
26 BRP-RC-1-RB-006 B
SAW ER-316 ITT 1
27 BRP-RC-1-RB-006 C
SAW ER-316 ITT 1
28 BRP-RC-1-RB-006 D
GIAW ER-316 ITT 1
29 BRP-RC-1-R&-006 E
SAW ER-316 ITT 1
30 BRP-RC-1-RB-006 F
SAW ER-316 ITT 1
31 BRP-MI-1-PB-002 FW9 GTAW ER-316 BR 32 BRP-Mi-1-RB-002 FW7 GTAW ER-308I/ER-308 BR 33 BRP-Mi-1-RB-002 FW6 GTAW ER-316/ER-316L PR 34 BRP-Mi-1-RB-002 FWSA GTAW ER-316/ER-316L BR 35 BRP-MI-1-RB-002 FW3 GIAW ER-316/ER-316L BR 36 BRP-MI-1-RB-002 2-1 GTAW ER-316/ER-316L BR C-6
'iABIE C-1 (CONT)
NO.
LINE WEID PROCESS FILIER SCXJRCE SPOOL NO.
PIECE NO.
37 BRP-E-1-RB-002 W2 GIAW E-316/ER-316L BR 38 BRP-RH-1-RB-002 N1 GIAW ER-316I/ER-316 BR-39 BRP-RH-1-RB-002 11 GTAW ER-308 BR 40 M P-RH-1-RB-002 A
SAW 316IR/ER-316 ITT 1
41 BRP-RH-1-RB-002 A
SMAN IR-316/E316-16 ITT 2
42 M P-RH-1-RB-002 B
SAW ER-316 ITT 2
43 BRP-RH-1-RB-002 C
GIAW ER-316 ITT 2
44 BRP-RH-1-RB-002 A
GIAW ER-316 ITT 3
45 BRP-RH-1-RB-002 B
GIAW ER-308 ITT 3
46 E P-SI-1-RB-037 FW7 GIAW ER-31G/ER-316L BR 47 BRP-SI-1-RB-037 EW20 GTAW ER-316I/ER-316 BR 48 BRP-SI-1-RB-037 FW5 CIAW ER-316 BR 49 E P-SI-1-RB-037 FW4A GTAW ER-316 BR 50 BRP-SI-1-RB-037 1-2A GTAW ER-316I/ER-316 BR 51 EP-SI-1-RE-037 FW1 GIAW IR-308 BR 52 BRP-SI-1-RB-037 FW3 GIAW ER-31.6 BR 53 BRP-SI-1-kB-037 FW2 GIAW ER-316I/ER-316 BR 54 BRP-SI-1-RB-037 1-1A GIAW ER-316I/ER-316 BR 55 BRP-SI-1-RB-037 FW13 GIAW ER-316/ER-316L BR 56 BRP-SI-1-RB-037 FW14 GIAW ER-316/ER-316L BR 57 BRP-SI-1-RB-037 14-4A GTAW ER-316/ER-316L BR 58 BRP-SI-1-RB-037 19-1 GIAW ER-316/ER-316L BR 59 BRP-SI-1-RB-037 IW6 GTAW ER-316/ER-316L BR 60 BRP-SI-1-RB-037 A
GTAW ER-3C3 ITT 1
61 BRP-SI-1-RB-037 B
GTAW ER-308 ITT 1
62 BRP-SI-1-RB-037 C
GIAW ER-316 ITT 1
63 BRP-SI-1-RIH)37 D
GIAW ER-316 ITT 1
64 BRP-SI-1-RB-037 E
GIAW ER-308 ITT 2
65 MP-SI-1-RB-037 A
SMAW ER-316/ER-316-16 ITT 2
66 EP-SI-1-RB-037 B
SAW 3161R/ER-316 ITT 2
67 EP-SI-1-RB-037 C
GIAW ER-316 ITT 2
68 BRP-SI-1-RB-037 A
SAW 316IR/ER-316 ITT 3
69 BRP-SI-1-RB-037 A
GTAW ER-316 ITT 5
70 BRP-SI-1-RB-037 B
SAW 316IR/ER-316 ITT 5
71 BRP-SI-1-RB-037 C
SAW 316IR/ER-316 ITT 5
72 BRP-SI-1-RB-037 D
SAW 316IR/ER-316 ITT 5
1 73 E P-SI-1-RB-037 E
SAW 316IR/ER-316 ITT 5
l 74 BRP-SI-1-RB-037 F
SAW 316IR/ER-316 ITT 5
l 75 BRP-SI-1-RB-037 G
SAW 316IR/ER-316 ITT 5
l 76 BRP-SI-1-RB-037 A
SAW 316IR/ER-316 IIT 6
l 77 BRP-SI-1-RB-037 B
SAW 316IR/ER-316 ITT 6
l C-7 l
,-r-- -,
.,--,-,-r---
,,,e-
TABLE C-1 (CONT)
NO.
LINE WEID HOCESS FILLER SOURCE SPOOL NO.
PIECE NO.
78 E P-SI-1-RB-037 C
GTAW ER-316 ITT 6
79 BRT-SI-1-RB-037 D
SAW 316IlyER-316 ITT 6
80 BRP-SI-1-RB-037 E
SAW 3161H/ER-316 ITT 6
81 EP-SI-1-RB-037 F
SAF 316IH/ER-316 ITT 6
82 BRP-SI-1-RB-039 FW6 GTAW ER-3161/ER-316 m
83 BRP"SI-1-RB-039 FW5 GTAW ER-316 BR 84 BRP-SI-1-RB-039 1-2A GIAW ER-308I/ER-308 BR 85 E P-SI-1-RB-039 1-1A GIAW ER-308I/ER-308 BR 86 BRP-SI-1-RB-039 FW1A GIAW ER308/ER-308L BR J7 BRP-SI-1-RB-039 13A GIAW ER316/ER-316L BR 88 E P-SI-1-RB-039 FW2 GIAW ER-316 BR 89 BRP-SI-1-RB-039 FW3 GL'AW ER-316 BR 90 BRP-SI-1-RB-039 IW4B GIAW ER316I/ER-316 BR 91 BRP-SI-1-RB-039 A
GIAW ER-316 ITT 1
92 BRP-SI-1-RB-039 B
GIAW ER-316 ITT 1
93 BRP-SI-1-RB-039 A
GIAN ER-316 ITT 2
94 BRP-SI-1-RB-039 B
SAW 316IR/ER-316, ITT 2
95 EP-SI-1-RB-039 A
96 EP-SI-1-RB-039 YO GIAW ER-316*
ITT 3
97 BRP-SI-1-RB-039 ZO GIAW ER-316 ITT 3
98 BRP-SI-1-RB-039 A
SMAW 3161H/ER-316/
E-316-16 ITF 4
99 BRP-SI-1-RB-039 B
SMAW 3161H/ER-316/
E-316-16 ITT 4
100 E P-SI-1-RB-059 FW4 GIAW ER-316/ER-316L BR 101 E P-SI-1-RB-059 3-1 GIAN ER-316/ER-316L BR 102 MP-SI-1-RB-059 FW3 GIAW ER-316/ER-316L BR 103 E P-SI-1-RB-059 FW2 GIAW ER-316/ER-316L BR 104 BRP-SI-1-RB-059 1-1A GIAW ER-316/ER-316L BR 105 M P-SI-1-RB-059 FW1 GIAW ER-316/ER-316L BR 106 BRP-SI-1-RB-059 A
GIAW ER-316 ITT 1
107 E P-SI-1-RB-059 B
SAW ER-316 ITT 1
108 BRP-SI-1-RB-059 C
SAW ER-316 ITT 1
109 BRP-SI-1-RB-059 D
SAW ER-316 ITT 1
110 BRP-SI-1-RB-059 E
SAW ER-316 ITT 1
111 BRP-SI-1-RB-059 F
SAW ER-316 ITT 1
112 BRP-SI-1-RB-059 G
SAW ER-316 ITT 1
113 BRP-SI-1-RB-059 A
SAW ER-316 ITT 2
114 BRP-SI-1-RB-059 B
SAW ER-316 ITT 2
- SOIUTION ANNEAIED C-8
[
l TABLE C-1 (CONF)
NO.
LINE WEID P30 CESS FIIIER SOURCE S100L NO.
PIECE No.
2 115 BRP-SI-1-RB-059 C
GIAW ER-316 ITF 116 BRP-SI-1-RB-059 D
SMAW ER-316/E316-16 ITT 2
117 BRP-SI-1-RB-059 E
SAW ER-316 ITT 2
118 BRP-SI-1-RB-059 F
SAW ER-316 ITT 2
119 BRP-SI-1-RB-040 FW6 GIAW ER-316I/ER-316 BR 120 BRP-SI-1-RB-040 FW5 GIAW ER-316I/ER-316 BR 121 BRP-SI-1-RB-040 FW4 GIAW ER-316 BR 122 BRP-SI-1-R;.r-040 FW3 GIAW ER-316 BR 123 BRP-SI-1-RB-040 FW2 GIAW ER-316 BR 124 BRP-SI-1-RB-040 FW1 GIAW ER-303 BR 125 BRP-SI-1-RB-040 A
GIAW ER-316 ITT 1
126 BRP-SI-1-RB-040 B
SMAW E316-16 ITT 1
127 BRP-SI-1-RB-040 A
GTAW ER-316 ITT 2
128 BRP-SI-1-RB-040 B
SMAW 316LH/ER-316/
E-316-16 ITT 2
129 BRP-SI-1-RB-040 A
SMAW 3161H/ER-316/
E-316-16 ITT 3
130 BRP-SI-1-RB-040 A
EAW 3161H/ER-316 ITI 4
131 BRP-SI-1-RB-040 B
SAW 316IH/ER-316 ITT 4
132 BRP-SI-1-RB-060 FW15 GrAW ER-308 BR 133 BRP-SI-1-RB-060 FW9 GTAW ER-316 BR 134 BRP-SI-1-RB-060 FWB GIAW ER-316I/ER-316 BR 135 BRP-SI-1-RB-060 8-1A GTAW ER-316I/ER-316 BR 136 BRP-SI-1-RB-060 FW16 GTAW ER-316I/ER-316 BR 137 BRP-SI-1-RB-060 1-1 GTAW ER-308 BR 138 BRP-SI-1-RB-060 10-1A GTAU ER-316'q/ER-316 BR 139 BRP-SI-1-RB-060 A
GTAW ER-316 ITT 1
140 EP-SI-1-RB-060 B
SAW ER-316 ITT 1
141 BRP-SI-1-RB-060 C
SAW ER-316 ITT 1
142 BRP-SI-1-RB-060 D
SAW ER-316 ITT 1
143 BRP-SI-1-RB-060 E
SAF ER-316 ITT 1
144 BRP-SX-1-RB-060 F
SAW ER-316 ITT 1
145 BRP-SI-1-RB-060 G
SAW ER-316 ITP 1
146 BRP-SI-1-hB-060 A
SAW ER-316 ITT 2
147 BRP-SI-1-RB-060 B
SAW ER-316 ITT 2
148 BRP-SI-1-RB-060 C
GIAW ER-316 ITT 2
149 BRP-SI-1-RB-060 D
SAW ER-316 IT1' 2
150 BRP-SI-1-RB-060 E
SAW ER-316 ITT 2
151 BRP-SI-1-RB-060 F
SAW ER-316 ITT 2
C-9
TABIE C-1 (CONT)
NO.
LINE WELD FIOCESS FILLER SOURCE SPOOL NO.
PIECE 10.
152 BRP-SI-1-RB-038A W6 GIAW ER-316/ER-316L BR 153 BRP-SI-1-RB-038A FW5 GIAW ER-316/ER-316L BR 154 BRP-SI -l-RB-038A W4A GTAW ER-316/ER-316L BR 155 BRP-SI-1-RB-038A FW3A GTAW ER-316/ER-316L BR 156 BRP-SI-1-RB-038A FW2 GIAW ER-316 BR 157 BRP-SI-1-RB-038A W1 GIAW ER-316 BR 158 BRP-SI-1-RB-038A A
GIAW ER-316 ITT 1
159 BRP-SI-1-RB-038A B
GTAW ER-316 ITT 1
160 BRP-SI-1-RB-038A A
GIAW ER-316 ITT 2
161 BRP-SI-1-RB-038A B
SAW 3161H/ER-316 ITT 2
162 BRP-SI-1-RB-038A A
SAW 316LH/ER-316 ITT 3
163 BRP-SI-1-RB-038A B
SAW 316IH/ER-316 ITT 3
164 BRP-SI-1-RB-038A A
SMAW ER-316/E-309-15/
E-316-16 ITT 4
165 BRP-SI-1-RB-038B FW9 GTAW ER-316/ER-316L BR 166 BRP-SI-1-RB-038B FW11 GIAW ER-316/ER-316L BR 167 BRP-SI-1-RB-038B FW17 GIAW ER-316 BR 168 BRP-SI-1-RB-038B 16-2 GIAW ER-308 BR 169 BRP-SI-1-RB-038B W10 GTAW ER-316/ER-316L BR 170 BRP-SI-1-RB-038B A
GIAW ER-316 ITT 6
171 BRP-SI-1-RB-038B B
SAW ER-316 ITT 6
172 BRP-SI-1-RB-038B C
GIAW ER-316 ITT 6
173 BRP-SI-1-RB-038B D
SAW ER-316 ITT 6
174 BRP-SI-1-RB-038B E
SAW ER-316 ITT 6
175 BRP-SI-1-RB-038B F
SAW ER-316 ITT 6
176 BRP-SI-1-RB-038B XR GTAW ER-316 ITT 6
177 BRP-SI-1-RB-038B YR GIAW ER-316 ITT 6
178 BRP-SI-1-RB-038B ZR GTAW ER-316 ITT 6
179 BRP-SI-1-RB-038B A
GTAW ER-316 ITT 5
180 BRP-SI-1-RB-038B B
SAW ER-316 ITT 5
181 BRP-SI-1-RB-038B C
SAW ER-316 ITT 5
182 BRP-SI-1-RB-038B D
SAW ER-316 ITT 5
183 BRP-SI-1-RB-038B E
SAW ER-316 ITT 5
184 BRP-SI-1-RB-038B F
SAW ER-316 ITT 5
185 BRP-SI-1-RB-038B G
SAW ER-316 ITT 5
186 BRP-RC-1-RB-026 FWS GTAW ER-316/ER-316L BR 187 BRP-RC-1-RB-026 FW4 GTAW ER-316/ER-316L BR 188 BRP-RC-1-RB-026 3-1 GIAW ER-316/ER-316L BR 189 BRP-RC-1-RB-026 FW3A GTAW ER-316/ER-316L BR 190 BRP-PC-1-RB-026 FW2 GIAW ER-316/ER-316L BR 191 BRP-RC-1-RB-026 FW1 GIAW ER-308 BR SMAW ER-308/E308 W
4/5 192 EPR-RC-1-RB-026 n
C-10
APPENDIX D HIGH STRESS LOCATIONS i
1 1
l l
l l
'IABIE D-1 CPSES-1 HIGI STRESS IDCATIONS KR MIIR7LT IBB ANAIXSIS LINE IDADING FORCES (lbs) 1GENIS (in-lbs)
CONDITIONS Fx Fy FZ Mr My Mz SIS 10-INGI DEADWEIGff 1300 2200 100
-29850 7470 28110 l
'IHERMAL 3420
-2590
-12090 39570 828630
-74620 SSE 22430 15060 14720 430670 466610 649630 SAM 140
-70 50 970 480
-3910
?
P RHR 12-INGI DEADWEIGff 30
-1970 10 3120 920 93380
'IHERMAL
-9400 1500
-3290
-171470
-2790 388700 SSE 23490 1093G 10640 359210 440970 686360 l
-120 20 140 1840 5940 930 l
l RCS 14-INGI DEADWEIQff 2370 2070
-60
-28840 49640 105510
'IHERMAL
-4470 15370 14360
-1568690
-896720 706880 l
-5490 7180 27340 413200 1272360 480810 SAM (see Note 1) 1 l
Note 1: 'Ihe surge line is dynamically coupled to the reactor coolant loop and the pressurizer and, therefore, soi-:mic anchor motions are included in SSE results.
l l
l
APPENDIX E LEAK DETECTION SYSTEMS Identified leakage is comprised of:
o Reactor head flange and valve packing leak-off that is captured and conducted to the reactor coolant drain
- tank, o
Reactor coolant leakage through steam generators, o
Pressurizer safety and relief valve leak-offs, o
Reactor coolant pump seal leak-off number 2, o
Reactor coolant pump seal leak-off number 3, and o
Leakage to the Component Cooling Water System from the RHR, SIS, and RCP thermal barriers.
For various forms of identified leakage, methods used to detect identified leakage include:
o The blowdown process sample monitor and the condenser vacuum pump gas monitor for the steam generator primary-to-secondary leakage, o
Surface mounted temperature detectors for the reactor head flange leak-off system and the pressurizer safety and relief valves, o
A flow totalizer downstream of the RCDT level control valve, and o
Component Cooling Water System (CCWS) radiation monitors and CCWS surge tank level increase.
A description of the primary unidentified leakage monitors is provided below along with a description of other. leakage indications available to the plant operator.
Containment Air Particulate Monitor -- Air particulate monitors take continuous air samples from the containment atmosphere and meeDare the particulate activity collected on a filter paper system.
After passing through an iodine and noble gas monitor downstream of the particulate monitor, the air returns to the Containment.
The sensitivity of the containment air particulate monitor to an increase in reactor coolant leak rate is dependent upon the magnitude of the normal baseline leakage into the Containment.
Sensitivity is greatest where baseline leakage is lowest.
Radioactive Gas Monitor -- The radioactivity gas monitor indicates the presence of containment gaseous activity originating from fuel-cladding defects.
It measures the gaseous beta radioactivity by continuously sampling the containment atmosphere.
The radioactive gas monitor is less sensitive to an increase in reactor coolant leak rate than the containment particulate monitor.
E-1
Containment Sump Flow Monitoring -- After collection in one of the two Containment sumps, the collected leakage is pumped to the floor drain tank outside the containment.
The combined sump pump discharge flow is recorded in the Control Room.
The sumps are also provided with level switches to alert the operator to 1 gpm or more leak to the sumps and to high level conditions in the event of sump pump malfunction.
Sump flow rates are determined and logged on a shiftly 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.
The sump discharge line may be sampled from outside of the Containment to provide additional aid in identifying the leakage source.
Specific Humidity Monitors -- Specific humidity monitors are sensitive to vapor originating from the reactor coolant, steam, feedwater, and auxiliary systems in the Containment.
Therefore, these monitors provide a means of detecting ur. identified leakage from both radioactive and non-radioactive sources.
Humidity detection is accomplished either by measuring the condensate from the Containment air cooling coils or by monitoring the dewpoint temperature in the Containment.
These methods are described as follows:
(1)
Condensate Flow Rate Measurement -- Humidity detection is accomplished by measuring the condensate flow rate from the Containment cooling coils.
The containment specific humidity increases proportionately with time and leakage until the dewpoint is reached at the Containment recirculation unit cooling coils.
If the specific humidity increases above this point, the heat removal needed to cool the air-steam mixture to its dewpoint temperature increases above this point.
Therefore, since the cooling coils are designed to remove heat at a constant rate, an increase in specific humidity results in increased condensate flow.
The condensate measuring system consists of a vertical standpipe with an internal self-siphoning device which empties the condensate in the standpipe to the sump when the standpipe is nearly full.
The condensate measuring systen permits measurement of the condensate flow rate from each Containment recirculation unit by means of a derivative unit which measures the rate of change in the standpipe level.
Should the leakage inside the Containment increase, the condensate flow E-2
also increases, thereby increasing the rate of change of the standpipe level.
The rate of level change in the standpipe is continuously recorded on strip chart recorders in the Control Room.
An alarm for high rate of level change is provided to warn Control Room personnel of an increase in the condensate flow rate.
An alarm is also provided if condensate flow is greater than the amount of flow that the siphon can discharge to the sump.
Through accurate measurements of condensate flow, a reliable estimate of the total leakage rate to the Containment can be made.
(2)
Containment Dewooint Monitors -- The containment humidity sensing system consists of dewpoint sensors, signal conditioning units, cabling, indicators, and recorders, all packaged in a system capable of the continuous, unattended, automatic operation for remote monitoring of the dewpoint of the Containment atmosphere.
Dewpoint sensors are strategically located in five reoresentativo areas of the Containment and are capable o'/ detecting and reading out a change of 1 F in dewpoint.
The signal conditioning units provide a linear output signal for transmission to a Control Room board-mounted analog indicator.
A separate output signal is provided to a Control Room board-mounted five point recorder.
Signal conditioning units have provisions for individually adjustable alarm contacts.
The range of the dewpoint temperature measuring system 0
is O to 100 F, Its accuracy is within 2 F.
Containment Temperature Monitors -- An increase in Containment temperature can indicate a leak of high temperature fluid from the RCPB or other high temperature systems.
Containment Pressure Monitors -- An increase in Containment pressure can indicate a leak of high temperature fluid from the RCPB or other high temperature systems.
Gross Leakace Indications -- Gross leakage in the Containment can be indicated by:
o Decrease in pressurizer level, o
Increase in the rate of supply of reactor coolant makeup water, Containment temperature monitors, o
o Containment pressure monitors, and o
Containment sump level high alarm.
Examples of both identified and unidentified leakage diagnosis techniques include:
Leakage occurring from the RV head to vessel closure o
joint is identified by a temperature rise in the leak-off line, E-3
o Leakage occurring from the RCPB is identified by a simultaneous rise in condensate and radioactivity monitor indications, o
Dewpoint temperature recordings assist operators in locating leakage points because of the various locations of the containment dew cells, Steam generator primary-to-secondary leakage is o
detected by the steam generator blowdown process sample monitor and the condenser vacuum pump gas moni-tor, o
Leakage of the reactor coolant to the CCWS is detectable by means of the CCWS radiation monitor, and o
The increased frequency of sump pump operation is also an indication of leakage as is the 1 gpm leak alarm in the control Room.
1 E-4
TABLE E-1 LEAK DETECTION SYSTEMS SENSITIVITY AND RESPONSE TIME
SUMMARY
REMARKS SYSTEM SENSITIVITY RESPONSE TIME
. Air Particulate
-Most gensitive 1 gpm in
-Ref nuclide Monitor
-5x10
'pci/cc
< 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> is Cs-137 Radioactive Gas 1 gpm in
-Ref nuclide Monitor
< 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> is Xe-133 Condensate
-very sensitive 1 gpm in
-best for Measuring
-0.1 to 10 gpm
< 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> leakage with System very low RC activity levels Dewpoint
-1*F or greater Temperature Monitors Containment
-alarms for 3" 1 gpm in
-higher leak Sump 1 gpm change in 45 45 minutes rates alarm Leak Alarm minutes
< 45 minutes E-5
. _. = - _ -, -.. - _,._.
TABLE E-2 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 Increast No No Humidity / Dew-Change Change point Temper-atures Containment Increase Increase Increase Increase Sump Flow rate Containment Increase Chromates l
Sump Chemistry Activity in Sample l
Analysis RCS Water Net No No No Inventory Loss change Change Change 1
l l
l l
l t
E-6 i
1
APPENDIX F CPSES-1 LEAK RATE CURVES l
l l
1 l
l i
l 1
l r
-~
r,-
r.-,--n,----
-en----
- - -, -,,, - ~,,, - -.,
12,00..........CPSES-1 SIS la.. INCH LINE, NORMAL LOAD LEAXAGE, CIRC
..........r..............................
1.0 L
E 8n 0
...................... )
i
)
R n0 m
p....................:.....................q..................y..>....................q.....................l a
,/
e f
E
,..f 4auQ.
..........................................}......r........................................................,
.=
.f P
./
p
,7 2a
.......................;.....................).............../.....>....................(.....................y...................4
.,f
.A.-
'i,00 2.00 4,00 00 CRACKLENGTH(INCHES) i Figure F-1 CPSES-1 Icak Rate Curve for 10-inch SIS Pipe, Circunferential Crack under Nornni Operating loads
12.00............CPSES-1 RHR 12-INCH LINE, NORMAL LOAD LEAXAGE, CIR
/
r
/
1, i
L
,i r
8.
i R
.t 1r
................>...............>...............)...............q...............(............,I...;.............../...............g g
i
./
m l
r E
/
l
,f 4,
...............<...............<................y...............)...............)...,s.........4................(..............<
/.
P M
2,
.................q...............q...............<.................y............>...............)...............).............i.
4
.J
,,,,,r-
-t. _
1 l
._s A
.00 2,00 4,'00 6,00 B,'00
~
CRACKLENGIH(INCHES) i t
I l
Figure F-2 CPSES-1 Irak Rate Curve for 12-inch RHR Pipe, Circunferential Crack mder Normal Operating Ioads i
I., an.................CPSES-1 SURGE LINE, NORMAL LOAD LEAXAGE, CIRCUM fa, H_ g.
l
,,t j
.i
./
1,................
j,
,f.
l a
j.
L.
.I f
I
,f 4
i l
)
A,.
a 6...............
,)
Q; i
y 31
}
s.
(
[
1
,c
..............4...............{................(................>...............:.....,...........)...............)..............4
.c 6
t
/
I W
i
,/
j 3
l 4s-..
...............)...............)..............
4...............<.........f.....<................,................................i A
t1 y
t.
=
.c 1
,J
_.............;................................p..........,a...q..............4.................................................;
sa s v
,, ',7 00 2,00 4,00 6,00 8,00 CRACHLENGTH(INCHES)
Figure F-3 CPSES-1 Leak Rate' Curve for 14-inch RCS Surge Line, Circumferential Crack under Normal Operating Loads
12aww............CPSES-1 SIS 10-INCHLINE,NORMALLOADLEAKAGE, AXIAL
........$....................g.....................p....................q.....................p.......,............g
[
[
I 1s
..........................................{.....................)..................../.....................).....
,i L
/
r r
./
.?
A 8e
......................)....................<.....................{.....................>..................j.4......................:
,/
f
.t R
n,n
,0
..............................a..........
- ~
/
i 4,0Q.....................'.....................'......................'....................:;f...................
[
s j
P l
f f
l e
2.
- b. r-
?""
w O'Y,00 2.'90 4.00 6.00
=
f I
f 8
f f
CRACKLENGTH(INCHES)
Figure F-4 CPSES-1 Icak Rate Curve for 10-inch SIS Pipe, Axial Crack under Normal Operating Ioads
12.M..............CPSES-1RHR12-INCHLINE,NORNALLOADLEAKAGE, AXIAL 1,
. i............
.i L
Or u,
R 6
.................{...............q...............<................>.......
.......>...............)...............)...............;
I tn E
4:
................)...............)...............{...............<..............<................>...............................l P
M 2e
................>...............>...............)........67si"..{...............q...............<.................................
.t
.t -
r t-kh' I)00 2.00 4.00 6,00 8'0 t.
0 4
CRACHLENGTH(INCHES)
~
~
~
Figure F-5 CPSES-1 leak Rate Curve for 12-inch RHR Pipe, Axial Crack inder
~~
Noum1 Operating Ioads m
12aa-................CPSES-1SllRGELINE,NORMALLOADLEAKAGE, AXIAL _
........................................................................................... n............
UU j
./
10.
...............................,z.................
).
E l
i
~
)
/
8
.........................y........,................
)
i f
R l
m
.................>...............>...............)...............{..............4.................:..!.............:................!
i i
f Q%
.c E
f.
e.
- 4. 8............... <.
.<................>...............)...............)...........'...q...............q................,
P
)
t
(
t 2a
................q...............q...............<................>..............,y.,
a.........)...............)...............(
1 f
4 5
6 5
4 3
6
.00 2.00 4,00 6.00 8,00 CRACHLENGIH(INCHES)
Figure F-6 CPSES-1 Icak Rate Curve for 14-inch RCS Surge Line, Axial Crack under Normal Operating loads
I-APPENDIX G ll MODIFIED LIMIT LOAD ANALYSIS BY THE MASTER CURVE METHOD Because the three
._?3-1 piping systems under consideration for WHIPJET leak-beforc break (LBB) analysis are entirely wrought austenitis stainless steel, it is appropriate to use a' modified limit load analysis to demonstrate acceptable LBB margins [1].
This appendix describes the methodology, called the Master curve approach, used in this modified-limit lead analysis.
This approach only applies to postulated circumferential cracks.
After the screening process for the LBB candidates has been successfully completed (Section 4), reasonable best-fit and lower bound material properties have been established (Appendix B), and the necessary leakage-size cracks have been poe#mlated (Section 7), the Master curve limit load analysis detailed below establishes the following two margins of safety:
first, that a margin of 2 exists between the leakage-size crack (10-gpm crack size under normal operating loads) and the critical crack size for the pipe when under normal plus seismic loads (i.e.,
L /L10 1 2; this is the flaw size margin) ; second, that a ratio Sk"at least one exists between the critical crdck size for c
a margin cr. loads of 1.4x(normal plus seismic) and the leakage size crack (i.e., L /L10gpm 2 1)*
c The Master Curve approach begins with the construction of a master curve where a stress index, SI, given by SI = S + MPm j
is plotted s a function of total through-wall flaw length, L,
defined by L = 2dR where 2af S=
- (2 sin
- sin #],
p
= 0.5 ((x-8) - x(P /ag)],
m l
e is the half angle in radiane of the postulated tnrough wall circ.erential flaw; R
is the pipe mean radius; P
is the combined membrane stress including pressure, m
deadweight, and seismic components (combined algebraically for this analysis);
G-1 Y_
M is a margin associated with the load combination method (in this case, M = 1.4 because of the algebraic load combination); and, og is the flov stress for austenitic steel material categories (42.8 ksi for base metal /TIG weld and 51.0 ksi for SAW/SMAW) (see Appendix B for flow stress determination).
If (e + p) is greater than w, then 2af S=
[ sing]
K where
=
x(P /#f) m For 'ach high stress node, therefore, two master curves are con,,ucted for each austenitic steel material category (one curve with or for base metal /GTAW and one curve with og for SAN /SMAW).
These curves are provided in this appendix as Figures G-1 through G-6 together with the specific values of the above variables upon which they are based.
After construction of the appropriate master curves, two Stress Indices are calculated for each of the three cases:
BASE METAL /GTAW, SAW, and SMAW.
The calculated Stress Indices are then used to enter the master curve of interest and find the allowable flaw lengths (critical crack sizes) used for flaw size margin and load margin calculations.
For the BASE METAL /GTAW case, the Stress Index is given by SI = M'(P,+ P )'
b where the combined membrane stress, including pres"ure, P
=
r~
daadweight, and seismic components (combined
)
algebraically);
Pb=
the combined primary bending stress, including deadweight, and seismic components (combined algebraically); and a margin, 1.4 for load margin calculations or 1.0 for M'
=
flaw size margin calculations.
G-2
~
l For both the SAW and SMAW cases, the Stress Index is given by SI = M'(Pm+Pb+P) Z' e
where P,
and M' are as defined above; P,
b m
combined expansion stress at normal operation' P
=
e (combined algebraically);
1.15(1.0 + 0.013(OD-4)] for SMAW:
Z
=
1.30(1.0 + 0.010 (OD-4)] for SAW; Z
=
and OD =
pipe outer diameter in inches.
Note the addition of the factor Z to the SAW/SMAW Stress Index equation.
The Z-factor is a conservative adjustment for the less ductile / tough flux welds, whose fracture is described by elastic-plastic fracture mechanics.
The Z-factor approach allows application of the modified limit-load approach (which assumes failure by plastic collapse of the cracked section without any crack extension) to the flux weld metal.
The validity of the Z-factor approach is established in Reference 2; Z-factors were determined from detailed elastic-plastic fracture analyses.
Note that for base metal and GTAW welds, Z = 1 which indicates that a limit load failure criterion (1) is being used; therefore, the terminology of crack instability includes either that of limit load plastic collapse or elastic-plastic fracture mechanics-based instability (using the Z-factor modification to limit load).
The following is summary of the modified limit-load (master curve) methodology applied at each high stress node:
1.
Assume BASE METAL /GTAW material properties.
- a. Construct the appropriate master curve.
- b. Enter the master curve with two Stress Indices as follows:
m + P ), and (1)
Let M' = 1.4, calculate SI = M'(P b
if the allowable flaw length from the master curve is at least equal to the leakage size flaw (L100pm), then the margin on 1.4 x (Normal + SSE) loads is met.
G-3
~
t
+ P ), and (2)
Let M' = 1.0, calculate SI = M'(P b
if the allowable flaw length from,the master curve is at least t'five d.c leakage size flaw (L10gpm)'
then the cargin on flaw size is met.
2.
Assume SAW/SMAW material properties,
- a. Construct the appropriate master curve (master curves for SAW and SMAW are identical).
- b. Enter the master curve with four Stress Indices as follows:
(1)
For Z(SAW), let M' = 1.4, calculate the value SI = M'(Pm+Pb + P ) Z, and if the e
allowable flaw length from the master curve is at least equal to the leakage size flaw (L10gpm), the the margin on load is met.
(2)
For Z(SAW), let M'
= 1.0, calculate the value SI = M'(Pm+Pb + P )Z, and if the e
allowable flaw length from the master curve is at least twice the leakage size flaw (L ogpm), then the margin on flaw size is me (3)
For Z(SMAW), repeat (1) and (2) above.
Results of the above methodology for CPSES-1 LBB lines are provided in Tables 8-1 and 8-2 of the basic report.
l REFERENCES I
1.
Standard Review Plan 3.6.3, Leak-Before-Break Evaluation Procedures, Section 10.1, Federal Reaister, August 28, 1987 (for public comment).
2.
- Norris, D.M.,
et al, Ey31uation of Flaws in Austenitic Steel _
PiDinc, NP-4690-SR, Special Report, EPRI, Palo Alto, California, July 1986.
3.
Norris D.M.,
and Chexal, B.,
PICEP:
PiDe Crack Evaluation Procram (Revision 3), NP-3596-SR, Revision 1, Specie.1 Report, EPRI, Palo Alto, California, December 1987, i
l l
G-4 l
S 15.00-SURGE LINE MASTER CURUE : BASE /GIAW T
u-i R
E
--FlowStress:42.8ksi i
S
~-M : 1.4 (Algebraic Load Conhination) i S
60.00
--Pipe Mean Rautts : 6.297 inches
-i
' s.
--ConhinedMenbraneStress: 4091.08 psi i
I 4
N
~~'.,
M 45.00 1.
)(
a l
tn P
30.00
'~.
-i
?
i 1
i i
15.00
~~
1 1
g
-_.4 0 A3 t
3 I80 5 '00 10l00 15.00 20'80; 25.00 CRACKLENGIH(inches)
Figure G-1 CF3ES-1 RCS Surge Line Master Curve for Base /Gtaw Material Properties
S SURGELINEMASTERCURVE:SAW/SMAW T
75.00r ri R
i
--FlowStress:51.0ksi i
E
--M : 1.4 (Algebraic Load Conhination) i S
S 60.00
~.
--PipeMeanRatitts:6.297 inches
~
s
~ --ConbinedMembraneStress: 4091.08 psi j
I j
~.
i n
i 45.00
-i
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t y
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n L
.P 30.00
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i
. ~
s.
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i 1
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l l
i 15.00 I
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i 0
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y e
8.ca 9.60 5.'00 10.'00 15.'00 20.'00 25.00 5
3 CRACKLENGIH(inches) i Figure G-2 CPSES-1 RCS Surge Line Master Curve for SAW/SMAW Material Properties
)
S 15.0B 10-INCH SIS MASTER CllRVE : BASF/GIAW T
r-i R
E
--FlowStress:42.8ksi i
S
--M : 1.4 (Algebraic Load Conhination) i S
60.00
--PipeMeanRaous:4.875 inches J
=...'..
--ConhinedMenbraneStress: 5265.07 psi 1
N i
D 45.00
-i E
i.
o a
s w
30.00
-i s
s.
.s 1
4 15.00
' s.
-i f
. ~ ~
0 y
a g.m 3
2 00 5.00 10.00 15.00 20.00 25.'00 CRACKLENGIH(inches)
Figure G-3 CPSES-110-inch SIS Line Master Cuwe for Base /GrAW Material Properties
S 75,gg.........................,...10-INCH SIS MASTER CURYE : SAWSMAW 7
R i
-s
--FlowStress:51.0ksi E
N..N.
--M:1.4 (AlgebraicLoadConhination) i S
S 60.00 N
--Pipe Mean Radras : 4.875 inches
-i N.
--ConhinedMenbraneStress: 5265.07 psi i
I
-s.
s N
N i
'\\
D 45.00 N
-i E
N i
X N.
i a
a P
30.00
\\.
s i
~~
N
~.
i
' s, 15.00
-i
.~~.
[
g E
~' s_
j
,, ~-
i 8.00 3
v.00 5.00 10.00 15.00 20.00 25.00 CRACXLENGIH(inches)
Figure C-4 CPSES-110-inch SIS Line Master Curve for SW/SMRJ Material Properties
S 15.0B-12-INCH RHR MASTER CURVE : BASE /GIAW T
E m.
i E
--FlowStress:42.8ksi i
i S
--M : 1.4 (Alcebraic Load Conhination) i S
60.00(~.
--PipeMeanRai.ius:5.8125 inches
-i
--ConbinedMenbraneStress: 5282.05 psi i
I H
's.
D 45.00 E
l
?
e P
30.00
-i
' ~.
s 1
1 i,
1500 i.
._~.
i 1
i 0
y l
4 x
g,e.
3 t.00 5.00 10.00 15.00 20.00 25.00 CRACXLENGIH(inches)
Figure G-5 CPSES-112-inch RHR Line Master Curve for Base /Gr/R Material Properties 4
S 75.00 m 12-INCH RHR MASTER CURYE : SAW/SMAX T
R E
--FlowStress:51.0ksi i
- s S
N.
-M:1.4 (AlgebraicLoadCoMbination) i S
60.00 N.
--Pipe Mean T,thus : 5.8125 inches
-i
.N
--ConhinedMembraneStress: 5279.86 psi i
I
\\.
i N
N.
D 45.00 N.
-i E
s X
a s
f P
30.00
\\,.
s i
N.
i
-s_
15.00 s_-s 1
~ ~ _
i 0
~_
i 9.0B 3
W00 5.9'0 10'00 15'90 20'00 25.00 CRACXLENGTH(inches)
Figure G-6 CPSES-1 12-inch MIR Line Master Curve for SAH/SMAW Material Properties