ML20151M344
| ML20151M344 | |
| Person / Time | |
|---|---|
| Site: | Summer  |
| Issue date: | 06/30/1988 |
| From: | WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19292J188 | List: |
| References | |
| SG-88-07-009, SG-88-7-9, WCAP-11858, NUDOCS 8808050117 | |
| Download: ML20151M344 (94) | |
Text
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WESTINGHOUSE CLASS 3 WCAP-ll858 SG-88-07-009 i
t TUBESHEET REGION TUBE ALTERNATE PLUGGING (L*)
CRITERIA FOR STEAM GENERATORS IN THE V. C. SUMMER NUCLEAR STATION WESTINGHOUSE ELECTRIC CORPOR'tTION NUCLEAR SERVICE DIVISION P. O. BOX 3377 PITTSBURGH, PENNSYLVANIA 15230 JUNE, 1988 Work Ferformed Under Contract:
Purchase Order No. W593071 l
W General Order No. CH-61370 l
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i 8808050117 seoso; !
DR ADOCK 05000395 l PDC I 1 l t
024W:49/0719M l l
ABSTRACT An evaluation was performed to develop alternate plugging criteria, for determining whether repairing or plugging of full depth hardroll expanded steam generator tubes is nocessary for degradation that has been detected in the portion of the tube located within the tubesheet. Criteria based on adequate axial strength and potential leakage (with the applicable length designated as L*) were developed. An L* criteria value of 0.50 inch was established as sufficient for continued plant operation for certain types of tube degradation below the L* distance. The evaluation demonstrates that application of the L*
criteria for indications of tube degradation within the roll expansion affords a level of plant protection commensurate with that provided by Regulatory Guide 1.121 for degradation located outside of the tubesheet region.
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TABLE OF CONTENTS t
- IDflG E8GE i
1.0 BACKGROUND
10 2.0 TECHNICAL APPROACH 13
- . 3.0 TEST FOR THE DETERMINATION OF RESISTANCE TO 19 LEAKAGE OF DEGRADED ROLL EXPANSIONS 4.0 CRITERION AND TEST RESULTS FOR LEAKAGE FROM DEGRADED 28 i
ROLL EXPANSIONS 5.0 TEST FOR THE DETERMINATION OF AXIAL LOADBEARING 39 CAPABILITY OF DEGRADED ROLL EXPANSIONS
- 6.0 CRITERION AND TEST RESULTS FOR THE AXIAL LOADBEARING 49 CAPABILITY OF DEGRADED ROLL EXPANSIONS
- 7.0 PULLOUT LOAD REACTION LENGTH 85 j 8.0
SUMMARY
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9.0 REFERENCES
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FIGURES NO. TITLE PAGE
. 1-1 Configuration for Tubesheet Region L* Alternate 12 Plugging Criterion for Roll Expanded Model D Steam Generator Tubes-Single Band Degradation 2-1 Configuration for Tubesheet Region L* Alternate 17 Plugging Criterion for Roll Expanded Model D Steam Generator Tubes-Multiple Band Degradation 3-1 Roll Expansion Leakage Tesc Sample for Model D 24 Steam Generator Normal Operation and FLB Conditions 3-2 Roll Expansion Leakage Test Sample for Model D 25 Steam Generator LOCA Conditions 4-1 Normal Operation Leakage Test: ~ Leak Rate Per 33 Degraded Roll Expansion vs. X 4-2 Feedline Break Leakage Test: Leak Rate Per 34 Degraded Roll Expansion vs. X 4-3 . LOCA Leakage Tert: Leak Rate Per 35 Degraded Roll Expansion vs. X 4-4 Number of Tube Ends per Steam Generator 36 Depositioned by L* - Normal Operation 4-5 Number of Tube Ends per Steam Generator 37 Despositioned by L* - Feedline Break 4-6 Number of Tube Ends per Steam Generator 38 l Depositioned by L* - LOCA 5-1 Geometry of Slots in Collared Tube Specimen - 43 Ultimate Strength Test 5-2 Geometry of Slots in Decollared Tube Specimen - 44 Ultimate Strength Test 5-3 Geometry of Slots in Unexpanded, Nevu-Collared 45 Tube Specimen - Ultimate Strength Test 5-4 Sketch of Zetac Tension and Compression Testing 46
. Machine 4
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FIGURES (Cont'd)
NO. TITLE PAGE 6-1 Pull Force as a Function of Pull Displacement for a 55 Collared, Nondegraded Tube (Specin:en No.11) ;
6-2 Pull Force as a Function of Pull Displacement for a 56 Collared, Nondegraded Tube (Specimen No. 16) -
Replication of Specimen No.11 6-3 Pull Force as a Function of Pull Displacement for a 57 ;
Collared Tube (Specimen No. 3), with (15)-30 Degree Slots, Tops of Slots 0.'54 Inch Below the Bottom of Roll Transition .
6-4 Pull Force as a Function of Pull Dr s p : n ement for a 58 Collared Tube (Specimen No. 8), wiUi (15)-30 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition - Replication of Specimen No. 3 6-5 Pull Force as a Function of Pull Displacement for a 59 Decollared Tube (Specimen No.12), with (15)-0 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition 6-6 Pull Force as a Function of Pull Displacement for a 60 Decollared Tube (Specimen No.17), with (15)-0 Degree .
Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition - Replication of Specimen No. 12 6-7 Pull Force as a Function of Pull Displacement for a 61 Deco 11ared Tube (Specimen No. 13), with (15)-15 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition 6-8 Pull Force as a Function d Pull Displacement for a 62 Deco 11ared Tube (Specimen th.18), with (15)-15 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition - Replication of Specimen No. 13
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6-9 Pull Force as a Function of Pull Displacement for a 63 Decollared Tube (Specimen No.1), with (15)-30 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition 6-10 Pull Force as a Function of Pull Displacement for a 64 Decollared Tube (Specimen No. 7), with (15)-30 Degree -
Slots, Tops of Slots 0.25 Inch Below the Bcttom of Roll Transition - Replication of Specimen No.1 ,
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FIGURES (Cont'd)
N0. TITLE PAGE
. 6-11 Pull Force as a Function of Pull Displacement for a 65 Decollared Tube (Specimen No. 2), with (15)-45 Degree Slots, Tops or Slots 0.25 Inch Below the Bottom of Roll Transition 6-12 Pull Force as a Function of Pull Displacement for a 66 Decollared Tube (Specimen No.14), with (30)-0 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition 6-13 Pull Ferce as a Function of Pull Displacement for a 67 Decollared Tube (Specimen No. 19), with (30)-0 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition - Replication of Specimen No. 14 6-14 Pull Force as a Function of Pull Displacement for a 68 Decollared Tube (Specimen No. 15), with (30)-15 Degree Slots, rops of Slots 0.25 Inch Below the Bottom of Roll Transition 6-15 Pull Force as a Function of Pull Displacement for a 69 j Decollared Tube (Specimen No. 20), with (30)-15 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition - Replication of Specimen No.15 6-16 Pull Force as a Function of Pull Displacement for a 70 Decollared Tube (Specimen No. 4), with (30)-30 Degree I Slots, Taps of Slots 0.25 Inch Below the Bottom of i Roll Transition )
6-17 Pull Force as a Function of Pull Displacement for a 71 Decollared lube (Specimen No. 9), with (30)-30 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition - Replication of Specimen No. 4 6-18 Pull Force as a Function of Pull Displacement for a 72 Decollared Tube (Specimen No. 5), with (30)-45 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition l 6-19 Pull Force as a Function of Pull Displacement for a 73 Never-Collared / Expanded Tube (Specimen No. 6), with ,
(30)-30 Degree Slots '
6-20 Pull Force as a Function of Pull Displacement for a 74
. Never-Collared / Expanded Tube (Specimen No.10), with (30)-30 Degree Slots, - Replication of Specimen No. 6 6
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FIGURES (Cont'd)
NO. TITLE PAGE 6-21 Ultimate Pull Force as a Function of Slot Angle for 75 '
Expanded, Collared and Decollared Tubes,15 Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition 6-22 Ultimate Pull Force as a Function of Slot Angle for 76 Expanded, Decollared Tubes, 30 Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition; Never-Collared, Non-expanded Specimens Also Shown 6-23 Ultimate Pell Force as a Function of Slot Angle for 77 Expanded, Collared Tubes, With and Without Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition 6-24 Ultimate Pull Force as a Function of Slot Angle for 78 Expanded, Decollared Tubes, Tops of. Slots 0.25 Inch Below the Bottom of Roll Transition; Never-Collared, Non-expanded Specimens Shown for Comparison 6-25 Ultimate Pull Force as a Function of Number of Slots at 79 0,15, 30, and 45 Degrees for Expanded, Collared and Deco 11ared Tubes, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition 6-26 Ultimate Pull Force as a Function of Slot Angle for 80 Collared and Decollared Tubes, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition; Never-Collared -
Specimens and Pull Strength Model Also Shown 6-27 Model for Plastic Collapse 81 6-28 Comparison of Load-Displacement Records, Computed 82 vs. Measured for 30 Slots at (=45' 6-29 Comparison of Load-Displacement Records, Computed 83 vs. Measured for 30 Slots at p=30' 6-30 Model D Steam Generator Degraded Tube Pull 84 Strength Design Curve for L* '
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TABLES i
N0. TITLE PAGE 2-1 Technical Approach to Meet Regulatory Requirements 18 3-1 Roll Expansion Leakage Test Results 26-27 5-1 Selection of Bounding Features for Axial 47 Loadbearing Test of Degraded Roll Expansions 5-2 Axial Loadbearing Strength Tests of Degraded 48 Roll Expansions 7-1 Calculation of Pullout Load Reaction Length 87-90 and Tube Length to be RPC Inspected
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NOMENCLATURE (Terms are listed in the approximate order of appearance in the report)
. RT Roll transition BRT Bottom of roll transition RE Roll expansion N. Op. Normal Operation LOCA Loss of Coolant Accident E Indication length TS Tubesheet T/TS Tube to tubesheet ECT Eddy current test X Distance from top of indication to BRT n Number of linear indications per band p Linear indication inclination angle from the tube axis TTS Top of tubesheet RPC Rotating pancake coil S/G Steam generator
. ID Inside diameter l OD Outside diameter I
- dpm Drops per minute ECI Eddy current indication LTL Lower tolerance limit FLB Feedline break SLB Steamline break sp T/TS interfacial radial contact pressure T Torque MA Mill-annealed SRE Sound roll expansion PLRL Pullout load reaction length SBD Single band degradation N Number of sound roll expansion portions 9
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TUBESHEET REGION TUBE ALTERNATE PLUGGING (L*)
CRITERIA FOR STEAM GENERATORS IN THE V. C. SUMMER NUCLEAR STATION
1.0 BACKGROUND
Existing plant Technical Specification tube repairing / plugging criteria which have been applied throughout the tube length do not take into account the reinforcing effect of the tubesheet (TS) on the external surface of the expanded portion of the tube. The presence of the TS will constrain the tube and will complement tube integrity in that region by essentially precluding tube deformation beyond the expanded outside diameter. The resistance to both tube rupture and tube collapse is significantly strengthened by the TS. In addition, the proximity of the TS significantly affects the leak behavior of throughwall tube cracks in this region. Based on these considerations, the establishment of alternate plugging criteria specific to the roll region of the tubes is justified. l The roll expanded length of the tube can be considered to consist of two zones, j the roll transition (RT) region and the roll region. The roll transition l region is defined as that portion of the tube where the roll expanded length transitions to the unexpanded length. The RT is approximately
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[ ]a,c,e in, long, axially. The bottom of the roll transition is located in the vicinity of the top of the tubesheet (TTS). The roll region of the tube I is that length between the bottom of the roll transition and the tube end at the bottom of the TS. The roll region is also referred to as the roll expansion (RE) or hardroll. Refer to Figure 1-1. ,
l 1.1 F* CRITERION l
Previously approved for use as a alternate plugging criteria in the V. C.
Summer steam generators is the F* criterion. The F* criterion permits operation with any amount of tube degradation below a distance F* equal to 1.6 inches belcw the bottom of the roll transition (BRT). However, it also requires no indication of degradation within the F* distance. The F* distance was analyzed to provide sufficient frictional force between the expanded tube i 10 0245H:49/071388
and tubesheet to resist pullout due to normal operation and postulated accident conditions. The use of the F* criterion does not require any assessment of I tube degradation other than elevation and any type of degraoation below the F*
distance is acceptable. The limitation of the F* criterion is that it does not address eddy current test (ECT) indications found in the top 1.6 inches of the .
roll expansion.
1.2 L* CRITERIA To address eddy current indications (ECI's) of tube degradation in the top portion of the roll expansion an additional alternate plugging criteria designated the L* criteria has been established. This report is written to document development of the L* criteria. This criteria is based on the fact that leaks from throughwall tube degradation will be substantially reduced by a length of tube expandea into the tubesheet shorter than the length required for structural integrity. Also, some types of potential tube degradation, specifically axia, or near axial cracks within the tubesheet region, will provide the required structural integrity even when the cracks have propagated throughwall. Use of the L* criteria will require more definition of the tube degradation and probably more sophisticated eddy current inspection techniques .
than required for the F* criterion but will address a significant number of the tubes which do not qualify for the F* criterion. Finally, a combination of the ,
L* and F* analysis methods can be shown to address the existence of tube degradation in the form of axial ECI's at a single elevation or at multiple elevations in the tube near the top of the roll expansion.
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CONFIGURATION FOR TUBESHEET REGION
. L* ALTERNATE PLUGGING CRITERIA' FOR ROLL-EXPANDED STEAM GENERATOR TUBES SINGLE BAND DEGRADATION a.,c,e 9
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. Figure 1-1 12 0245H:49/071388
1 2.0 TECHNICAL APPROACH l
2.1 GENERAL .
The proposed roll regibn plugging criteria has been evaluated for the four tube modes recommended to be considered by NRC Regulatory Guide 1.121, for three steam generator conditions. This evaluation included a combination of tests and analyses, outlined in Table 2.1. The pull and proof tests in the table are documented in Ref. 1. The degraded-RE collapse issue is also addressed in Ref. 1. (Note that Ref. 2 is the non-proprietary version of Ref. 1 and is included for information only.)
The eddy current indications in the roll region of the V. C. Summer steam generators (S/G's) can be conservatively characterized as one of two limiting types of degradation and addressed with the F* and L* tube alternate plugging ;
criteria. This report is written to support the incorporation of the L*
criteria into the technical specification.
2.2 F* CRITERION In the alternate plugging criterion previously approved for use at
. V. C. Summer, eddy current indications were treated as throughwall, 360* l circumferential indications and it was shown that, if the degradation occurred below a distance, F*, below the BRT of the tube-to-tubesheet (T/TS) joint, l operation is expected to be safe and reliable. The distance from the BRT to )
the indication top was determinable by standard bobbin coil ECT data, within a i i
small but conservative tolerance. The F* tube plugging criterion is based on a j semi-empirical method of quantifying the axial loadbearing capability of the rolled joint, resulting from the radial contact preload pressure and the associated friction between the tube and TS. The presence of the T/TS radial l pressure, s p, which consists of the as-manufactured pressure and as-modified by operating loads and temperatures, also causes significant resistance to the leakage of primary-to-secondary and secondary-to-primary water. It has been determined by previous tests (Ref. 1) that sound RE's, of F* 1ength, are
. essentially leaktight. The use of the F* criterion obviates the necessity of I 13 024 W 49/071388
l determining the ECI depth, number, inclination, length and circumferential spacing. Only the distance of the ECI uppermost extent to the BRT needs to the ,
determined. In short, the nature and extent of tube degradation need not be determined. Refer to Figure 1-1.
2.3 L* CRITERIA 2.3.1 LEAKAGE The selection of an L* 1ength has as a primary consideration the goal of minimizing significant leakage from tubes which have been accepted using the L*
criteria. Significant leakage would be an aggregate leakage from the L* tubes which would cause a shutdown or operation at reduced power due to safety, regulatory, or operational reasons. Determination of the leak rate through the interface between the expanded tube and the tubesheet is not conducive to calculational methods. The resistance-to-leakage (RTL) of the T/TS interface extending over a portion of the sound roll expansion portion (SRE) immediately t'elow the BRT was determined by test. For some tube joint ECT data analyzed, more than one degradation array, or degraded roll expansion (DRE), per joint was found. Refer. to Figure 2-1. The uppermost DRE would be expected to exhibit the most stringent, i.e., highest, leakage because the uppermost DRE, in the limit, has no T/TS radial contact pressure and forms a manifold which ,
intercepts all leakage flow , if any exists, in the T/TS interface. The uppermost SRE is also farthest removed from the TS neutral axis. Therefore, it undergoes the greatest reduction in ps , and therefore in reduction of resistance to leakage during the most stringent i.e., upward, bending mode of the TS. All of the other SRE's above the TS neutral axis have larger s r's and therefore lower leakage in this mode.
2.3.2 STRENGTH Determination of the loadbearing capability of DRE's was based on the limiting i assumption that the L* region recommended length of 0.50 in., plus ECT +
uncertainty, provides no axial fixity. Therefore, the most atringent axial loads, viz., pullout heads, were assumed to be applied directly to the ,
uppermost DRE. The criteria also considers the degradation to be axial or to 14 0245H:49/071388
have an inclination angle, 4, of up to [ ]*,a,c,e and to be throughwall. In this case, the minimum axial length of sound, undegraded RE between the BRT and the top of the highest-elevation degradation, designated as L*, for operation with acceptable leakage was determined. In this criteria,
. the tube axial loads are borne by the sound portion of the tube above L*, the uppermost degraded portion, the sound portions, other axial or netr-axial degradation portions, if any, below the uppermost degraded portion and extending to the T/TS weld. Determination of the characteristics, e.g., ECI number, inclination angle, etc., of all of the degraded regions in a given tube joint would be necessary if the tube pullout load were to be applied to the l T/TS weld. However, this load need not be assigned to be reacted by the SRE within the L* distance and the weld. Instead, it can be reacted by the appropriate aggregate length of SRE's of the tube below the uppermost degradation array. Below this pullout load reaction length (PLRL), the degradation need net be quantitated to the same extent as within the length; it need be quantitated only to the same extent that degradation is quantitated below the F* distance per Ref.1. In the most stringent case, these SRE's are interspersed with DRE's and several interspersed SRE's may be assigned to react the pullout load. Part of the pullout load is reacted by the SRE within the L*
distance. Each DRE must transmit part of the pullout load to the SRE(s) and DRE(s) below it; the pullout load to be transmitted decreases with decreasing
. elevation. Therefore, the required DRE strength could also reduce with elevation. However, for the sake of simplicity, all DRE's were treated the same. Calculation of the holding force from each SRE was performed in the same way that the holding force was calculated for the F* distance in Ref.1. This involved calculation of the axial frictional force, preventing pullout, as the product of the T/TS radial contact pressure (s r) of the area of the SRE and i the T/TS static coefficient of friction. The reduced holding force of the two ends of each of the SRE's was also considered. Therefore, it was concluded that the configuration involving multiple bands of DRE's could be determined by calculation and did not need not be tested.
1 Thus far in this discussion, the strength aspects of the uppermost portion of the T/TS joint has involved strictly axial pullout loads on the tube. Other l
. loads, such as axial compression, bending and torsion about the tube vertical axis were considered but were judged to be negligible. For example, the axial 15 0245M:49/071388
l compression, downward, load occurring during large break LOCA, hereinafter referred to in this report as LOCA, was judged to affect the tube at the upper SRE to an insignificant extent. If the T/TS interface within the L* distance '
were not leaktight and permitted the LOCA pressure to penetrate to the DRE(s),
radial compression would act on the DRE(s). However, in Ref. 1 is discussed ,
the within-TS tube collapse strength characteristics in the presence of axisl and circumferential indications. It was pointed out that significant margin exists between tube collapse strength and the limiting secondary-to-primary pressure differential (LOCA).
The effects on the tube in the vicinity of the TTS, of other applied loads such as tube static and dynamic bending were also considered. Static bending occurs as a result of TS bending during heatup to the N. Op. condition. Dynamic bending results from the small-amplitude vibration of the tube span between the TTS and first support plate. In general, tube bending loads at the TTS, i.e.,above the BRT, and the resulting axial stress are low. This conclusion is based on the lack of circumferential indications in this area. The manifesta' tion of excessive bending stress above the BRT, circumferentical indications, have not been reported for these S/G's. Similarly, circumferential indications have not been reported below the BRT, in the ,
intended L* region. This indicates that even if bending, i.e., axially tensile, stresses were significant in the intended L* region, the tube is ,
sufficiently compressed by the interference fit with the TS to prevent circumferential degradation related to tensile stresses there.
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CONFIGURATION
, FOR TUBESHEET REGION
. L* ALTERNATE PLUGGING CRITERIA FOR ROLL-EXPANDED STEAM GENE"ATOR TUBES MULTIPLE BAND DEGRADATION a.,c,e I
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- Figure 2-1 17
.. *'"# #"* ?** - . - - - _ . . . - . . _ . . __ - , . _ . -.
lABLE 2-1 TECHNICAL APPROACH TO MEET REGULA10RY REQUIREMFNTS LEAK POSTULATED TUBE TUBE TUBE BEFORE E RUPTURE LEAK COLLAPSE BREAK CONDITION Normal Operation o Preload Analysis o Preload Analysis o Pull Tests o Proof Tests N/A '
o Proof Tests o Leak Tests o Preload Analysis
'FLB + SSE o Preload Analysis o Preload Analysis I o Leak Tests o Pull Tests o Proof Tests N/A o Proof Tests o Leak Tests 00
- LOCA + SSE o Preload Analysts o Limit Analy.
N/A o Leak Tests o Existing. N/A
- Test Data
- FLB: Feedline Break (Note: FLB is most stringent Faulted Condition); SSE: Safe Shutdown Earthquake
- LOCA: Loss of Coolant Accident i
0245M/071388:49
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3.0 TEST FOR THE DETERMINATION OF RESISTANCE TO LEAKAGE OF DEGRADED ROLL I EXPANSIONS 1
3.1 BACKGROUND
. l There is generally a correlation, for a given tube-to-tubesheet joint design, between tube pullout strength, i.e., force, and tube roll expansion thinning; l there is a desire ( thinning range for maximum strength. The desired thinning is achieved by the use of a certain roll expansion tool operated in a given l stalling torque range. This facet of power plant heat exchanger design decreased in importance approximately two to three decades ago, with the j widespread use of welded-and-rolled joints because of the addition of the weld to the structure. The RTL of a given roll-only design was generally consistent with the maximum strength requirement for the design. However, the RTL of the expanded portion of a welded-and-rolled design typical of nuclear steam generators has not been extensively studied because of the very good sealing propertios of the weld. Therefore, the RTL of short lengths of this design was not available and a test was designed to determine the RTL of the V. C. Summer Model D Steam Generator rolled joint design.
3.2 OBJECTIVE The objective of this test was to determine tube-to-tubesheet joint resistance to leakage for the L* case defined in Section 2.0. The tests were performed with T/TS test specimens fabricated of short sections of prototypical tubes rolled into collars which provided the same structural compliance as a unit cell of the TS. The tests were performed at prototypical pressures, temperatures and tube axial loads for normal operation (N. Op.), the most stringent faulted primary-to-secondary condition, i.e., feedline break (FLB),
and the most stringent faulted secondary-to-primary condition, Loss of Coolant Accident (LOCA).
The effect of prototypical loads on joint leakage was addressed in the leak test. For example, a load such as primary-to-secondary pressure differential during N. Op., causes direct pressure effects on the tube in the L* region to i
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increase sp . It also causes tubesheet bending. The differential pressure increases the T/TS radial contact pressure for all tubes in the TS, peripheral as well as interior, i.e., away-from-periphery, tubes. However, the upward bending also causes a decrease in radial contact pressure between tube and tubesheet for interior tubes and therefore a decreased RTL. This bending
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effect, which is detrimental for interior tubes, can be accounted for in the laboratory.
The acceptable leakage for a reasonable number of tube joints was established for the N. Op. and faulted conditions. It is obvious that for all acceptable-leakage cases, the joint must also exhibit acceptable strength. It was concluded that the (L*) strength testing could be decoupled from the leakage testing and the strength limits were determined by a separate test.
The primary to-secondary pressure differential effect and a secondary-to-primary pressure differential effect were expected to have negligible impact on the strength test results. However, prototypical radial contraction or expansion of the tube, causing a reduction or increase in the T/TS radial contact pressure (sr ), was achieved in the leakage test.
3.3 TEST EQUIPMENT
- 1. Tubes, Inconel 600 (1600) mill annealed (MA), [ ]a,c,ein. ,
wall (nominal).
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- 2. Tubesheet Simulants (Collars): (1) Roll Torque Reduction Test - Cold Rolled Carbon Steel, AISI 1018, 4.5 in. long and of [ '
]a,c,e in. 00 (nom.) (2) Leakage Test - all parameters to be the same as for the roll torque reduction test except the collar length to be 7.0 in. Refer to Figures 3-1 and 3-2.
- 3. Airetool Roll Expander, No. [ ]a,c,e rolls per original l (shop) fabrication specification.
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- 5. Laboratory leak test equipment such as pressurizing systems, furnaces, etc.
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. 3.4 TEST MAJOR STEPS 3.4.1 Determine roll torque reduction to simulate the effect of TS bending on reduction of tube-to-tubesheetp s for interior tubes.
3.4.1.1 Apply (hoop) strain gages to collar ODs.
1 3.4.1.2 Roll expand tubes to achieve minimal, T/TS radial contact pressure, i.e., interference fit. Record the corresponding hoop strains and I calculate T/TS srby means of Lame's thick-walled cylinder I relationships.
3.4.1.3 Roll expand at successively higher stalling torques i.e., thinning, and calculate respective sp values.
3.4.1.4 Analyze data to determine torque reduction to account for lack of TS bending effect for laboratory samples. Determine torque j
- reduction for N. Op. and FLB. The LOCA torque will be the mid-range torque used during shop fabrication of the S/G. Although TS downward (LOCA) bending increases the T/TS contact pressure for interior tubes, it doesn't affect the pressure for the most stringent case, i.e., for peripheral tubes.
3.4.2 Fabricate Leak Test Samples with [ ]a,c,e in, of sound roll between the BRT and the DRE. This distance is referred to as "X". Each sample has a separate "X". Refer to Figure 3-1. For X
= 0.5 In., the desired length of sound RE above the degradation, concentrate the most samples. An X of [ la,c,e In. is important for trending purposes only; X values of [ ]a,c,e In, were expected to be leaktight or to have negligible leakage.
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3.4.2.1 Estimate collar ID surface finish in area of leak path. During the RE process, the tube becomes fully plastic and flows into very small depressions occurring on the S/G tubesheet hole surface, or on the tubesheet siniulant (collar) hole surface in this case.
Conceptually, the amount of tube flow, determined by hole surface '
roughness, may be related to joint RTL. Ascertain that the collar finish is on the order of the prototypical finish, i.e.,
[ ]a,c,e RMS.
3.4.2.2 Fabricate T/TS samples per Figures 3-1 and 3-2. Set roll expansion motor torque to [ ]a,c,ein-lb.
torque range specified during S/G manufacturing, as reduced to account for TS upward bending. The torque reduction from the
[ ]a,c.ein-ik. shall correspond to the s caused r by TS ,
bending, as 1.isted in Ref. 1.
3.4.2.2.1 Partial RE (non-interference fit). Remove tube from collar.
3.4.2.2.2 Drill holes in tube to simulate DRE crack top tips. Reinsert tube in collar. .
3.4.2.2.3 Finish (hardroll) RE at [ ]a,c e in-lb. torque. ,
3.4.2.3 Heat T/TS samples in furnace at [ ]a,c,e hours to duplicate the most stringent conditions caused by post-weld heat treatment of the channelhead-to-TS weld during shop fabrication. '
3.4.2.4 Leak test the FLB primary-to-secondary configuration with [
Ja,c,e in. of sound roll between the BRT and the degraded portion of the joint. Refer to Figure 3-1. The limiting case for the degradation is a 360' circumferential, throughwall "crack",
machined in the tube. However, because this machining may loosen the tube for small X values and therefore possibly bias the testing, a better method is to use the discrete-hole approach. The number of holes [ ]a,c,e is chosen to be large, to approximate ,
or exceed the number of linear ECI's observed in these S/Gs. Refer l
l 1
22 0245M:49/071888 l
i I
l to Table 3-1. After the FLB samples are tested, convert the samples to the N. Op. configuration by rerolling to the higher, N. Op., torque i.e., [
]a,c,e hours and leak test at the N. Op. conditions.
3.4.2.5 Leak test the secondary-to-primary (LOCA) configuration using the same sequence as used in primary-to-secondary configuration above.
The LOCA samples are completely separate from the FL8/N. Op.
samples. Refer to Figure 3-2.
1 I
3.5 TEST FACILITY I The tests were performed at the Westinghouse Research and Development !
Laboratory. A four-cell manifold was used to apply pressurized, deionized water simultaneously to the tube ID's of four specimens for j primary-to-secondary tests; i.e., the N. Op. and FLB conditions, and to the ;
tube 00's for secondary-to-primary tests, i.e., the LOCA condition. System pressure was measured by a transducer system with a 5 psi accuracy. The electrical resistance h.aters and inw1ation were positioned on the T/TS specimen such that the temperature acting over the leak path and ending at the l BRT, the tcp end of the leak path in the plant, was [ ]a,c,e.F for the l
- N. Op and FLB conditions. Above the BRT, the temperature dropped as designed by placement of the insulation so that any leakage was captured and condensed as droplets. The droplets were manually counted during the test period, usually 60 minutes. For the LOCA tests, the temperature acting over the leak path and beginning at the BRT was [ ]a,c,e.F. In this case, all leakage was captured from the tube ID and condensed as droplets and counted manually during the 60 minute test period.
3.6 TEST PROCEDURE l l
The test was performed in accordsnce with a written procedure, a referenceable document, under Qitality Assurance (QA) surveillance. The tube specimens were fabricated from QA-controlled stock; the collars were fabricated of i
. known-strength AISI 1018 cold rolled carbon steel and were also fabricated under QA surveillance.
I 23 0245M:49/071188 l
-, -. , -- -. .I
FIGURE 3-1 ROLL EXPANSION LEAKAGE TEST SAMPLE
, FOR MODEL D STEAM GENERATOR NORMAL OPERATION AND FLB CONDITIONS
%C>C l l
l I
H0_If1:
- 1. Primary Side source, dionized water, subcooled
- 2. Step roll 4 passes, include overlaps, partial roll and hardroll
- 3. Roll 1st Pass (outside of test region) at high torque to ensure leakt195tness.
. 4. Plug this end of tube.
- 5. Test FLB condition first, then roll top 1 or 2 passes only, at appropriate higher torque, to include X, to achieve N. Op. T/TS radial pressure, then leak test N. Op. condition.
24 0245M:49/071368
FIGURE 3-2 i ROLL EXPANSION LEAKAGE TEST SAMPLE FOR ;
MODEL D STEAM GENERATOR LOCA CONDITIONS a.,c,s
- 1 i
NOTES '
- 1. Secondary Side source, deionized water, subcooled
- 2. Step roll 4 passes, include overlaps, partial roll and hardroll
- 3. Plug this end of tube. -
i
- 25 )
1 0245M:49/071388 ;
1 TABLE 3-1 ROLL EXPANSION LEAKAGE TEST RESULTS ROLL *SIFFERENTIAL SAMPLE EXPANSION . PLANT PRESSURE. TEMP. 7 NO. LENGTH IN. CONDITION PSI LEAK RATE.
PLANNED AQTUA( PLANNED ACitlAL DPM J J 4
i i
024'A:49/071388-26 i
-Q TABLE 3-1 (Continued) f ROLL EXPAWSION LEAKAGE TEST RESUi.TS 1
l 1
ROLL *0IFFERENTIAL SAMPLE EXPANSION . PLANT PRESSURE. TEMP. *F NO. LENGTH. IN. CONDITION PSI LEAK RATE.
PLANNED ACTUAL ftANNED ACTUAL DPM gg, (- gr l
I PO
%4 1
i
~
.j 024SM:49/071388-27
' b s y
_______ _ _ _ _ _ _ . - _ _ -w c
4.0 CRITERION AND TEST RESULTS FOR LEAKAGE FROM DEGRADED ROLL EXPANSIONS
~
, 4.1 LEAKAGE CRITERION
- The leakage acceptance criterion was developed for evaluation of the leak rate test results. Individual acceptance values were developed for test differential pressure conditions simulating normal operation, feedline break and LOCA.
4.1.1 NORMAL OPERATION The total primary-to-secondary leakage allowed in any one steam generator is specified in the plant Technical Specification as 0.35 gpm for a throughwall crack in the free span of a tube. It was determined based on commercial and operational considerations that the initial leak acceptance criterion for N. Op. would be based on having the total leakage from L* tubes to be [
]a,c,e of the 0.35 gpm limit ( ).a,c,e Using only a portion of the Technical Specification allowable leaktge to estas.ish the leakage criterion for L* tube ends accommodates leaktge from other locations. The
. number of tube ends to which this limit is to be applied was set at
[ ].a,c,e This number was established based on previous oddy current test results and trends with provisions for additional L* tubes to be added in the ,
future. Using [ Ja,c,e tube ends per steam generator establishes an average leak acceptance criterion of (
]a,c,e drops per minute (dpm) per tube end. (This is based on the equivalence of 75,000 drops per gallon.) l 1
The use of ( ]a,c,e of the limit in the Technical Specification, i.e., an allocation factor of ( ]a,c,e to determine the leakage rate criterion is an arbitrary decision for operational flexibility and does not require that the current Technical Specification limit be altered. The initial !
leak rate acceptance criterion may be reevaluated in the light of actual test
. results and the extent of degradation in the steam generator. Use of a larger fraction of the Technical Specification leakage limit and/or use of a smaller
- leak test acceptance criterion closer to actual results can be used to support application of the L* criteria to a number of tubes larger than the number on which the initial leak acceptance was based.
28 0245M:49/071988
4.1.2 FAULTED CONDITION: FEEDLIhE BREAK l Postulated feedline break conditions provide the maximum primary-to-secondary differential pressure across the tube. The steamline break (SLB) provides the -
most stringent radiological conditions for postulated accidents involving a
^
I loss of pressure or fluid in the secondary system. To establish the leak rate acceptance criteria for faulted conditions, the assumed SLB leakage rate is used with the feedline break pressure differential. This is the most stringent case because the associated primary-to-secondary side differential pressure causes the greatest amount of TS bending upward. This, in turn, causes the greatest amount of TS hole dilation for the hole interior locations, which causes the greatest, i.e., bounding, reduction in sp. The smallest s r leads to the most likely conditions for the largest primary-to-secondary leakage. The analysis of the steamline break described in the FSAR uses the assumption that the primary-to-secondary leakage in the steam generator in the faulted loop is 1.0 gpm. Using the same allocation factor as for normal operation the total leakage from ( ]c,c,e L* tubes at test conditions simulating SLB conditions must be less than [
]a,c,e dpm per tube end.
As with the N. Op. test leak acceptance criteria, use of a larger fraction of
- the allowable leakage and/or an acceptance criteria closer to actual results will support a larger number of tubes to which L* is applied. A site specific determination of acceptable leakage during a SLB event using the criteria of 10CFR100 would be expected to result in a significantly larger value for the aggregate leak rate.
The primary-to-secondary pressure differential for this condition will be
[ ]a,c,e psi &. It should be noted that dynamic leads on the tube joint, viz, secondary side fluid drag on the tube during the postulated accident, need not be added to this value. This is because the safety valve relieving event and the maximum fluid drag event are not concurrent. Therefore, the
[ ]a,c,e psia is the most stringent condition for test and an additional pressure difiirential, corresponding to the drag load need not be considered. -
This sequence of events also eliminates the need for augmenting the axial load 29 0245M:49/071988
i during the test to determine the axial loadbearing strength of DRE's. The axial load during the leak test is prototypical because it is the end cap load caused by the prototypiul differential pressure.
. 4.1.3 FAULTED CONDITION: LOCA Typically the small leakage from steam generator tubes into the primary system !
is not a significant consideration in the analysis of a postulated LOCA. To determine a leak acceptance criterion for test conditions simulating LOCA conditions, the values used in the determination of the acceptance criteria for N. Op. of total leakage per steam generator [ ]a,c,e and number of ,
tubes [ ]a,c.e to which the L* criteria can apply are'used. This results f in a conservative value of [ ]a,c.e per tube end for the LOCA leak acceptance criterion, the same as the criterion for N. Op. conditions.
The test conditions used to simulate the LOCA conditions of ( ]a,c.e psi secondary- to-primary differential pressure and a temperature equal to normal operating temperature are conservative for analyzed LOCA conditions.
, 4.2 ROLL TORQUE REDUCTION TEST RESULTS This test related srto rolling torque (T). It was determined that the s p was essentially directly proportional to T over the range of T's of from ( l
]a,c,e in.-lb., based on averaged OD hoop strains from two strain gages located at the middle of one of the four roll passes on each of six T/TS !
specimens. This relationship was determined to be [ Ja,c,e psi /in.-lb. I Simulation of the TS-bending effect for the subsequent leak tests in the laboratory was accomplished by reducing the T corresponding to the s r reduction in the plant.
The plant sp reductions as specified in Ref. 1 for N. Op. and FLB, were [
]a,c,e psi, respectively. Accordingly, the T for N. Op. was reduced l
. from the midrange value of (
Ja,c,oin.-lb. Therefore, the torque used to prepare the N. Op. leak test i - samples was ( Ja,c,e in.-lb. Similarly, the FLB leak test samples were ;
i l
l l
30 f 0245M:49/071888
prepared by ro' Ting at [ ]a,c e in.-lb. torque in the X, i.e., leak path, region. The same samples were used for the FLB and N. Op. conditions. The FLB condition was tested first, followed by rolling at the higher N. Op.. torque and then testing at the N. Op. condition. Between tests, the samples were -
maintained at a temperature above 212*F to prevent corrosion of the carbon steel collars and, therefore, possible extraneous RTL effects.
4.3.2 LEAXAGE TEST RESULTS 4.3.1.1 NORMAL OPERATION The N. Op. leakage test results shown in Table 3-1 were plotted in Figure 4-1.
The extreme values, i.e., the largest and the smallest values (short horizontal bar), and the average value (circled point) at each X are shown in the figure.
All of the average values were small, in comparison with the proposed criterton,i.e.,[ ]a,c,e DRE's (tube ends). Based on this criterion and the small average measured leak rates, [ ]a,c,e of the tube ends in the S/G could be dispositioned by an L* as small as ,
[ ]a,c e in., as shown in Figure 4-4. However, it's advisable not to consider any distance less than 0.50 in. without further study. The
[ ]a,c,o in, case was performed only to establish the leakage trend. The larger L* cases, [ ]a,c,e in., were also performed to establish the leakage trend in these instances, at very low leakage rates, and therefore, -
for a large number of tube ends, viz, [ ]a,c.e of the tube ends in the S/G. The case of the L* distances being greater than the F* distance generally would have no application for a plant such as V. C. Summer which is already using F*, The F* for V. C. Summer is 1.6 In., including ECT uncertaisty. The F* value is 1.06 In. without ECT uncertainty and m set by the FLB condition.
4.3.2.2 FEEDLINE BREAX The FLB leakage test results shown in Table 3-1 were plotted similarly to the N. Op. results and are shown in Figure 4-2. Althca:gh the FLB values are an
~
order of magnitude greater than the N. Op. values at respective X's, the FLB j leakages are small relative to the initial FLB leak acceptance criteria of ,
l 31 0245M:49/071988
[ ]a,c,e dpm per tube end. The fact that the FLB samples were made with smaller torques than the N. Op. samples correlated with the increased flow for the FLB samples. The slightly higher AP for the FLB test had a negligible effect on the flow because, based on Henry's Law for critical flow through
, small passages, the FLB flow would only have been expected to be larger than t'1e N. Op. by approximately the ratio of [
]a,c,e percent. Based on the FLB criteria and the small average measured leak rates, [ ]a,c e tube ends could be dispositioned at X's of
[ ]a,c,e and 0.50 in., respectively, and [ ]a,c,e the tube ends in the S/G at an X's of [ ]a,c,e in. Although the leakage is very small at X - [ Ja c,e in., and the number of tube ends which could be e
depositioned thusly is large, it is advisable not to consider any X less than 0.50 in., without further study.
4.3.3.3 LOCA The LOCA test results shown in Table 3-1 were plotted similarly to the N. Op.
and FLB results and are shown in Figure 4-3. As expected, the LOCA values were
[ ]a,c,e smaller than the N. Op. values
. at all X's. [ ]a,c e of the tube ends in the S/G could be dispositioned at all X's for the LOCA case, with the proposed N. Op. leakage criteria. The same
- reasoning with respect to minimum X value selection as discussed for N. Op. and FLB also applies to LOCA. Refer to Figure 4-6.
r 32 0245M:49/071988
NORMAL OPERATION LEAKAGE TEST:
LEAK RATE PER DEGRADED ROLL-EXPANSION VS.
X o
2 a.,c,e h
Figure 4-1
- 33 0245M:#9/071388
.. - _ - - . - . - . - . , . - . . - . . . . . _ -, --- - . - . . - , = , . . . . . . . . , - - . - - - . , . . . . - , , _ - . - - .
FEEDLINE BREAK-LEAKAGE TEST:
LEAK RATE PER DEGRADED ROLL EXPANSION VS.
X a.,c,e b
~
r 4
1 l
I Figure 4-2 34 0245M:49/071388
LOCA LEAKAGE TEST:
i LEAK RATE PER DEGRADED ROLL EXPANSION e VS. ,
X i
a,c,e I
I l
1 i
I l
1 I
1 4
Figure 4-3 I
- 35 0245M
- 49/071388
__ _ _ _ , _ - .__ . , . - . _ . - . _ . . = - . . _ _
- . . ,. = . _ . . ..
NUMBER OF TVBE ENDS PER STEAM GENERATOR DISPOSITIONED BY L*-NORMAL OPERATION i
P i
a.,c,e d
l
. ~
Figure 4-4 l
(
36 0245M:49/071388
NUMBER OF TUBE ENDS PER STEAM GENERATOR DISPOSITIONED BY L*-FEEDLINE BREAK a.,c,e l
i Figure 4-5 37 0245N:49/07138S
NUMBER OF TVBE ENDS PER STEAM GENERATOR DISPOSITIONED BY L*-LOCA a.,c,e 4
I t
l Figure 4-6 i
38 0245M:49/071388 a--+ w s, ,
8 5.0 TEST FOR THE DETERMINATION OF AXIAL LOADBEARING CAPABILITY OF DEGRADED ROLL EXPANSIONS D
5.1 INTRODUCTION
Tests were conducted to determine the axial loadbearing capability of the -
Model D steam generator degraded roll expanded T/TS joints. The simulated degradation consisted of a single band of axial or near axial slots. The results of the single band degradation (SBD) tests were also applicable to the most stringent, i.e., uppermost band of maltiple band degradation (MBD) in the 4 plant. The MBD configuration did not require testing. A series of tests was performed to determine the ultimate pull force-(i.e., tensile strength) for non-degraded and artificially degraded, prototypical, [ .
Ja,c,e in. (nominal) wall 1600MA tubes rolled into carbon steel AISI 1018 collars. The limiting test condition involved application of the pullout j load directly to the top of the SBD. The roll expansion was performed utilizing a prototypical Model D steam generator T/TS roller with regulated rolling torque to provide proper levels of tube thinning and joint preloads.
Slots were electric oischarge machined (EDM) into the wall of tube specimens to simulate throughwall indications. All aspects of the test were designed to bound the plant respective conditions. Refer to Table 5-1. For example, tubes were t1achined to have [ ]a,c,e slots inclined at angles of [- ,
]a,c e degrees from the axial centerline of the tubes, whereas, the plant indications involve a maximum of approximately ( ]a,c,e ECI's in the i
uppermost SBD and [ ]a,c e at lower elevations, all at approximately
[~]a,c,e degrees. The length of the slots was [ ]a,c e in, which was considered to bound 1he plant length by a factor of approximately 2, based on.
eddy current test assessments. The test axial distance "X" between the tops of the slots and the BRT was [ ]a,c e in.; the plant "X" for the intended tubes was approximately 0.5 in. The smaller (test) X was judged to bound the plant X; lower strength generally results from a smaller X. The collars were l geometrically sized to simulate the equivalmt structural reaction of a unit j cell of the steam generator tubesheet. Figures 5-1 and 5-2 illustrate the basic shape and geometries of the collared and decollared specimens, l
. respectively. The never-collared tube specimens are shown in Figure 5-3.
4 l 39 0245M:49/071388
All tests were performed at ambient atmospheric conditions employing a tension-compression machine shown in Figure 5-4. Because the plant tubes of
( Ja,c,e ksi ultimate tensile strength (UTS) at 650*F, will be j stressed at elevated temperature, the room temperature test results obtained with tubes of [ ]a,c,e ksi UTS were analytically adjusted (decreased) to '
I a count for the small effects of temperature and UTS, for the Pull Strength hdel discussed later in this report.
5.2 OBJECTIVE The objectives of this test were: (1) Determine the ultimate pull force necessary to structurally fail non-degraded and artificially degraded [
]a,c e inch wall Inconel 600 tubes rolled into collars that simulated a unit cell of the steam generator tubesheet. (2) Measure the ultimate pull force necessary to structurally fail degraded non-rolled Inconel tubes free of collars. (3) Evaluate the type of failure mechanism that induced structural failure in the tubes.
5.3 TEST EQUIPMENT 5.3.1 Tubes, prototypical 1600MA, [ ]a,c.e in, wall (nominal). Refer to Table 5-1 for specimen features, and to Table 5-2 ,
for specific specimen configurations tested.
5.3.2 Tubesheet simulants (collars), cold rolled carbon steel, AISI 1018, 4.5 In. long. Refer to Figure 5-1.
5.3.3 Airetool Roll Expander, No. [ ]a,c,e rolls per original (shop) fabrication specification.
5.3.4 Roll Expander Motor, of 30 to 95 in.-lbs. torque capacity, 1250 RPM no-load speed.
~
5.3.5 Zetac Tension-Compression Testing Machine. Refer to Figure 5-4.
5.3.6 Other laboratory strength test equipment such as furnaces, measuring instruments, etc.
40 0245M:49/071388
5.4 TEST MAJOR STEPS 5.4.1 COLLARED TUBES The collared tubes were prepared by rolling to approximately two percent thinning, then removed from the collar for EDM of the slots, if slots were required; the tubes were then replaced in the respective collars, hardrolled to the prototypical midrange torque, [ ]a,c,e in.-lbs. and welded. The specimens were then "pulled" in a configuration similar to that shown in Figure 5-4.
5.4.2 DECOLLARED TUBES The decollared samples were prepared by rolling the tubes to a small amount of wall thinning, a non-interference fit, then removed from the collar for EDM of the slots; the tubes were then replaced in the respective collars, and (hard) rolled to the specified midrange torque, [ ]a,c,e in.-lbs.
After hardrolling, the specimens were decollared, i.e., the collars were carefully removed by machining. This step prevented the collar from providing the known beneficial effects, primarily friction, during the strength test.
. The slot geometry and axial location are shown in Figure 5-2. Table 5-2 identifies the specimen configuration in terms of the number of slots, angle, and slot top location with reference to the BRT.
5.4.3 NEVER-COLLARED TUBES These tubes were prepared for pulling by EDM of the slots. The reason for including this configuration was to show that it involves higher, i.e.,
non-limiting ultimate strengths. The specimens were pulled in a configuration similar to that shown in Figure 5-4.
O 41 i 024 91:49/071388
l 5.4 TEST FACILITY 1
The tests ware performed at the Westinghouse Research and Deveopment Laboratory. A Zetac tension - compression test machine was used to apply a continuously incre.ising axial pull force to the tube specimens. Figure 5-4 is a sketch of the test machine. The tube top ends were pernitted to rotate as dictated by the slot angle, p. A certain amount of rotation about the tube vertical axis could occur in the plant in keeping with the conservative assumption that the SRE within L* provides no fixity to the TS, owing to the tube straight leg acting as a torsional spring and being anchored at the U-bend. It was judged that this providad icwer ultimate values than if axial rotation were prevented.
5.5 DATA ACQUISITION SYSTEM Axial pull forces were recorded on Cartesian coordinates as a function of tube top end displacement by employing an X-Y analog plotter.
5.5 PROCEDURE <
The specimens were vertically mounted in the Zetac testing machine as shown in Figure 5-4. A steadily increasing axial pull was applied to the specimens ,
until ultimate load failed the tube. A curve was simultaneously drawn by an X-Y analog plotter of the pull force versus tube top deflection. Curves were '
produced for non-degraded and degraded tubes; non-degraded tubes were tested to establish reference data in order to assess the strength of the degraded tubes. .
The puil specimens were tested per Table 5-2.
i i
' l l 42 0245M:49/071388 l
s a.,c,e -
d
.l
)
1 I
l
- Figure 5-1 !
Geometry of Slots in Collared Tube Specimen -
Ultimate Strength Test 43 0245M:49/071388
G,C,C
- l 1
i f
1 i
[
Figure 5-2 '
Geometry of Slots in Decollared Tube Specimen -
Ultimate Strength Test j
1 44 024sM:49/0nssa
1 a.,c,e-p k
l l
1 I i .
l l
l Figure 5-3 Geometry c,f Slots in Unexpanded, Never-Collared Tube Specimen - Ultimate Strength Test I
' I 45 l 024y:4Wo713ss j
. . -. =. .- - . . . - . _ . . - _ _ _ - - _ _ _ -
Q.,C,6 i
I a -I l
4 f
v i
9 i
! Figure 5-4 1 '
Sketch of Zetac Tension and Compression
) Testing Machine i
i 46 0245N:49/071388 1
, . , - . - . . - - - , . - , --v,.,. ..,r..-,- .. . - - - - , - - . _ .. ,
TABLE 5-1 SELECTION OF BOUNDING FEAfURES FOR AXIAL LOADBEARING TEST OF DEGRADED ROLL EXPANSIONS Asoect Plant Igit Boundina Feature a.,c,e l
l l
1 l
l 1
4 l 47 0245M:49/071356
TABLE 5-2 AXIAL LOADBEARING STRENGTH TEST OF DEGRADED ROLL EXPANSIONS Array Top Distance Slot Specimen
- Specimen Below Bott No of Angle, Ultimate No. Confia, of Trans.. In. Slots Daarees (qA W
%cf__
v i
1
. I 48 0245M:49/071388
l 1
l l
6.0 CRITERION AND TEST RESULTS FOR THE AXIAL LO'ADBEARING CAPABILITY OF l
DEGRADED ROLL EXPANSIONS j 6.1 MAXIMUM AXIAL LOAD CRITERION 6.1.1 SINGLE BAND DEGRADATION The maximum axial tensile load which a DRE must bear will be the most stringent of the N. Op. and FLB axial loads. The LOCA axial load is axially compressive and cannot cause tube pullout.
6.1.2 MULTIPLE BAND DEGRADATION If degradation occurs in discrete circumferential bands i.e.,- as single band degradation (SBD), or arrays, in an RE above the F* elevation, thereby preventing use of F*, it may also occur in multiple discrete bands below the F*
elevation. This is termed multiple band degradation (MBD). As discussed earlier, it's desired to avoid quantitating each of the possible several degradation bands to apply the pullout load to the T/TS weld. Instead, it's desired to apply the load only to the uppermost sound roll portions (expansions) which are interspersed with the uppermost degradation bands. For
. the purpose of simplicity, each degradation band requires the same strength.
This strength is the same as that for the SBD. Therefore, for the MBD case, the number of individual bands which must be quantitiated is one less than the number of SRE's (N) needed to react the pullout load. If N is 3, N-1, is 2.
Conservatism may be added to this calculation in the form of a larger N than is !
needed.
6.1.2.1 Normal Operation The tube ultimate load required for normal operation is approximately [
].a c e This load was increased slightly for additional margin, to [ ]a,c.e 1bs., per Ref. 1. A safety factor of three is applied to this load. Therefore, the tube axial
. loadbearing requ'rement is [ ]a,c,e1b. This load is the highest for the i three conditions considered and is plotted in Figures 6-21 through 6-26.
49 0245M:49/071388
6.1.2.2 Feedline Break The tube ultimate load required for FLB is approximately [
Ja,c,e A safety factor of [ *
]a,c e for allowable stress for faulted conditions) is proposed for this load. Therefore, the tube axial load bearing requirement for FLB is [ ]a,c.e1b. As discussed in Ref. 1, this load need not be further augmented for dynamic loading effects during FLB because of the sequence of events during an FLB.
6.1.2.3 Loss of Coolant Accident An axial loadbearing requirement for a LOCA event acts to move a tube downward, the opposite of pullout. Therefore, this condition does not apply to pullout.
(However, the DRE must not collapse under LOCA conditions. This condition was tested with prototypical, i.e., compressive, end cap loads, in the leak test section.)
6.2 STRENGTH TEST RESULTS Individual strength test results are presented in Figures 6-1 through 6-20.
These figures are arranged in approximate order of increasing severity and the ,
results were previously summarized in Table 5-2. [
Ja,c.e Structural failure of the tube was attributed to the axial and shear forces acting through the tube wall as a ,
function of the vectorial relationship between the axial pull force and the '
I angle of the respective slot to the tube axial centerline. !
I i
50 I
0245M:49/071388
. . . . __ .. -_,~. _
The strength test results discussed above dealt with pullout loads. It should be recalled that LOCA i.e., compressive loads on DRE's were discussed in Ref.1
. and that no deleterious structural effects are exnected.
. The basic objective of the L* program was to deonstrate the acceptability, for leakage and strength, of degradation arrays of known or bounding dimensions and configurations. The leakage issue was addressed elsewher6 in this report. The strength design curve will be developed here and is based on the strength test results.
6.3 EVALUATIONS AND CONCLUSIONS 9
i 6.3.1 PULL STRENGTH MODEL J The pull strength of Model D S/G tubes containing circumferential arrays of axial or near axial ECI's has been empirically established in Section 5.0 and
]
the results were discussed in Section 6.2. The deformation behavior of tubes .
- in these tests is modelled as follows. As the tube is loaded, the initial r deformation is elastic followed by plastic yielding. Depending on the nmober ;
l . of cracks and crack angle, yielding may first occur in the virgin unrolled tube ;
- section or in the crack array. Large numbers of cracks and high crack angles (4) favor yielding of the crack array. Yielding of the crack array
- proceeds by plastic bending of the ligaments between cracks and subsequent rotation of the cracks toward the longitudinal axis of the tube. As rotation ;
- occurs, the applied axial force must increase substantially since the effective .
j moment are for ligament bending is continually decreasing. Strain hardening of
! the ligaments adds to the geometric hardening produced by crack rotation. As i l the load is increased, general yielding of the. Virgin unrolled tube section ;
! develops, foilowed by strain hardening and then yielding of the rolled bu6 i uncracked section of the tube. At some point, depending on the crack morphology, fracture will terminate the process of plastic deformation.
. The equation describing the uhl load required to yield a crack array, as sketched in Figure 6-27, is d.,Cg I
! 51 1
! 0245M:49/0M388 l
_ _ _ _ _ - -_- , . . - . _ - _ . _ _ -- _ . . _ - . , _ _ - . ....1
The axial plastic displacement resulting from yielding and then rotation of the crack array is given by
[ ja,c.e ,
If the crack array plastic displacement, as a function of axial load, is added the baseline load-displacement record of an otherwise identical tube without a slanted crack array then the computed load-displacement records closely approximate actual measurements as illustrated in Figures 6-28 and 6-29.
The prediction of ultimate pull strength load as opposed to yield loads is an elastic-plastic fracture mechanics issue rather than a plastic collapse issue because maximum loads are caused by crack tearing (beyond some minimum crack angle). A J integral approach has been applied to the issue of estimating the onset of crack tearing with good success. The equation describing the plastic load-displacement behavior of the crack array was integrated to relate plasi.ic work, crack length and displacement. The compliance definition of J was then used to compute J. Since the critical value of J for the onset of crack tearing in 1600 is very high, the neglect of elastic contributions to J is appropriate. ,
Pull test results of tubes with slanted EDM notches at high crad nr.gics are consistent with a critical value of J of ( Ja,c.ein.lbs./in.2 Mill annealed 1600 plate exhibits a JICvalueofabout[ ]a,c.e in, lbs./in.2 The empirical critical value of J is reasonable for thin walled tubing which is cold worked several percent by hard rolling. The EDM notches had a root radius of less than [ ]a,c.e inches. Most notch root radii were less than this value since they were rolled after machining. Since the notch root radii were many times less than the critical crack opening displacement (COD) associated with a critical J of [ ]a,c.e in, lbs./in.2, the EDM slots provide a reasonable representation of the behavior of natural cracks.
52 0245M:49/071388
The measured pull strength of a tube rolled into a collar in the absence of any EDM slots as plotted in Figure 6-1 is on the order of ( Ja,c.e pounds.
This load is the load required to yield a tube which has been cold worked several percent by rolling into a collar. As yielding of the rolled tube
. develops, the tube shrinks away from the collar and frictional resistance between the tube and collar continually diminishes until finally the weld at the bottom of the collar is subjected to the full applied axial load. Fracture of these welds limited the pull strength to [ ]a,c,e pounds as opposed to an ultimate load in the vicinity of ( ]a,c.e pounds for a virgin length of tubing. A reasonably conservative bound to the pull strength in a collar or tubesheet is therefore the axial yield load of a tube with an expected amount of cold work due to ralling. For the heat of material tested and an equivalent strain of ( ]a,c.e pounds. Pull strengths less than this value will occur when fracture at the crack array intercedes.
The J integral fracture analysis indicater that there is a worst case number of evenly spaced cracks for each slanted crack angle. Loads at a J critical of l
[ ]a,c.e in, lbs/in.2 are plotted in Figure 6-26 as "Pull Strength l
. Model" as a function of crack angle, The worst case number of cracks is assumed. It is evident that the strength curas based on rolled tube yielding l at low crack angles and crack tearing at higher crack angles provides an )
excellent low bound to the actual test data and a very good pull strength ;
model. This model was repeated as the Model D S/G Degraded Tube Pull Strength Design Curve for V. C. Summer, plotted in Figure 6-30. l Indication lengths (l's) of ( Ja,c,e in, were tested because the ECT ,
data showed l's of approximately [ ]a,c,e in, and a bounding curve was needed. Observation of axial degradation in skiprolls and RT's discussed in Ref. 1, shows that the degradation tends to arrest when it reaches the edges of l the skiproll discontinuity or RT. It is expected that axial degradation below i 1
the BRT, e.g., related to roll overlaps and repaired skiprolls, will also j l
. arrest when it reaches the edges of the RE womaly. Because tubes with shorter l's have higher ultimate pull strengths, the design curve is very l
l 53 -
i I
0249t:49/071388
conservative. The design curve was determined for the V. C. Summer tubes which have UTS's of ( ]a,c.e ksi, (LTL) at 650*F. Thi's parameter was analytically adjusted. The UTS for the tubing used in the test was [ ]a,c.e ksi. The Design Curve also includes a small reduction for the lower ultimate strength at [ ]a,c,e.F vs. the ultimate strength at room temperature ,
which was used for test and for the fracture mechanics analysis, the results of which are shown in Figures 6-28 and 6-29.
The Design Curve was developed to be used for comparison with tube axial loads, multiplied by suitable factors of safety, caused by the most stringent normal and faulted operation conditions. The highest load thus calculated,
[ ]a,c,e 1b., resulted from normal operation conditions, from Ref. I and is shown plotted as SF = [ Ja,c.e, in both Figures 6-21 through 6-26 and in Figure 6-30. This load is significantly below the Design Curve, for all p's up to ( Ja,c,e degrees, as shown in Figures 6-26 and 6-30, 6.4 RESULTS Reference to Figure 6-30 shows that the most stringent axial load, (
Ja,c,edegree ,
inclination angle, larger than the largest angle noted in the plant.
The curve should be applied to tubes with relatively well characterized ECI's, e.g., tubes for which substantiated rotating pancake coil (RPC) ECT data are available. The strength criterion per this curve, must be met for tubes with eddy current indications.
6.5 CONCLUSION
With sufficiently quantified information about degradation in the anchor regions of the V. C. Summer S/G's roll expansions, use of the Design Curve is l
expected to provide safe and reliable operation. (The leakage criterion for j this case is discussed elsewhere and involves an upper limit on the number of ,
tubes per steam generator which may be dispositioned thusly.)
l I
54 U245M:49/0?!S88 i
a., C,e 1
l t
i N
a .
i B
1
- l 4 L 4
i 1
4
! Figure 6-1 1
Pull Force as a Function of Pull Displacement j for a Collared, Nondegraded Tube (Specimen No. 11) i 1
55 0246M:49/071388
4 Q., C,6 i i
o i
J Figure 6-2 i Pull Force as a Function of Pull Displacement for a Collared, Nondegraded Tube (Specimen No. 16)
- Replication of Specimen No. 11 56 0246M:49/071388
Q , c,e i
t
(
]-
Figure 6-3 Pull Force as a Function of Pull Displacement for a Collared Tube (Specimen No. 3), with (15)-30 Degree Slots, a
Tops of Slots 0.25 Inch Below the Bottom of Roll Transition i
I
- 57 0246M:49/071388
Q.,C,C
! +
l
'I t
! t i
d 4
~ -
Figure 6-4
" 9 j Pull Force as a Function of Pull Displacement i for a Collared Tube (Specimen No. 8), with (15)-30 Degree Slots, ,
! Tops of Slots 0.25 Inch Below the Bottom of Roll Transition -
Replication of Specimen No. 3 i
58 0246M:49/071388
. _ .-. . _ . - - .- - _-. -_ ..-- . _~ . __ _,-._ - - . . .
43.,C ,C .
1 i
m e
Figure 6-5 Pull Force as a Function of Pull Displacement for a Deco 11ared Tube (Specimen No.12), with (15)-0 Degree Slots.
Tops of $1ots 0.25 Inch Below the Bottom of Roll Transition 074m/0713M:49 _ __ _ _ _ _ . _ _ . , , , _ _ , , . , _ _ _ . _ , . , , , . , _ .. _ , _ . _ _ _ . _ _ . _ .
.. - - - . _ _~ _ _ _ - = - -_ . . - _ - . - - _ . .- _ - - , . . = - .._ . . _ - - . _ . ..
Q.,C,e 9
m o
t j
j 1
)
F6gure 6-6 -
Pull Force as a Function of Pull Displacement for a Deco 11ared Tube (Specimen lio. II). with (15)-0 Degree Slats, Tops of Slots 0.25 Inch Below the Bottsa of Roll Transition - Replication of Specimen 110. 12 0746M/073388:49 * *
- i a
a.,c,e 4
k 7
1 4
l 1
l l
l j - -
Figure 6 7
- . Pull Force as a Function of Pull Displacement ,
I j for a Decollared Tube (Specimen No. 13), with (15)-15 Degree Slots, l
Tot t of Slots 0,23 Inch Below the Bottom of Roll Transition !
i 61 l I
2 0246M:49/Mi$l8 I
000 1 1 ,
i i
f e
^
Figure 6-8 Pull Force as a Function of Pull Displacement -
for a Decollared Tube (Specimen No. 18), with (15)-15 Degree Slots,
~
Tops of Slots 0.25 Inch Below the Bottom of Roll Transition
- Replication of Specimen No. 13 62 0246M:49/071388
a.,c,e e
. i l
Figure 6-9 l Pull Force as a Function of Pull Displacement for a Decollared Tube (Specimen No.1), with (15)-30 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition 63 i 0246M:49/071388 i
- .. . . _ _ - _ ~. -
b ,
I i
{
~
Figure 6-10 .
Pull Force as a Function of Pull Displacement-for a Decollared Tube (Specimen No. 7), with (15)-30 Degree Slots, .j Tops of Slots 0.25 Inch Below the Bottom of Roll Transition
- Replication of Specimen No.1 I
64 0246M:49/071388
l I
a,c,e l
l 4 ' i e
l 1
1 I
l
-1 l
, l 1
Figure 6-11 1 Pull Force as a Function of Pull Displacement for a Decollared Tube (Specimen No. 2), with (15)-45 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition I
l 65 0246M:49/071388
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8,_ -
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s (
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. 1 f 9 r 6
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. ig c n n e f F u m o F i c m a ep to
. s S t a ( o B
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d w l
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o h c c _
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9 4:
8 8
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Q.,C,6 I
i i
1 Figure 6-14 Pull Force as a Function of Pull Displacement' for a Decollared Tube (Specimen No. 15), with (30)-15 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition i I
( 68 ;
- 0246M
- 49/071388 l l- - - - -- --
l
- 0.,c, e _
o l
l 1
I l
l
. Figure 6-15 Pull Force as a Function of Pull Displacement for a Decollared Tube (Specimen No. 20), with (30)-15 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition
- Renlication of Specimen No.15 l 69 0246M:49/071388
l l a.,c,o 1
l l
l 1
)
9
{
I I
f I
l
~
Figure 6-16 Puli Force as a Function of Pull Displacement 1 l
I for a Decollared Tube (Specimen No. 4), with (30)-30 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition
ajc,a _
1 l
P
~
Figure 6-17 Pull Force as a Function of Pull Displacement for a Decollared Tube (Specimen No. 9), with (30)-30 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition
- Replication of Specimen No. 4 71
. _ _ _ 02 W s m issa _ _
0.,C,e P
Figure 6-18 Pull Force as a Function of Pull Displacement ,
for a Decollared Tube (Specimen No. 5), with (30)-45 Degree Slots, Tops of Slots 0.25 Inch Below the Bottom of Roll Transition i
72 0246M:49/071388
l l 1
- a,c,o
- l. _
o l 1 1 i l
> l i
i l
i l
I Figure 6-19
- Pull Force as a Function of Pull Displacement for a Never-Collared / Expanded Tube (Specimen No. 6),
with (30)-30 Degree Slots 73 0246M:49/071388
.. . - - . . ~.. . _ . . .= . - - _ _ _ _ _ _ _ _ _ _ _
a.,c,0 4
Figure 6-20 Pull Force as a Function of Pull Displacement ,
for a Never-Collared / Expanded Tube (Specimen No. 10),
With (30)-30 Degree Slots, - Replication of Specimen No. 6 0246M:49/on388
. ~ . . . _ . _ . . . _ , - _ _ -_ -.
5 J b I
G,C,e l :
1 I.
i
) !
i
?
1-i i
N m
i 1
i 4
1-1 1
i
- t i Figure 6-21 ,
t .
I' Ultimate Pull Force as a Function of Slot Angle for Expanded, Collared and Decollared Tubes, 15 Slots, Tops of Slots 0.25 Inch Below the Bottom of. Roll Transttton
(
l l
0246M/071388:49
' 43.,C,C I
a -
1 N
G i
I Figure 6-22 Ultimate Pull Force as a Function of Slot Angle for Expanded Decollared Tubes '30 Slots Tops of Slots 0.25 Inch Below the Bottom of Roll Transition Never-Collared, Non-expanded Specimens Also Shown 024fJe/070888:49 -
- l II 6
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l l
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m n d o n o e i r t e r
t t a oB t 6 c l S 2 n l
- u o e l 6 F c h l e t u e a D P r w u s d o d l ig a na e a n F e B s
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M
. 6 4
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_ 0 mO
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.-- .. . .. . . _ _ .___ =_ _ _.
gc,e - J t
l l
f 1
! Figure 6-27 Model for Plastic Collapse
)1 81 0246M:49/enssa ,
-.__. = - . - . . . . . . - . - . .- . -
i -
'a.,$e l
! l l
l ,
. l l
I I
l !
4 Figure 6-28 Comparis9n of 1.oad-Displacement Records, ,
Computed Vs. Measured for 30 Slots at (=45' 82 0246M:49/071388
a.,c,e I
l Figure 6-29 Comparison of Load-Displacement Records, Computed Vs Measured for 30 Slots at (=30'
, 83 -
0246M:49/071388
l I G,Cp t -
i
) s i
M i Figure 6-30 ,
Model D Steam Generator Degraded Tube Pull Strength Design Curve for L*
074fM:49/07r3sa . * . 9 '
7.0 PULLOUT LOAD REACTION AND RPC INSPECTION LENGTHS
. 7.1 GENERAL The methods used to inspect, evaluate, and define the tube degradation for the application of the L* alternate plugging criteria may be quite time consuming and require considerable personnel resources. Inspection of the degradation must be done by a RPC eddy current probe or other advanced inspection method.
Data collection using such methods is considerably slower than standard methods and the evaluation is less automated. If sufficent length of sound expanded tube is extant in the upper portion of the tubesheet, the lower portions of the tube contribute relatively little to the structural integrity and resistance to leakage of the tube. Criteria have been developed which limits evaluation to the more important upper portion of the tube expansion.
The analysis to support limited evaluation of the degradation in a tube is based on a combination of the methods used to qualify the F* and basic L*
criteria. The analysis demonstrates that sound portions of expanded tube between degraded sections can provide for structural intergrity of the tube.
In the same manner that the friction force between the tube and the tubesheet in an F* tube will resist pullout, the sound portions of a tube with degradation meeting the L* criteria will provide a frictional force to resist pullout. The calculation of the length with degradation within the portion of the expanded tube required to resist pullout includes the possible reduction in friction force next to the degraded portions due to end effects. The length of tube required to resist pullout plus the length of the degraded tube portions establishes the length of tube which must be inspected and evaluated by rotating pancake coil or other advanced inspection methods. Therefore, rather than survey the entire length of RE, to the T/TS weld, only that portion near the TTS which will provide adequate anchoring for all axial loads need be surveyed. This length, in addition to the SBD or MBD within it, is the minimum length, beginning at the BRT and extending downward, and is referred to as the pullout load reaction length (PLRL).
f 0246M:49/071388
7.2 SINGLE BAND DEGRADATION The PLRL for this case consists of four parts, the SRE within L* consisting of the central and end-effect portions, the recommended 0.5 In., plus the required .
SRE immediately below the SBD. Calculations for N. Op. and FLB are shown in
~
Table 7-1. The limiting case was FLB; a PLRL of [
]a,c,e in of RE to be RPC inspected.
7.3 MULTIPLE BAND DEGRADATION The PLRL for this case consists of six parts, the SRE within L*, i.e., the recommended 0.5 in. consisting of the central and end-effect portions, the uppermost two degradation bands and the two SRE's below these two respective bands. Calculations for the N. Op. and FLB are shown in Table 7-1. The PLRL's for N. Op. and FLB conditions were the same. A PLRL of [ ]a,c,e in. for each was determined. Addition of the two degradation band lengths and the ECT uncertainty of [ ]a,c,e in. to the PLRL resulted in a length of
[ ]a,c,e in of RE to be RPC inspected.
7.4 TECHNICAL SPECIFICATIONS REQUIREMENTS .
To facilitate the specification of Technical Specifications for the length of degraded tube required to be inspected by an advanced inspection technique, the information provided above can be combined into a single set of criteria. The length of sound tube calculated to be required in both cases above is less than the F* value of 1.6 in. Inspection of the top [ ]a,c.e in, rounded off to
[ ]a,c.e in., of the tube should determine if'there is at least
[ ]a,c,e in., plus ECT uncertainty, of sound tube with no more than l
[ ]a,c,e bands of axial degradation separating the sound portions and with an L* of 0.50 in. plus ECT uncertainty.
I 86 0246M:49/071988
l TABLE 7-1 CALCULATION OF PULLOUT LOAD REACTION LENGTH AND TVBE LENGTH TO BE RPC INSPECTED A. Single Band Degradation - Normal Operation
- 1. Calculation of Pullout Load Reaction Length PLRL Axial Reaction Reaction Portion Lenath. In. force. lb./ih force. 'Ib. a ,c,e l
d l
I I
i 87 l 0246M:43J071388
1 l
l TABLE 7-1 (Contd) l CALCULATION OF PULLOUT LOAD REACTION LENGTH AND TUBE LENGTH TO BE RPC INSPECTED i
I 1
B. Single Band Degradation - Feedline Break
- 1. Calculation of Pullout load Reaction Length PLRL Axial Reaction Reaction Portion Lenath. In. force. lb./in. force. lb.
CL ,C,6
""" mmm O
O l
88 024M:49/071368
l l
TABLE 7-1 (Contd)-
CALCULATION OF PULLOUT LOAD REACTION LENGTH AND L TUBE LENGTH TO BE RPC INSPECTED I C. Multiple Band Degradation - Normal Operation
- 1. Calculation of Pullout Load Reaction Length PLRL Axial Reaction Reaction Portion Lenath. In. force,lh,/in, force. lb. ct,c e l
l i
)
e M
i I
89 i 024sy:49/0713sa j
TABLE 7-1 (Contd)
CALCULATION OF PULLOUT LOAD REACTION LENGTH AND TUBE LENGTH TO BE_RPC INSPECTED
- 0. Multiple Band Degradation - Feedline Break
- 1. Calculation of Pullout Load Reaction Length , ,
PLRL Axial Reaction Reaction Portion Lenath. In. force. lb./in, force. 16. C.3 C,C
- 1 6
I 1
I
)
90 0246N:49/071388
8.0
SUMMARY
8.1 GENERAL
= ,
, The semi-empirical evaluation uncoupled the leakage and strength effects. The issues were determined separately.
8.2 LEAKAGE 8.2.1 NORMAL OPERATION a, c,e l
=
l
- 91 0246M:49/071358 i 1
a.,c,e 1
l 8.3 STRENGTH The axial strength of DRE's far exceeds the required axial strength for the most stringent plant condition. Based on the limiting assumption that the L*
region, of length 0.50 In. provides no axial fixity, the uppermost degradation band was tested for strength. It was understood that the plant uppermost a.ge With the axial fixity afforded by the SRE within the L* region, 0.50 In. long, only approximately [
),a,c e 8.4 RECOMMENDED L* CRITERIA With sufficiently quantitated information about degradation in the upper approximately [ ]a,c,e in, which may be rounded off to 3.00 in., of degraded roll expansions, use of the recommended L* value of 0.50 in., plus ECT uncertainty ([ ]a,c,e in, was used for an example), is expected to provide for coatinued safe and reliable operation of the V. C. Summer S/G's. ,
i l
92 0246M:49/071988
Inspection of the 3.0 in, distance should determine if there is at least 1.3 in., plus ECT uncertainty, of sound tube with no more than two bands of axial
, degradation separating the sound portions. The largest ( permissible for the ECI's in a degradation band is 30*. A conservative number of tube eads in a S/G which may be dispositioned thusly is 2648.
4 i
93 0246M:49/072588
9.0 REFERENCE
, 1. WCAP-11228, Rev.1, "Tubesheet Region Plugging Criterion for the Scuth Carolina Electric and Gas Company, V. C. Summer Units 1 and 2 Steam
' Generators", Westinghouse Electric Corp., October 1986. (Proprietary)
- 2. WCAP-11229, Rev.1, "Tubesheet Region Plugging Criterion for the South Carolina Electric and Gas Company, V. C. Sunner Units 1 and 2 Steam Generators", Westinghouse Electric Corp., October 1986. (Non-Proprietary) a a
94 0246M:49/071988