ML20082N988
| ML20082N988 | |
| Person / Time | |
|---|---|
| Site: | North Anna  |
| Issue date: | 08/31/1991 |
| From: | Wootten M WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19312B437 | List: |
| References | |
| SG-91-08-028, SG-91-8-28, WCAP-13035, NUDOCS 9109100143 | |
| Download: ML20082N988 (326) | |
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n - Westinghouse Class 3 WCAP 13035 SG-91-08-028 -. North Anna Unit 1 Steam Generator Operating Cycle Evaluation i i
.z August 1991
\b Approved by: \N M. J. Wootten, Manager Steam Generat r Technology & Engineering l
. -l i
s.
,4 C 1991 Westinghouse Electric Corporation
TABLE OFCCNTENTS SantiQD T1112. E002
1.0 INTRODUCTION
11 2.0
SUMMARY
ANDCONCLUSIONS 21 2.1 Overall Conclusions 2-1 2.2 Summary of Principal Results 23 2.3 Summary of Report Sections 2-8 3.0 PULLEDTUBE EXAMINATIONS 3-1 3.1 1991 Pulled Tube R11C14 3-1 3.2 Prior Tube Exams from North Anna Unit 1 SGs 3-4 3.3 Crack Morphology for Tube integrity Assessments 3-6 4.0 EDDY CURRENT DATA REVIEW 4-1 4.1 1991 Inspection Summary 41 4.2 WEXTEX Expansion Region 4-2 4.3 Axial Cracks at Tube Support Plates 4-5 4.4 Circumferential Cracks at Tube Support Plates 4-7 4.5 Detection Sensitivity for Circumferential Cracking 4-11 5.0 PROJECTED EOC CIRCUMFERENTIAL CRACK DISTRIBUTIONS 5-1 5.1 Detection Thresholds 5-2 5.9 Growth Rate Distributions 5-2 5.3 Eddy Current Uncertainties 5-5 5.4 Estimated Number of 1992 EOC Ind' cations 56 5.5 Methods Evaluation Against 1991 Crack Distribution Results 5-7 5.6 EOC Circumferential Crack Estimates at 90% and 95% 5-8 , Probability Levels 5.7 Projected EOC WEXTEX Crack Distributions 5-9
- 5.8 Projected EOC TSP Crack Distributions 5-10 5.9 Distribution of Indications at First TSP Elevation 5-10 5'3 Summary of Probabilities for Exceeding Acceptable Through 5 10 Wall Crack Angles I
4 TABLE OF CONTENTS (continued) Section Iltle Ea;te 6.0 TUBE-!NTEGRITY ASSESSMENT: TUBE BURST AND LBB 6-1 6.1 Burst Strength of Tubing With Circumferential Cracks 6-2 , 6.2 Burst Strength for Axial Cracks: Sealing Bladder Considerations 6-4 6.3 Leakage Modal 6-5 6.4 Tube Burst Capability 6-11 6.5 L'ak Before Break (LBB) Capability 6-15
. 6.6 SLB Leak Rate Estimates 6 18 6.7 Conclusions 6-19 T.0 - TUBE VIBRATION ASSESSMENTS 71 7.1 Overview of Analysis Methods and Models 7-2 7.2 Damping for Dented Tube Conditions 7-12 7.3 Interpretations and Recommended Tube Damping 7 29 7.4 Umiting U/fnD for Predicting the initiation of Fluideiastic 7 32 Vibration in Tube Arrays in Water -
7.5 Results from Normal Operation Straight Leg Tube Vibration 7-38 Evaluations 7.6 Influence of Ugaments on Tube Vibration and Crack Propagation 7-40 7.7 Crack Growth Rates Associated With Vibration Induced 7-42 Crack Propagation 7.8 Leak Rate Versus Time for Vibration Induced Crack Propagation 7-44 7.9 Summary and Conclusions 7 46 7.10 References 7 48 8.0 EVALUATION OF POTENTIAL CRACK PROPAGATION UNDER SLB FLOW 81 CONDITIONS 8.1 Steam Line Break Loads Analysis 81 . 8.2 Vibration Analysis Methods for Steam Line Break Conditions 85 8.3 Straight Leg Large and Small Steam Line Break Vibration Evaluation 8-6 8.4 Conclusions 8-7 li
.)
TABLE OF CONTENTS (continued) Section Ilite Eagg 9.0 COMBINED ACCIDENT ANALYSIS 91 , 9.1 SSE Analysis 9-1 9.2 LOCA Analysis 93 9.3 SSE + LOCA Effect on Umiting Tube Condition 95 l 9.4 SSE + SLB Effect on Limiting Tube Condition 98 l l l t lii
x F
1.0 INTRODUCTION
This report describes a R.G.1.121 evaluation for North Anna 1 tube degradallon b' identified by eddy current inspection following Cycle 8 operation. Degradation identified in this outage included initial identification of circumferential cracking at r tube support plate (TSP) edges as well as previously identified PWSCC WEXTEX expansion transition circumferential cracks and PWSCC axial cracks extending outside + of the TSPs. The PWSCC cracks at the TSPs have had previous R.G.1.121 evaluations ; and are included in this study only for demonstration of burst capability and leak before l break (LBB) margins. The principal emphasis of this report is an evaluation of the circumferential cracks for operating cycle length considerations. Utilizing inspection data from the 1989 and 1991 outages, end of cycle (EOC) crack angle projections have been performed to 1 i provide crack distributions for the tube integrity assessments. The EOC projections are j based on cracks left in service below eddy current detection thresholds and include growth to the EOC and eddy current uncertainties for consistency with R.G.1.121
]
guidelines. Both deterministic and probabilistic evaluations are performed. A I deterministic evaluation was previously performed in February,1991 to support return to power and emphasized tube burst margins and leak before break considerations. This report extends the deterministic evaluation to include tuce vibration and accident condition evaluations for circurnferential cracks at the WEXTEX transitions and straight leg TSP Intersections. Circumferential cracking has not been identified above the 5th TSP. Thus U bend tube considerations are not required for the circumferential crack evaluation. To further assess the potential risk associated with the low probability, large crack angle tail of the EOC distributions, a probabilistic evaluation of through wall crack angles associated with vibration induced crack propagation has also been performed. Circumferential cracks can be considered " strong" for tube burst considerations in that through wall crack angles of [ ]bc meet three times normal operating pressure differenfials and through wall angles greater than [ ]b,c meet SLB burst pressure cor.ditions. North Anna inspection practices and plugging of all circumferential cracks reduce the likelihood of tube burst at SLB conditions to extremely low levels. Thus the principal objective of this report is to extend the avaluation of circumferential cracks to include potential crack propagation by tube vibration under normal operating and 1-1
a SLB flow conditions. This evaluation extends efforts of the WEXTEX Subgroup of the f Westinghouse Owner's Group (WOG) to include circumferential cracking at TSP i ,= locations, Evaluation of loads under accident conditions including combined events are also included in this report. - The EOC cred. hngle projections are based upon 8x1 and RPC inspections and represent EOC RPC angle projections. The RPC data provide very limited data on crack depth and - { generally overestimate the' length of the through wall portion of' the indicated [ circumferential crack length. Tube vibration assessments are particularly sensitive to through wall crack angles as small ligaments such as 10% remaining wall or ligaments
- between microcracks add substantially to tube stiffness and can effectively eliminate tube vibration concems. The tube vibration analyses to assess potential propagation of corrosion cracks are most confidently performed for assumed through wall cracks.
. Thus it is necessary to bridge the information gap between RPC crack angles and
- potential through wall crack angles which, if present, would be considerably smaller than the RPC crack angles. Based on the extensive North Anna inspections for circumferential cracks, it is reasonable to expect little through wall crack penetration in Cycle 9 'particularly with operation limits on leakage of 50 gpd. However, it is the -
. goal of this evaluation to provide conservative but realistic assessments of potential [ crack propagation due to tube vibration. To this end, two evaluations of vibration induced corrosion crack propagation are performed in this report. The deterministic evaluation is based upon a segmented crack model for WEXTEX cracks and conservative
- estimates of the fraction of RPC angles that might be through wall for ODSCC. This evaluation is supplemented by a probabilistic assessment of the potential propagation of through wall cracks. The probabilistic assessment has been very conservatively based l on an assumption that up to 10% of the RPC crack angles are through wall, in reality,
; it is judged that this percentage is notably less than 1%. Mechanical damping tests for dented tubes with WEXTEX expansions ano for corrosion fatigue cracked tubes have been r performed to improve input to the vibrat,on analyses.
r To perform an assessment of potential propagation under SLB flow cond% ions, small and O large break transient analyses were made to obtain more detailed flow distribution data than previously'available. These flow distributions are used to assess tube vibration ,
- j. during the short term, high flow conditions of the SLB event. _ Finite element analyses of
- Series 51 S/Gs have also been performed to improve definition of tube and TSP loads in l a seismic event. In addition, Series 51 TSP crushing tests were performed to refine 1-2
assessments of tube deformhtion under combined LOCA + SSE events. These data are . Incorporated into the accident condition ar.alyses. Potential leak rates under SLB conditions are also estimated for comparisons with acceptable limits. Section 2 of this report ir'tegrates the detailed results of this report into a summary tube integrity assessment. Overall conclusions of this evaluation are also given in Section 2. North Anna pulled tube examination results are described in Section 3. Results of the 1991 inspection are given in Section 4. Section 5 develops the EOC crack angle projections and probabilities associated with exceeding acceptable threshold crack angles. The tube integrity assessment for tube burs. 1 LBB is developed in Section 6. Evaluations of tube vibration are included in Seu . Sections 8 and 9 provide the evaluations for potential crack propagation under SLB flow conditions and the accident condition analyses. e a 1-3
2.0-
SUMMARY
ANDCONCLUSIONS This report describes a R.G.1.121 evaluation for North Anna Unit 1 steam generator tube degradation identified by eddy current inspection following Cycle 8 operation. Circumferential tube cracking at the TSPs was identified at this outage and, together with circumferential cracking at the WEXTEX transitions, forms the primary subject for this evaluation. A c'aterministic evaluation was performed to support return to power and emphasized tube burst and leak before break considerations. This evaluation, although incomplete with regard to flow induced vibration and accident conditions,
' indicated that full cycle operation was feasible. This report extends the evaluation to include tube vibration and accident condition evaluations for circumferential cracks at the TSPs and WEXTEX expansion transitions. The scope encompasses a complete R.G.
.1.121 evaluation. in addition to evaluating expected EOC crack sizes, the evalu3 tion is .
extended to include assessment of the low probability of large crack sizes and their associated impact or potential risk on plant operation. The evaluations are performed principally for the very conservative assumption that the detectable crack sizes are uniformly through wall cracks. It is then shown that the probabilities for potential crack propagation are small, even for the assumed through wall cracks. For the expected crack morphology of ligaments between cracks or partial depth cracks, large margins exist against even the maximum projected crack angles, in this section, the overall conclusions of this report are provided together with a summary of the supporting test and analysis results. 2.1 Overall Conclusions From the results of this study, it is concluded that: Based on the evaluations for normal operating and accident conditions, the EOC crack considerations satisfy R.G.1.121 criteria and full cycle operation is acceptable. 9 Reg. Guide 1.121 guidelines for tube burst at 3aP N .O. are satisfied for greater than . 95% of the projected EOC indications. Even the most limiting projected indications provide burst capability within a few hundred psi of the 4300 psi guidelines for 3APN.O. and very large burst margins for SLB conditions. Review of the 1991 2-1
inspection results has indicated no occurrences of mixed mode (intersecting axial-and circumferential) cracking such that burst capability is dictated by the limiting " axial or circumferential cracks.
- The administrative leak rate limit of 50 gpd provides large leak before break margins against tube burst at 3.iPN .O. for assumed through wall cracks and against burst at SLB conditions even for leak lithiting segmented crack morphologies.
- Corrosion crack propagation due to tube vibrations has a very low probability of occurrence (-1% chance of 1 tube in 3 S/G at EOC 9) even under the conservative assumption that 10% of the projected RPC crack angles represent through wall circumferential cracks, Applications of crack models representative of crack morphologies found in pulled tubes [
ja.c result in determinit, tic margins against crack propagation even for the most limiting EOC crack angles.
+
Even under the low likelihood event that crack propagation due to tube vibration , should occur, the North Anna leakage monitoring system and administrative procedures would provide detection and plant shutdown prior to a large leakage , event approaching a tube rupture. The North Anna leakage monitoring procedures would initiate plant shutdown following leakage exceeding 50 gpd in a single S/G with plant shutdown completed within 2 hours. Accident conditions including SSE, LOCA, SLB flow conditions and combined events
.(LOCA' + SSE) were evaluated and found to not result in propagation of tube degradation including circumferential cracks at WEXTEX transitions and affected - TSP intersections. +
Potential leakage at SLB conditions is expected to be bounded by the leakage from the limiting crack leaking at the 50 gpd adminis'rative leakage limit. Evaluated at 95% confidence limits, the associated SLB leakage would be less than [ ]b,c, [ 2-2 L i . - ..
1 2.2 Summary nf Principal Results The RPC probe, which is used as the basis for crack angles in this study, measures the crack angle above a detectable threshold depth. The threshold depth for an RPC is typically an average depth of about 40% which could increase to about 50% for OD circumferential cracks. The through wall portion of the indicated circumferential arc length cannot be reliably determined from an RPC inspection. Crack angle growth is measured c.s the growth in crack angle for depths greater than the detection threshold I ja.ce. End of cycle crack projections obtained by summing detection threshold, growth and EC uncertainty represent the crack cngle over which penetration is greater than 40 to 50% depth. The crack morphology for the EOC crack argles with depths greater than 40% can be quite complex and is typically different between PWSCC and ODSCC cracks. For tube integrity analyses of the crack angles >40% depths, it is necessary to define models representative of the crack morphology based on pulled tube examinations. Crack models representative of PWSCC axial cracks have been discussed extensively since
~
1986. Circumferential PWSCC cracks and ODSCC circumferential cracks are the principle types of degradation analyzed in this report. However, the tube vibration analyses are directed toward determining the crack threshold angle at which vibration amplitudes are sufficient to result in potential crack propagation. Crack propagation results are most confidently performed for through wall cracks, particularly for this report wherein all straight leg tubes and tube spans are evaluated to search for tubes sensitive to propagation of circumferential cracks. The resulting through wall crack angle for propagation, with associated stress amplitude and tube stiffness, is then compared with relative values for propagation of the crack model to estimate the crack angles for propagation of realistic crack morphologies. For example, the tube stiffness [ ja,c.e. Based on these methods, the overall analyses develop the threshold crack angles for the PWSCC and ODSCC crack models representing RPC measured crack angles and for assumed through wall crack angles. 2-3
The propagation threshold angles for the crack models are then compared with projected-EOC crack angles at a 95% cumulative probability to obtain a deterministic tube , integrity assessment. To further assess the risk of crack angles larger than the 95% values,'a probabilistic evaluation has been performed. This evaluation estimates the . probability that one tube in the 3 S/Gs exceeds the thrsahold craci; angle for crack propagation by vibration. A comparison of RPC crack angles with the tube examination results for the North Anna pulled tube is shown in Table 2.1. For PWSCC in WEXTEX transitions, the RPC angles agree well with the destructive examination results. The small ligaments separating the average 21' to 33' microcracks resulting from the multiple init!ation sites are too small for RPC detection. Thus an appropriate model for PWSCC is a segmented crack model assumed to have 30* to 49' through wall cracks separated by ligaments. The ligaments are sized in the model to remain elastic at normal operating pressure conditions. The burst pressures for two North Anna pulled tubes with WEXTEX circumferential indications exceed the burst pressures obtained for the segmented crack model, thus providing support for this crack model. No significant cracking outside the primary crack was found for the PWSCC cracks. Thus the segmented crack model can assume undegraded tubing around the circumference bounding the segmented macrocrack. For ODSCC circumferential cracks at TSPs, the North Anna experience of Table 2.1 indicated RPC results overestimating crack angles for significant through wall cracks and sometimes underestimating the crack angles with 40-90% depth. Due to the limited pulled tube data base for ODSCC circumferential cracks, a conservative ODSCC crack model is defined from the results of Table 2.1. Since the 51% average depth crack was not detected, it is assumed that a 50% deep crack can exist outside the primary crack. The data indicate that the through wall crack angle is less than 60% of the measured RPC crack' angle. Thus, pending additional pulled tube data for ODSCC circumferential crac'ts, a simplified ODSCC crack model can be defined as [ Ja,c.e. With this model, vibration analysis results for a through wall crack angle plus 50% deep crack can be compared with a projected EOC RPC angle by dividing the analysis angle by [ ]a,c.e. The 24
simplifications of the ODSCC model need not be applied for tube burst considerations as the largest projected EOC crack angle meets R.G.1.121 burst critoria even if the total crack angle is assumed to be through wall. It can be noted that the defined ODSCC model is more conservative for the present analyses than alternatives such as assuming average RPC crack depths up to 90%. Scoping analyses indicate that crack depths up to at least 90% would not result in vibration induced crack propagation. Utilizing the above crack model, Table 2.2 summarizes the results of the deterministic analyses. The projected EOC RPC crack angle at 90% probability represents the sum of detection threshold plus growth to EOC and EC uncertainty. Each of these three terms is evaluated at the cumulative 95% value of the respective distributions developed in this report. The projected EOC crack angles are notably lower than acceptable crack angles for tube burst. For potential crack propagation due to tube vibration, only the RPC crack angle of [ )b,c for initiation of turbulence propagation at the bottom of the first TSP is less than 360'. The [ )b,c angle exceeds the projected EOC RPC angle of [ ]bc. All straight leg tube spans up to the top TSP were evaluated for tube vibration. No tubes were found to have the potential for crack propragation up to the largest through wall crack angle of [ ]b,c evaluated. For the WEXTEX transitions, the bending stiffness provided by the ligaments in the segmented crack model are sufficient to preclude crack propagation even up to a 3GO' macrocrack in the most limiting area of the steam generator. Although not shown in Table 2.2, the PWSCC axial cracks extending outside the TSP were also evaluated for tube burst capability. Utilizing the crack length distribution found in the 1991 inspection, 95% of the indications had crack lengths less than 0.59 inch. The burst capability at three times normal operating pressure capability is [ ]b,c for a segmented axial crack. Thus the R.G.1.121 guideline is satisfied also for the PWSCC axial crack distribution at the 95% level. Overali, the deterministic results of Table 2.2 satisry the guidelines of R.G.1.121 for End of Cycle 9 operation. Accident conditions were also evaluated for potential propagation of circumferential cracks in straight leg tubes and found to result in negligible potential for crack propagation. The evaluation included combined accident conditions such as LOCA + SSE. To assess the potential for propagation due to crossflow in the first pass under SLB flow conditions, small and large break SLB transient analyses were performed to develop 2-S l
. more detailed mapping of flow distributions than available from prior SLB analyses.
These evaluations show small vibration amplitudes even for cracked tubes with up to . [ ]b,c through wall _ cracks over the short time frame of high velocities in an SLB event. To further assess the potential risk of operating with crack angles in the largest - 51% of the distribution, a probabilistic analysis was performed to estimate the number of tubes exceeding a threshold angle for crack propagation. When this number is less than one, it can be defined as the probability that one tube in 3 S/G at the EOC exceeds a threshold crack angle for propagation. Emphasis in this assessment is on the potential for propagation due to tube vibration. With tube burst capability at SLB conditions exceeding [ ]b,c for through wall cracks, the likelihood of the burst of a circumferential crack is exceedingly small given the North Anna Unit 1 inspection practices with 8x1 and/or RPC inspection of potentially susceptible regions. Thus the principal concern for circumferential cracks is the potential risk that the corrosion induced cracP will propagate as a consequence of tube vibration. All tubes are expected to be fluiow.astn, ally stable with small turbulence amplitudes in the undegraded condition with dented TSP latersections. However, large through wall circumferential cracks reduce the tube stiffness and increase the potential for vibration propagation of corrosion cracks. Thus the major emphasis of this operating cycle evaluation is applied - to assessing the risk for vibration Induced propagation of a circumferential corrosion crack. This evaluation builds upon R.G.1.121 evaluation efforts performed for the WEXTEX Subgroup of tho Westinghouse Owner's Group (WOG). This effort is extended to include circumferential crackr at the straight leg TSP intersections in this report. Corrosion cracks at the top (7th) TSP intersection have not been found at North Anna 1 i. i- so that U bend vibration issues are not applicable. Only one axial crack at the top TSP l' edge was found in the 1991 inspection even though significant numbers have been found at lower TSPs over the last three cycles. The highest TSP found with a circumferential crack was the 5th TSP :n 1991. It is expected that the reduced tube temperatures at higher elevations in the bundle significantly reduce the occurrence of corrosion degradation at the higher elevation TSPs. l_ Projected EOC RPC crack distributions are used to estimate the probability of a crack
~
h exceeding a threshold angle for propagation. The vibration analyses yield the number of tubes potentially subject to crack propagation at a threshold through wall crack angle. ' i If the PWSCC and ODSCC crack models are applied, the crack angles for vibration propagation given in Table 2.2 exceed the maximum projected crack angl6 such that propagation probabilities are exceedingly low. An alternate approach to the crack 2-S 1
l l models is to estimate the probability that the projected EOC RPC crack angles' represent
. through wall crack angles. This probability is likely to be very low and notably less than 1% althouf there is insufficient pulled tube data to fully substantiate this value.
For this study, a 'ry conservative 10% chance that the RPC angle is through wall supports low probability estimates for vibration induced crack propagation. The probability that any tube has a circumferential crack is obtained by [ I
]a,c.e. For resulting numbers less than one, the final result can be interpreted as the proba :lity of one tube in the three Nort" .ina-1 S/Gs exceeding the threshold crack angle for vibration propagation.
Table 2.3 summarizes the results of the probability analyses. For tube burst considerations, there is a negligible 15% chance that one WEXTEX through wall crack angle would exceed the [ ]b,c 3 a PN.O. guideline of R.G.1.121 for the largsst q projected RPC crack angle of [ }b,c. For ODSCC at TSPs, the maximum projected crack angle of [ ]b,c is much less than the [ ]b,c guideline. The WEXTEX vibration results are separated into values for the central sludge region where more than 99.5% of the circumferential indications have occurred and the peripheral region which has the highest crossflow velocities. In the central region, with. propagation' threshold angles of [ ]b,c for turbulence and fluideiastic vibration, respectively, there is less than a 0,1% chance for a crack probability tail at large crack angles for the WEXTEX crack angle distribution. In the peripheral region, the through wall crack angles for propagation are [ ]bcfor turbulence and [ ]b,c for fluidelastic vibration. The corresponding probabilities for these through wall crack angles are 0.38% and 0.14%, respectively. For ODSCC at TSPs, ihe threshold through wall angle of [ ]b,c for fluideiastic , vibration represents the largest angle analyzed and significantly exceeds even the low probability tail for projected crack angle distribution. This results from the fact that peripheral region crossflow velocities are high near the tubesheet and even the cracked tube remains stable for very large through wall crack angles. All straight leg tube 2-7
spans up to the top TSP were evaluated including cracks at the top and bottom of the mid span tubes; cracks up to [ )b,c are not propagated by vibration. Thus propagation of ODSCC circumferential cracks at TSPs by fluidelastic vibration is not expected. The turbulence results indicate propagation of through wall crack angles of ( ]b,c without.a 50% deep additional crack, and [ )b,c with an additional 50% deep crack around the tube circumference. The [ ]b,c through wall crack has a low (<1%) estimated chance of occurrence. Overall, the results indicate a very low or about 1% chance of one tube in the three North Anna 1 S/Gs propagating a crack by turbulence induced vibration at the end of cycle 9 operation, Even if this event should occur, the rate of crack propagation is very low. The North Anna leakage monitoring system would easily detect this event, initiate plant shutdown at the 50 gpd administrative leak limit, and plant shutdown would be completed without a large leakage event. The potential for propagation due to fluidelastic vibration is limited to the WEXTEX region with a negligible probability of about 0.1% at EOC 9. Even if this event were assumed to occur, crack propagation rate analyses demonstrate adequate time for plant shutdown prior to a tube rupture. . 2.3 Summary of Report Sections , This sections provides a summary of the evaluations provided in the following sections of this report. Puned Tube Evamination Aesults 1991 Tube Exam - ODSCC Circumferential Cracks at TSPs Tube R11C14_ was pulled from S/G B in 1991 to assess circumferential indications found by 8x1 and_ RPC probe inspections at the first TSP. Circumferential cracks, predominantly ODSCC with some IGA components, were found at locations corresponding to the top and bottom edges of the TSP. At the top edge, two cracks of 124* and 113' - with depths greater than 40% including a maximum through wall length of 90* were found. Crack depths of 8% to 33% were found between the two deeper circumferential
- cracks. At the bottom edge of the TSP, circumferential cracks of 135' and 112' above 40% depth were found with no significant crack depth between these cracks. The maximum depths of the two cracks were 98% and 61% The cracks were centered near 2-8
1 l the apex of the ovalized dented (28 mil ovality) tube. At the WEXTEX transition, a
- semicontinuous network of shallow (21% maximum depth) OD circumferential and i axial microcracks was found. The maximum axial crack length was 0.04 inch and maximum continuous circumferential extent was about 80'.
1987 Tube Exam - WEXTEX PWSCC Circumferential Cracks in 1987, two tubes were pulled for circumferential indications in the WEXTEX transition region. One tube was found to have a PWSCC circumferential macrocrack of 176' with up to 90% depth. The second tube had a 128' PWSCC macrocrack with through wall penetration. The macrocracks for both tubes showed multiple initiation sites with numerous parallel microcracks, many with remaining ligaments. Burst pressures for both tubes were close to that for undegraded tubing. 1985 Tube Exam PWSCC Axial Cracks at TSPs in 1985, two dented tubes were pulled to examine distorted eddy current indications at the edges of the TSPs. Both tubes showed PWSCC axial cracks extending outside the TSP. One tube showed two axial cracks separated by about 180* at the minor diameter at the ovalized tube. The most extensive axial crack pattern extended from 0.8 inches above the TSP to 0.2 inch below the TSP. Maximum depths of the macrocracks were 50%, 85 % and 100 %. The PWSCC cracks showed multiple microcracks comprising the macrocracks. ODSCC within the TSP crevices was also found with depths ranging from 15 to 40% through wall. Based on the pulled tube crack morphology, PWSCC cracks can be modeled as segmented cracks with deep or through wall crack segments (aspect ratios of 4/1 to 6/1) separated by ligaments with multiple segments adding to the total crack angle. ODSCC is modeled as a deep or through wall crack with a potential partial depth crack around the remaining tube circumference. Conservatively, the ODSCC analytical modeling assumes through wall crack angles with and without up to a 50% crack depth around the remaining tube circumference. These models are based on the North Anna pulled tubes, together with laboratory data and tubes removed from steam generators at other plants. The data base on ODSCC circumferential cracks is limited and the above ODSCC model may be conservative for crack growth above the detection thresholds in one operating cycle. 2-9
Eddy Current Data Review The 1991 inspection included: 100% full length bobbin coil inspection; 8x1 inspection for 100% of the S/G B hot leg,100% to 4th TSP and 20% to the top TSP hot legs of S/Gs A and C; 100% RPC inspection of hot leg WEXTEX transitions; and RPC confirmation of potentially pluggable bobbin coil and 8x1 indicatior's. The results show 216 circumferential Indications in the WEXTEX transition. Based on phase angle evaluations, 200 are classified as PWSCC,15 cannot be distinguished - between ID or OD origin and 1 is possibly ODSCC. The largest single crack angle is 247*. The circumferential cracks are predominantly located below the top of the tubesheet with an average distance 0.07" below the tubesheet top and 0.12" above the bottom of the transition (highest contact point in the tubesheet). The totals include 36 cracks classified as multiple circumferential indications (MCis) based on the RPC amplitude returning to the null point between indications. All but one of the MCis are centered below the top of the tubesheet for which the
~
tubesheet constraint helps to prevent tube burst. The one tube with an MCI above the tubesheet had crack angles of 78* and 92* separated by 65* and 125' ligaments. The
- smallest ligament measured had a 7* arc length. Tests on EDM notches show that ligaments greater than [ ]b,c are necessary for the RPC amplitude to return to the
- null point and clearly resolvo separate cracks. Thus the small ligament arc lengths can be assumed to represent . east a [ }b,c ligament if the cracks are through wall or, more likely, the cracks are not through wall at least near the ends of the cracks. Based on the MCis being pRdominantly located below the top of the tutesheet and being separated by large ligaments if through wall, the single COls (circumferentially oriented indications) are more challenging to structural integrity than the MCis. For this reason, only the single circumferentialindications are used to project EOC crack
- angles for evaluating tube integrity considerations such as tube burst, vibration, etc.
A total of 108 circumferential indications were found at the edges of the TSPs in the , 1991 inspection. All were tietected by the 8x1 probe and confirmed by the RPC probe. An additional 433 TSP intersections were inspected by the RPC probe and none were found to have circumferential indications which supports the 8x1 detectability at TSP locations. Both indications showing crack-like signals and those with a volumetric appearance are included as circumferential indications. More than 90% of the 2-10
circumferential cracks were found at it'e first two support plates with the remainder . extending up to the 5th TSP. The largest slagte crack reported was 212*. Nine (9) indications were found at the bottom of the 1st TSP out oi 45 indications at this plate. None of these tubes had circumferential indications at the WEXTEX transition. Five tubes were found to have both circumferential cracks and axial cracks outside the same edge of the TSP. These indications were found to be azimuthally separated by 80* to 120* and were found at highly ovalized, dented intersections. It is expected that the ODSCC circumferential indications are located at the major diameter of the ovalized tubes and the PWSCC axial cracks at the minor diameter found for the North Anna-1 pulled tubes. Overalt, no intersecting mixed mode cracks were found and similar azimuthal separation is expected in the future. Thus, mixed mode cracking is not expected and therefore has not been considered in the present tube integrity evaluation. Circumferential crack detection thresholds for the 8x1 and RPC probes were developed based on North Anna experienco, industry experience and the 8x1 probe features. Dutection distributions are denned as 0% at 35' to 100% at 75* for the 8x1 probe and 50% at 23* to 100% at 50* for the RPC probe. These detection thresholds are used in projecting crack distributions at the end of the current operating cycle. The distribution of axial crack lengths outside the TSPs has been essentially the same over the last 3 inspections. The maximum axial crack length in 1991 was 0.82 inch which is slightly shorter than prior outages. The number of indications decreases sharply with TSP elevation and associated primary temperature reduciions, it is anticipated that the 1992 EOC length distributions will be similar to the 1991 distributions. Projected EOC Circumferential Crack Distributions EOC crack angle distributions have been developed by Monte Carlo sampling methods combining distributions for: EC detection thresholds developed from the eddy current review, industry experience and 8x1 probo characteristics; growth rates developed from 1989 and 1991 inspection results; and EC uncertainties developed from fitting 19918x1 coil hit results to the measured RPC crack angles. This projection thus represents EOC RPC crack angle distributions. The data base and Monte Carlo methods have been checked by utilizing the 1989 data to project the 1991 RPC inspection 2-11
results. The results show good agreement between projections and actual crack angles for predicting the large crack angle distributions. The WEXTEX projections are . somewhat conservative and overestimate the maximum 247* actual angle with a [ ]b,c projection. The TSP 1991 projections show excellent agreement and predict - the actual maximum angle of about. 215*. Crack angle distributions projected for the 1992 EOC Indicate maximum crack angles of about [ ]b,c for the WEXTEX transition and about [ ]b,c for the indications at TSPs. These distributions are used to calculate the probability that a circumferential crack angle is greater than a specified angle. Conservative estimates for the number of indications at the EOC are made using log normal projections. It is estimated that the number of circumferential EOC indications will be less than 150 at the WEXTEX transitions and 100 at the TSP locations summed over all 3 S/Gs. Probability estimates for indications exceeding threshold crack angles are developed from the EOC projections and the tube vibration results as described in Section 2.2. Tube Intecrity Assessment- Tube Burst and LBB Tube burst tests for EDM notch simulations of circumferential cracks were performed for varying conditions including: through wall slots with and without lateral restraint at the TSP elevation, segmenter' crack conditions, circumferential through wall slots combined with 50% deep slots over the remaining tube circumference and simulated cracks at the top and bottom of a span. The results show that the presence of the TSP increases the burst capability significantly over that without lateral motion restriction. Segmented cracks show the expected increase in tube burst capability over that of uniform through wall cracks. Including 50% deep cracks in addition to the through wall crack does not affect burst capability at 3 times normal operating pressure differentials but slightly decreases burst capability at SLB conditions. Circumferential cracks at the top and bottom of a tube span do not significantly change the burst capability compared to a crack at only one end of the span. The test data are adjusted for temperature and lower material tolerance limits for application to the tube
~
integrity assessments. Table 2.4 summarizes the tube burst capability for circumferential and axial cracks with various crack morphologies. For TSP locations, the maximum projected EOC circumferential crack is [ ]b.c which provides margins against the 3aP burst capability of [ jb,c for through v all 2-12
cracks with a 50% deep crack around the remaining circumference. The projected WEXTEX crack angles show a maximum angle of [ ]b,c which results in margins against 3AP burst ([ )b,c) for the expected segmented crack morphology and only t.vo tubes would exceed 3AP at [ Ja,c for assumed through wall cracks. Large WEXTEX margins remain against the SLB burst capability of [ ]b,c for the assumed througn wall cracks. The distribution of axial crack lengths above or below the TSP edges has been essentially the same over the last 3 cycles with a maximum projected EOC length of [ ]b,c. This length provides large margins agalnst burst at SLB conditions which requires a [ ]b,c crack for the expected segmented crack morphology. A segmented burst capability of 4000 psi is expected for the [
)b,c ' crack compared to the 3AP burst capability of 4300 psi. Thus it is concluded that the maximum projected EOC cracks provide acceptable margin against tube burst. Projected average EOC cracks would exceed 3AP capability even assuming uniformly through wall cracks.
The leakage models uwd to demonstrate LBB and to estimate potential SLB leakage for axial cracks have been compared with experimental data from pulled tubes and laboratory tests. This comparison has permitted establishing both the mean (nominal)
- and 95% confidence levels on the leak rate predictions. For leak rates from circumferential cracks, the leak rate correlation has been established as a lower bound for the limited experimental data available. The LBB rationale provides assurance that a rogue tube with greater than projected crack growth will not result in unstable crack growth or tube burst in the unlikely event of a limiting accident such as a SLB event.
North Anna 1 has implemented an administrative leak rate limit of 50 gpd. For assumed 1 uniform through wall cracks, the 50 gpd limit and nominal leak rates provide for plant shutdown at crack lengths (axial and circumferential) less than that associated with 3AP tube burst capability and thus provide considerable LBB margin. At the lower 95% certainty on leak rates for through wall cracks, 50 gpd corresponds to a crack length of [ ]b,c inch which is much less than the [ ]b,c for burst at SLB conditions. For segmented circumferential cracks, the 50 gpd limit also provides large LBB margins against the 3AP crack angles. For axial segmented cracks, the 50 gpd limit corresponds to operating leakage for a [
]b,c macrocrack which provides
. LBB for SLB conditions with large margins against the [ ]b,c macrocrack for burst in an SLB event.
2-13
l J The operating leak limit of 50 ' ! 31so limits the potentialleakage in a SLB event. The limiting crack that has a 50 pp soak at lower 95% certainty is the l }DC .. axial crack at the TSP edges. ! Evaluating the SLB letkage at the upper 95% certainty results in an upper bound SLB j leali rate of [ ]b,c. Another SLB leakage calculation was performed to bound ; potential leakage in the unlikely event that a large population of tutss have either , through wall cracks or near through wall cracks that do not leak during normal operating conditions but may leak during a postulated SLB event. This method is
. inconsirtent with the operating leakage limits (50 gpd) in place at North Anna which lead to the [ jb,c limit previously identified. This analysis appiles the i segmented crack model for PWSCC adal and circumferential cracks and assumes the ODSCC :nodel with [ ja c. The results indicate a SLB leak rate of less than [ ]b.c, Tube Vibration AMessment Mechanical damping tests were performed for dented tubes with WEXTEX exr.ansions and for corrosion cracks rnechan!cally cvcied to tatigue remair.ing lipments. These tube damping results chart.cterize the tube damping for input to the tube vibraticit '
amplitude. Damping measurements are made as a function of vibration amplitude to ! facilitate calculation of vibration amplitudes for fluidolastic and tutbulence excitation. A literature review was conducted to establish U/fnD = [ ]b,c as a lower bouno
- ihreshold for fluidolastic instability at small values of the mtiss damping parameter.
Utilizing the test and evaluation results, tube vibration analyses were performed for eveiy straight leg tube in the tube bundle. The detailed analyses were performed for uniform through wall crack angles to estimate the cracP. angie poteritially resulting in crack propagat!on. Comparative analyses of tube stiffness and crack tip stress intensities between_ cracks with _ ligaments and through wail cracks indicate that small remaining ligaments can prevent [ , ja,c, 2-14
._..__._.________._m____ - _ _ _ _ _ _ _ _ _ . _ - . _ _ _ _
i I f t Limiting through wall angles for crack propagation were determined for overy straight i
. log span and tube in the S/G. For WEXTEX transition circumferential cracks, the !
t limitirw throuch wall cra,:k angles are [ 10.c ier turbulence and [ l.cfor b ;
=*
fluidolastic excitatfor cd are located in the peripheral tube region of tho S/G. The f associated probabl%es for these through wall crack anglos are about [ f
]b,c, respectively. For WEXTEX cracks in the contra'. sludge region, the limiting
{ crack anglos increase to [ ]b,c (0.1% probability) for fluidolastic vibration and l , [ ]b,c (.0% probability) for turbulenco excitation. The limiting crack ongles are } r reduced for combined circumferential cracks at the WEXTEX transition and at the bottom of thn 1st TSP. However, the probability for occurrence of those combined ; crack conditions is <10 5, i For circumferential cracks at TSP edges, the limiting through wall crack angles for propagation are ( ]b,c (0.95% phbability) for turbulence and [ ]b,c (.0% j precability) for fluidolastic vibration. No casos were found that resulted in fluidolastic instability for TSP circumforential cracks up to the largest angle of [ b l.c used in f the analysos. This results from the fact that crossflow velocities in the inlet pass are f high near the tuboshoot and low at the bottom edge of the 1st TSP, ! Crack propagation rate analyses woro performed for the limiting through wall crack angles to obtain crack angle and leak rate versus time. These results provido a basis for ovaluating the North Anna leakage monitoring system and proceduros. Turbulent driven
- crack propagation proceeds slowly and can readily be detected by loakage monitoring in f ample time to implement plant shutdown if the 50 gpd limit is exceeded. For f
fluidolastically driven crack propagation initiated at a [ ]D C crack angle, it takes j about [ )b,c to proceed to a tube rupture. Since North Anna providos ! continuous leakage monitoring and 3hutdown capability within 2 hours of exceeding the f 50 gpd administrativo leak limit, the fluidotastically driver, crack propagation would f
- also be deter,ted and the plant shutdown prior to a tube rupture, j
Overali, the vibration analysis results indicate that remaining ligaments in corrosion $ cracks are likely to prevent crack propagation due to tube vibration. This is particularly true for WEXTEX PWSCC cracks for which ligaments are found betwoon the i microcracks in pulled tube examinations. Thus WEXTEX crack propagation is unlikely [
- due to crack morphology considerations. Fluidolastic instability is not expected fer
{ circumferential cracks at TSPs due to flow distribution considerations. If large through 2 15 :
wall cracks and associated crack propagation should occur (<1% probability of one tube in 3 S/Gs at EOC), the most likely condition would be turbulence driven propagation of a circumferential crack at the bottom of the 1st TSP. This would result in a slowly increasing leak rate and prompt shutdown without a large leakage event. The North . Anna leakage monitoring system is capable of del ^cting leakage and shutting down the r plant prior to a tube rupture even assuming fluidelastically driven crack propagation. Potentist for Ornck Ptoonant!on Under SLB Flow Conditions TRANFLO SLB analyses for small and large breaks were performed to obtain crossflow conditions in the first pass to assess the potential for crack propagation under GLB conditions. These analyses modeled radial and axial nodes to characterize the flow magnitudes and distributions for vibration assessments. Previously available SLB anclyses did not provide adequate mesh definition to def:ne the crossflow conditions for the flow that passes from the tube bundle up the downcomer in the early seconds , following the break. The analytical results show very low vibration amplitudes for through wall circumferential cracks as large as [ . ]b c such that nyligible potential exists for crack propagation in the straight leg of the tubes under SLB conditions.
~
Combined Accident Condition Evaluntion A seismic analysis specific to a Series 51 S/G wa:; performed to obtain TSP loads and the t displacement time history response of the tube bundle. LOCA loads were developed for small breaks (accumulator lines, etc.) as North Anna has quallflod leak before break for primary piping. The combined LOCA + SSE loads were used to assess potential crack growth for a 185* circumferential crack at the TSPs. The results indicate minor crack growth to [ }b.c such that the tube will maintain its integrity under combined LOCA
+ SSE loading. The combined loads were also used to evaluate potential crushing of the TSPs with associated tube deformation. TSP plate crusn test data recently obtained fcr Series 51 TSPs were incorporated into the analyses. The analytical results show that no tubes will undergo permanent deformation in the event and thus no significant secondary to primary leakage in a cracked tube is expected for the combined accident conditions.
l 1 2-16 " i
._ _._m I
i I Table 2.1 l Comparison of RPC Crack Angles with Pulled Tubo f' Examination Results Pulled Tubo Results ! Locatlen on Tube RPC Anc'o Anote >40% MM.Denth Thruwall Ane!o [ f Tube R11C14. TSP #1(1): i Top Edge of TSP: ODSCC [
- Crack i 158' 124* 100% 90' [
+ Crack 2 74' 113' 100% (90% avg.) -10' & 20' [
Bottom Edge of TSP: ODSCC f
. Crack 3 80* 90* 135' 98% (74% avg.) -
- Crack 4 NDD -112* 61% (51% avg.) -- l Tube R17C31:
WEXTEX Transition: PWSCC 150' - 14 0. ( 2 ) 90% - Tube RiBC36: WEXTEX Transition: PWSCC 120' -118*(3) 100% 53*(4) Notos:
- 1. Depths from destructive examination are detailed in Section 3.
- 2. Total crack angle was 176* and included 2 tonslie torn ligaments on crack face with a maximum 33' between ligaments.
-3. Total crack angle was 128' and included 5 tensite torn ligaments on crack face with an average of 21* and maximum 57' between ligaments.
- 4. Laboratory corrosion test conducted on crack prior to destructivo examination may have increased crack dapth and resulted in reduced remaining ligaments. Loak rato beforo
.- corrosion test was 20 cc/hr and was - 500 cc/hr after testing at 680*F for 2 months in primary water. No discernibio boundary between field and laboratory corrosion could be seen.
2 17
t 4 Table 2.2 l Summary of Deterministic Assessment for Potential Circumferential Cracks , f PWSCO h1WEXTEX ODSCO at TSP Edces Tube intecrity Conflderation Seamented Crack Anole (1) RPC AMie 4 50% Crach bec Projected EOC RPC Crack Angles
+ o 95% Probability ;
Crack Angles for Tube Burst o 3xaP N .O. ! o APsta Crack Angles for Vibration Propagation o WEXTEX Transition . Turbulence
+
- Fluidelastic .
o ODSCC atTSPs Turbulence Fluidelastle Potential for Accident Condition Propagation o SLB Flow o SSE Loads o LOCA + SSE Loads >
~
Notes:
- 1. Segmented crack model for-5/1 aspect ratio (0.25") through wall crack separated by ligaments sized (0.05") to remain elastic at normal operating pressures, j
- 2. TW = Through wall crack angle
- 3. Values given are RPC angle for ODSCC model and (in parentheses) acceptable through
. wall crack angle as 260% of RPC crack angle. A through wall crack angle of [ ]b,e . was the largest angle evaluated in vibration analyses. I 2-18 m, -ww , , - - , --..a e - - s w., e v. e..w---,, - ~ n,--, w---_ - _ - - , . . - - ~ - , - - , , , ,e-w.,
r 1
- Table 2.3 Summary of Probabilistic Assessment for Potential Circumferential Cracks PWSCC atWEXTEX ODSCC atTSP Tube Infecritv Consideration Transition E@cs EOC Crack Angle Projections bc o Largest RPC Crack Angle o Percent of TotalIndications Tube Burst Assessment: Through Wall Cracks o Crack Angle for Burst at 3AP N ,0.
Prob.of any tube exceeding 3AP o Crack Angle for Burst at APSLB Prob.of any tube exceeding APSLB . Potential for Tube Vibration Propagation o Crack Angle for Crack Propagation Turbulence: Thruwall cracks Fluidelastic: Thruwall cracks o Probability of exceeding angle for propagation
- Turbulence
.. Fluiddlastic Notes:
- 1. Estimated to be -0 as angle is much larger than found in any Monte Carlo samples,
- 2. Velves given are for central region with values in parentheses given for peripheral region.
~
- 3. Values given are for uniform through wall crack with values in parentheses given for through wall crack plus 50% deep crack over remaining tube circumference.
2 19
Table 2.4 ,j Circumferential and Axial Crack Tubo Burst Capability Summary
.1 Crack Morchotoav Eg g Circumferential Crack Anctes
, b.c Segmented Cracks Segmented Cracks with 50% Deep Crack Single Crack Single Crack with 50% Deep Crack Axial Crack Longths inch 4
_ _ b.c Segmented Cracks ! Single Cracks 2 20
_ .. _ _ ._ _ _ .~._ - _ ..___ _ - 3.0 PULLEDTUBE EXAMINATIONS 3.1 1991 Pulled Tube R11014 Tube R11 C14 from Steam Generator B of North Anna Unit 1 was cut below the second hot leg support plate and pulled for examination to characterize corrosion cracking suspected to exist at the first support plate crevice reglon. This summary provides l examination data that is considered significant. ' Nondestructive Examination : Visual examination of the pulled tube showed circumferential cracking at and just below the location of the first upport plate top edge. The crack networks existed at two locations, one centered ricar 45' and the other centered near 220'. (Zero degrees faces the steam generator divider plate and 90* is clockwise of O' when looking in the , primary flow or upward birection). Dimensional characterization showed a maximum ovality of 0.028 inch at this location with the apex of the ovality occurring near the centers of the crack networks. Dimensional characterization also showed that the dented support plate (denting was observed by field oddy current data prior to the tube pull)
~
acted as a die during the tube pull, reducing the tube diameter above the support plate approximately 0.008 inch. At the tubesheet top, the explosive expansion transition was approximately 0.75 inch long with a diametral expansion of 0.018 inch. The center of the expansion transition was located approximately 0.1 inch above the estima*ed 'ocation of the tubesheet top. X. Ray radiography clearly showed a complex network cf circumferential cracking between 0.0 and 0.2 inch below the support plate top edge. The cracking appeared to be almost continuous around the circumference. in addition, fair't radiographic indications of circumferential cracking were sporadically observed around the circumference at the support plate bottom edge location. The cracking at the bottom edge location appeared to be restricted to a more narrow band than the cracking at the top edge location. No crack indicat'ons were observed at the tubesheet top location. Laboratory eddy current testing was performed using bobbin and RFC probes. Dobbin probe data showed a 99% deep indication at the first support plate crevice region. APC examinations found two major circumferential cracks at the support plate top edge , 31 , t a_.,,,.._ -
location, the same as in the field examinations. The cracks were 160* and 140' long. The cracking at the top edge location had RPC signals that were approximately 30 volts. . In contrast, 5 to 6 volt readings were obtained in the field, indicating that the cracks were significantly opened by tho tube pull. At the support plate bottom edge location, . laboratory RPC examinations found 3 to 5 regions around the circumference with circumferential crack indications. The largest of these mdications was 80* long and had a signal strength of 3.7 volts. A review of the field RPC data showed that at least one, i posslbly two, of these regions were observed at the support plate bottom location. No ! field or laboratory eddy current indications were called at the tube &heut top location. ' l Ultrasonic testing (UT) in the laboratory showed circumferential indications near the i support plate top edge and smaller circumferential indications near the bottom edge in both the radial and axial aim scans. No axial Indications were detected in the circumferential aim scan. Similar to the RPC data, two major circumferential indications were found at the top edge location where they were almost cohtinuous around the circumference. At the bottom edge, four to five locations, more or less uniformly located around the circumference, had circumferential indications; the largest was 80' to 90* long, in addition, two small circumferential indications were observed at other ' locatiens. One was located within the center of the support plate crevice region near 40* and the other 0.7 loch above the support plate top edge near 140*. (Destructive examination showed that these latter indications were not caused by corrosion). A number of axial and circumferential indications were observed above, below and near the tubesheet top location where the expansion transition was located. Only one of these was judged likely to be a corrosion related indication. it was a 0.2 inch long axial indication located 1 inch above the tubesheet top near 50*. (Subsequent destructive examination thowed that this indication location did not have corrosion degradation), The remaining UT indications at the tubesheet top region were judged to represent tube pulling marks andor surface deposits. Destructive Examination ThE first support plate crevice region- was found to have two regions with circumferential cracking of OD origin. One was located at the support plate top edge and I. the other at the support plate bottom edge, No cracks were found within the support plate crevice region except for the cracking at the support plate edges. The more extenslV9 Corrosion occurred at the support plate top edge where the cracking was 32 l l'
located from the support plato top edge to 0.2 inch below. The crack network at tho
. support plate top edge formed a contihuous 360' macrocrack with two regions with deep (greater than 40% deep) penetrating microcracks. One of the two regions with dooper cracks was 124' long and was located from 348' to 112*, The other was 113' long and was located from 168' to 281*. Both regions had locations were the cracking was through wall. The maximum longth of through wall cracking was 00'. Tho two roglons were more or less diametrically opposito at d woro contored near tho apox of tho tubo ovality. Table 3.1 presents a summary of the crack depths at solocted locations by SEM fractography, Figure 31 shows a sketch of the crack orlontations.
The circumferential cracking at tho support plate bottom edge occurred in two separate microcrack notworks which formed two continuous macrocracks separated by regions without any ctacking. The cracking was located from the support plate bottom odge to O.15 inch above the bottom. One region with cracking was 135' long and was located from 315* through 360/0* to 90*, The second region was 112' long and was located from 135' to 247*. Again, the two regions were contored near the apex of the local eva!!!y. Table 3.1 presents a suminary of crack depths found at the support plate bottom edge. The crack depths for one of the two cracking regions was determined by SEM fractography. At the other region, where the cracking was judgod to be loss deep based l: on vbual examination of deformed tube sections, the crack depths were dolormined by metallography. Consequently, the depth of cracking is not as extensively charactorized in comparison to SEM fractographic observations. The maximum depth of cracking at the st.pport plate bottom edge was 98% through wall. l The morphology of the individual circumferential microcracks within the support plate crevice region varied from only IGSCC to IGSCC with significant IGA components. Where the' individual microcracks joined together, only intergranular features woro usually observed. Only in a few cases did dimple rupture features prodominate on tho lodges located between the numorous individual microcracks. At the tubeshoot top location, within the conter of the expansion transition, a f semi continuous- network of circumferential and axial. OD origin, microcracks was found. .The maximum longth of axial c.acking was 0.04 inch. The maximum length of continuous circumferential cracking was approximately 30*. The circumferential cracking was deeper than the axial cracktrg and the maximum depth of circumferenilal cracking was 21% through wall. Table 3.2 prosents a summary of crack depths found at 3-3
the tu'.,esheet top region. The depths at the tubosheet top region were dolormined by both SEM fractography (on the half of the tube circumference judged to have the deepest ,; cracks based on visual examination of deformed tube sections) and by metallography. ' Summarv and Condusions Intergranular OD origin, circumferential stress corrosion cracking was found at the top and bottom edges of the first support plate crevice region of tube R11014. The ; morphology of the cracking ranged from SCC to SCC with significant IGA components. The cracking at the top edge was the deeper and the more extensively located. The cracking ! was continuously located around the circumference. Two regions, el 124' and 113' lengths, had cracking that was greater than 40% deep. The maxirnu:n length of through
- wa!! cracking was 90'. The cracking at the bottom suppori plate edge was also significant and had a cimilar distribution. The regions with it e doopest cracks wore contered near the apex of the tube ovality. The tube within the suppcrt plate region was dented (0.028 inch).
At the tubosheet top region, minor circumferential and axial OD origin, Jr.iergranular corrosion was sporadically present near the contor of the expare;.:on fransition, which was 0.75 inch long with a diametrical expansion of 0.018 inch. The circumferential cracking, which was up to 21% through wall, was deeper and more extensively distributed than the axial cracking. i Eddy current data provided an accurate description of the degradation found at the support plate top odge. At the support plate bottom edge, oddy current data, especially , the field oddy current data, provided a less complete description of the degradation even though the cracking was deep. Ultrasonic testing in the laboratory provided slightly more complete description of the cracking, but it also generated indications not caused by corrosion degradation.
- 3.2 Prior Tube Exams from North Anna Unit 1 Steam Generators 1985 Tube Pull t
Tubes R3 C41 and R9 C58 were pulled from the hot leg side of Steam Generator C to l examine the cause of clear and strongly distorted oddy current indications as well as dent 34 ,
. - _ ,_ _ w _ __ _ . . - _ _ - _ _ _ . _ . . - .
signals at support plate crevice locations. Results of the examination showod SCC on both the ID and the OD surfaces in the immodlate vicinity of the tubo supports. No significant corrosion was found in the tubosheet areas, which included the tuboshoot top and tho tuboshoot crevice region of these WEXTEX expanded tubes. Dimonsional charactorization of the support plate regions indicated a small amount of permanont dolormation of tho tubing at suppori plate Intersection locations. The maximum amount of deformation ranged from 4% ovality to 7% ovality depending upon the technique used to measure the deformation. Tube R3 C41 was secMnod above tho third support plato region and tube R9 C58 was sectioned above the first support plate region. Results of the metallurgical examination Indicated PWSCC that was 100% through wall at Support Plate #2 and 50% through wall at Support Plate #1 in tube R3 C41 and 85% through wall at Support Plato #1 in tube R9 C58. The most extensive ID IGSCC involved two paralloi axial patterns of cracks extending from 0.8 inch above Support Plato #2 to 0.2 inch below the bottom of the support plate in tube R3 C41. An anglo of 180' separated the two axial networks, and the cracks were located at the minor diamotor of the ovalized tube whereas the circumferential cracks in tube R11 C14 were located at the taajor or largo diamotor of the ovalized tube. The OD IGSCC found was confined to within the tubo support plato
~
crovice region and ranged from 15 to 40% through wall. Laboratory tests of reverse U band specimens indicated a low resistance to PWSCC. The microstructure was typical of mill annoaled Alloy 600. The PWSCC at the tubo support plato regions is attributed to high stresses in the tubing at the tube support plate intersections produced by tubo denting caused by corrosion of the carbon stool support plate matorial. The oddy curront signals generated by donts at the tube support plate regions woro distorted in some cases. Based on the tube pull information, this distortion may have arisen from the influence of the cracks in the tubing at the same locations. 1987 Tube Putt Tubes R17 C31 and R18 C26 woro removed from the hot tog side of Steam Generator A with tube R17-C31 cut just below the second support plate and tube R18-C36 cut just
, below the first support plate. F!gures 3 2 through 3 5 show SEM fractography of the pulled tubos. At the tubosheet top location, both tubes h,ad circumforontial PWSCC in the explosivo expancion transitions. The macrocracks were composed of numerous parallel microcracks confined to a narrow zone. Some microcracks had grown together 3-5
by intergranular corrosion while others had not, The macrocrack in l';be R17+C31 was . 176' long and up to 90% through wall. The macrocrack in tube R18 036 was 128' long and was through wall for 53*. The lattnr crack grew partially during plant operation and partially during an accelerated laboratory corroslor test. Apparently, the . laboratory corrosion growth component was confined to ligament interconhection and i possibly to increasing the depth of the macrocrack. Burst testing of the tubesheet top regions of both tubes showed that the tensile strength of the tubes was not greatly [ affected by the circumferential crack networks. The strength properties were cloye to ! those for nondegraded tubing. Tube R17 C31 burst at 10,700 psi and tube Ri8 030 burst at 9,250 psl. It was suspected that ligament corrosion interconnection during the - laboratory corrosion testing, decreased the burst strength of tube R18 C36. The tubesheet top sludge pile regions of both tubes also developed OD origin intergranular corrosion that was up to 10% deep for tube R17 C31 and up to 6% deep for tube R18-C36. The first support plate crevice region of tube R17-C31 t ad both OD origin and ID origin , IGSCC. The crevice region also had a 0.001 to 0.002 inch radial dent. The CD origin cracking was up to 28% deep and confined to just below the suppoit plate top edge. Thn extensive network of primarily circumferentially orientated mic4ocracks formed a 90' long macrocrack. The morphology of the CD origin cracks was lGSCC with some IGA characteristics. Many ' similarities exist between this OD origin circumferential cracking at the first support plate crevice region of tube R17-C31 and the OD origin circumferential cracking found in the 1991 examination of tne first support plate crevice region of tube R11-C14. The PWSCC found withm the first support plate region of tube R17-C31 was up to 22% deep. It was located from the support plate top edge to almost the mid support plate-region. Near the mid support plate region, the PWSCC was axially orientated. Near the support plate top edge, the cracking was tilted 45' to the major tube axis. 3.3- Crack Morphology.for Tube Integrity Assessments The North Anna Unit 11985 to 1991 pulled tube examina: ion results provide a basis to - define crack morphologies for the tube integrity assessmcats. Both tne PWSCC axial and circumferential cracks are. characterized as cracks wi'h multiple initiation sites and numerous microcracks adding up to the macrocrack length wnich is typically identified 36 -
I by RPC inspections. The Individual microcracks have aspect ratios (length / depth) of j
- 4/1 in 6/1 and are separated by ligaments tending to be nearly through wall. Tho {
I doopest cracks are typically within a macrocrack or adjacer.1 microcracks for which the separating ligament may have been lost by corrosion. No sr .. cant cracking is found other than the essentially continuous macrocrack. For analysis, the PWSCC cracks can : be represented by a ligament model with 0.2 to 0.3 inch long, deep (assumed through l wall) segments separated by ligaments (assumed wall thickness). An elastic ligament ( model is used to $120 the width of the ligaments cuch that the ligaments romain clastic ! under normal operating pressuro differentials, which results in ligaments of about 0.050 inch width for 0.2 inch through wall sogments. . I The ODSCC degradation shows many multiple initiation sites with Individual crack depths of 8% to 33% found for tube R11014. Based on the limited pulled tube data base for ODSCC circumferential cracks, such as R11C14, crack progression through wall appears to link up the individual microcracks by carrosion of the ligahionts such that the through wall crack - shows few remaining uncorroded ligamonts. Thus the I circumferential ODSCC cracks can be modoled as deep or through wall cracks with few ligaments other than the remaining ID wall thickness and with the potential for some , partial depth cracking around the remaining tube circumference. For conservatism in l the present analysis, ODSCC is modeled as a through wall crack with and without a 50% { deep crack around the remaining tube circumference. Applying this model to an RPC-
- indicated crack angle ignores the remaining wall thickness expected over all or part of f
app c on of this odel to crack p opagati n st e , it s necess to mit he through wall crack angle to a fraction of the RPC implied crack angle. Based upon the f R11C14 results, the through wall crack angle can be modeled as 60% of the RPC angle. r i + j e i i t 37 !
. . - - .. - - . _ - - - -___--- -_-.- -. -w
3 i i I Table 3.1 Depth Profiles of OD Origin IGSCC in Tube R11C14 Hot leg First Tube Support Plate Region ,. OR!ENTATION 1ST TSPTOP ?DGE 1ST TSP BOTTOM EDGE (degrees) (%) (%) J 0 92 5 98 h 11 's100 80 .* b 22 s100 f$ B0 $I ' 33 100 $$ 66 $ 45 100 80 )jg$ 56 67 100 100
) $l gy g5 76 76 33 g5 78 100 E 72 g3 90 101 100 100
$$ 64 J 0
112 67 / 0 ! 1E3 8 L 3 135 12 146 12 53M E-157 12 61M j$ 168 71 3
- SBM '?
.?
180 92 f- - 191 s100 jb
*N r f"E 202 91 44M OE 2$
213 96 43M 225 hE m" 48M 8 s100f 236 100 0E $ ~8 247 89 2$ 8 ' 258 96 0 270 81 0 ~3 0 281 70 / 0 292 33 0 303 17 0 e . 315 12 40 e x 326 337 8 21 E e 4 72 78 5Ej EE 348
*e
- 42) h 3$ 80 J h 2 $
8
- Crack network focation defined for cracks greater than 40% through wall. The cracking was "
continuously present within tne circumferential band shown. All depths were determined by SEM fractography unless "M" follows the depth. In the later case, the depth was determined , by metallography. 30
Table 3.2 , o i Depth Profiles of OD Origin IGSCC in Tube R11014 Hot Leg Top of Tubesheet Region j i TUBESHEET EXPANSON ORTENTATION TRANSITlON ' (Degrees) 0 0 11 4M ' 22 3M 33 . 45 .E , , 56 E '!i 67 $ '@ h 7d &Rs l 4 90 j *, + 101 *D" 112 E 123 )Nh.e '~ 6
, 135 0 146 ;
157 , 168 j l 180 2 t 191 8 202 6 213 5 225 8 i 236 8 247 21 , 258 8 i 270 0 i 281 4 ; 292 6 303 2 ' 315 12 326 4 337 8 348 8
' Crack network location defined for cracks greater than 40% circumferential through wall.
The crack was continuous ly present within the circumferential band shown All depths were determined by SEM fractography unless "M" follows the depth, in the later case, the depth was determined by metallography. 39
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l 4.0 EDDY CURRENT DATA REVIEW 4.1 1991 Inspection Summary The steam generator tubing inspection was comprised of three major components:
- 1. Full length oddy current inspection of all tubes in the steam generator using conventional bobbin coil probes. This technique is used for general purposes:
identifying the location of all fixtures, diameter changes, sludge and freo length degradation that is axial or volumetric in character.
- 2. 8 x 1 pancake coli probes were used on the hot leg straight lengths for 100% of the tubes in each steam generator, as follows:
S/G A: 100% to 4H,20% to 7H* S/G B: 100% to 7H S/G C: 100% to 4H,20% to 7H
* #H: Support Plate # and H for Hot Leg Degradation masked by dents at support plates or WEXTEX transition at the secondary tubusheet face is rendered more visiblo by this technique and circumferential cracking detection is much onhanced over the bobbin coil.
- 3. Rotating Dancake Coil (RPC) probes were employed to inspect 100% of the hot leg WEXTEX transitions -- i.e. 3 inches above and 3 inches below the top of the tubesheet, Row 2 U bend regions, all indications reported to be greater than or equal to 40% TWD by bobbin coil, and all possible indications (PI) by the 8x1 coll. This rotational technique largely eliminates the lift.off inter-forence signals associated with dents and the WEXTEX transitions, provides good detection for all types of degradation and serves to characterize the signals reported as pcssible degradation by 8 x 1 and bobbin probes.
4 1 1 __-__-__- n
4.2 WEXTEX Expansion Region 4.2.1 RPC Probe Characteristics The characterization of the hot leg WEXTEX expansion inspection results is based entirely on the results from the RPC probe for the reasons stated in 4.1. The circular symmetry of the EC field about the coil and the helical scan which results from axial translation of the rotating sensor enabler .'ffective detection capability for both axial ind circumferential cracks. The probe is ocployed at a rotational speed of 300 rpm and translated at a rate which yields a pitch of approximately 0.050 inches between scan lines. The coil diameters for the Echoram 2XRPC are 0.105" for each of the identical coils. The output of the two diametrically positioned coils is digitally combined by software to emulate the output of a single coil moving at half the translational velocity. 4.2.2 Inspection Results for WEXTEX Region As stated above, all hot leg tubes were inspected 13 inches about the top of the tubesheet, a length which includes the WEXTEX transition as well as the sludge zone just above the tubesheet. The RPC inspection zone also includes some expanded tube region below the top of the tubesheet (TTS), and the transition region as well as nominal diameter length above the tubesheet. The RPC findings in the expanded zone and the transition cannot usually be corroborated by the bobbin coll, but some correspondence between boboin and RPC data can be expected in the straight length above the tubesheet. For the purpose of this section, bobbin findings are excluded except if useful to assist in characterizing the orientation of a crack, since bobbin probes are optimized to detect axial cracks. Table 41 provides a summary of the WEXTEX region RPC results. Both axial and circumferential cracks are reported; crack origin is determined on the basis of apparent phase angle associated with the lissajous figure for the scan with maximum signal amplitude. Calibration of the phase angle dependence for this purpose depends upon setting the 100% TWD EDM notch standard response at 20* phase angle; flaw like signals with phase angles less than 20* wilt tend to have ID origins, while those whose phase angles exceed 20* tend to reflect OD origins. However, since the influence of transition non conformity cannot be excluded. Lroader phase spread is usually allowed 4-2 j
_ __ . _ _ . _ . - . _ _ _ _ - . _ . _ . _ - _ _ . _ ~ . _ _ for the ID origin signals. Thus signals with phase angles loss than 30' have been
, classified as PWSCC (lD origin SCC) whilo thoso greater than 50' are classified as ODSCC; those between 30' and 50' are regarded as indeterminate. (As ID signals
. psnetration depth is essentially 100%; as OD, TWD exceeds 80%)
. The total number of tubes with W'IXTEX transition cracking was 221. Of these, 200 are attributed to PWSCC all circumferential (three tubes, one in each S/G, woro originally reported as axial cracks, ID in origin: review of the dimensional data showed that the reported azimuthal extent for each one was greater than the axial longth for the purpose of this evaluation, all three have been transferred to the circumferential - category); 6 due to ODSCC I circumferential and 5 axial, and the remaining $5 are indeterminate circumferential indications. Five of these tubes alsc amibited > circumforential cracking at support plates which was given primacy in the hierarchy of tube plugging attribution: hence, the 216 tubes attributed to WEXTEX ti asition cracking in Table 41. 4.2.3 Ligament Sizes Between Circumferential Cracks
. I A baction of the WCXTEX transition circumferential cracks were reported as multiple circumferential indications (MCl): S/G A had 3 MCl's. S/G B 6, and 28 in S/G C. All but 1 of the MCI's reoorted weto centered below the top of the tubesheet; R25C43 in S/G C, with arc lengths of 78' and 92* separated by 65' and 125' ligaments is reported between 0.00" and 0.05" above the tubosheet, in several cases, circumferential cracking is reported at two axially displaced locations, or more properly in a band extending 0.07" axlally at maximum; this is equivalent to no more than two scan lines separating the crack centers. This may represent axial separation of the cracks or slight axial meandoring of the cracks, but in some cases it can be explained on the basis of scatter in the measurement of the axial conter of the cracks.
Where ligaments have been reported between circumferential cracks, this represents a situation in which the EC analysts have observed a full return to the null amplitude level between identilled indications. Values as low as 7* arc length and as high as 96* are reported for minimum ligaments between circumferential indications. Figure 41 displays the distribution of minimum ugament are lengths measured for the MCI's on a 43
E composite basis for the three steam generators. Given the 2.43" inner diameter represented by the 360* circumference scanned with each RPC rotation,7* arc length 4 is equivalent to 0.047". For EDM notches separated by ligaments in a circumferential
~
L plane, clear resolution is not seen until the ligament exceeds 0.2" (30*); Figure 4 2 illustrates the bd.iavior obtained with EDM notches. Sr.. aller reparations reported by the fir.N analysts for circumferential crach may be the result of dispa%s in (elative
, amplitudes for adjacent indications. In ar y case where the presence of a ligament is k confirmed, it is believed that the actual separation exceeds the apparent value since no corrections are made for field spreads at the crack tips. Additionally, EDM notches represent abrupt penstrations, while the through wall portion of a crack may be - limited to very short lengths. Typical examples of the more extended crack networks A ; are listed in Table 4-2 and illustretcd on Figures 4 3 through 4 8 Table 4-3 lists the indications and ligament arc lengths fcr all MCI's reportcd.
, 4.2.4 Distribution of Circumferentiai Crack Sizes [s [ The azimuthal arc lengths of the predominantly ID-originated circumferenilal cracks , F found in the WEXTEX transition zone are summari:ed graphically in Figure 4-9. The 1 histogram dicplays the crack arc lengths reported as single indications as well as the individual indications comprising the multiple indications; this provides the natural crack length distribution. The values indicated along tae abcissa are the upper values
~
for the bin; i.e., 300* includes all data from 270' to 300*, inclusive. R22C66 cf S/G s C was reported to have a sin 0le crac" -tn 7 arc length; review of the field data in the laboratory shows Inat the indicae n k 4ctually comprised of two smaller cracks, un one 92* and the other 168*, separated by a 15* ligament. R19C33 was reported to m have a single crack with 19* arc length at an elevation 0.3 inches below the top of the tubesheet. Althougn this tube was plugged for an MCI at an elevation of -0.1 inches, a laboratory review of this location did not confirm the presence of an indication
- 11 is also of interest to examine the combined lengths for multiple indications; this can be done in two ways. First the crack network arc length is presented in Figure 410; this dimension represents the ar.imuthal extent of cracking observed from beginning of the first indication to the end of the last indicaron excluding the largest ligament.
Alternatively Figure 4-11 represents the summation of individual ind: cation arc lengths excluding all ligaments; this gives the actual arc length of cracking, for companson with the total circumference. 4-4
I k 4.2.5 Distribution of Cracks Relative to the Top of the Tubesheet f l All but one of the indications of circumferential cracks were reposed at or below the i top of the tubesheet. Figure 412 presents a graphical summary of the distribution of [ these locations re'lative to the top of the tubesheet. The axla! locus of the circumferential cracks is above the bottom of the expansion transition, averaging 0.07 : I inch below the top of the tubesheet. The WEXTEX transitions typically extend approximately 0.65 inches, from approximately 0.19" below the top of the tubesheet (Figure 413) endir'g an average 0.46" above the top of the tubesheet. Thus the 3 circumferential cracks occur mainly from the inside diameter at an average 0.12 inches above the last contact point in the tubesheet. { [ 4.2.6 Position of WEXTEX Cracks Relative to Tube Bundle Center Line . I i: WEKiEX COls are generally located within the sludge deposition zone of the j 9
- vet. Figure 4-14 shows a composite tubesheet map of the WEXTEX indications. !
#h' all indications are contained within an arc radius of about 45" from the tube !
'soi : enter line. Figure 8-5 (Section 8) shows the low velocity zone at the tubesheet
..- I eave.n. The WEXTEX indications generally fat' within a zone of cross flow less than 3 ft/sec. One tube these 200 indications,- R4C89, is outside of this zone. This is a i i
single CO; with a ( ick angle of 115*. Similar grouping confinement of WEXTEX indications is also observed in the Series 51 steam generators at other plants. P 4.2.7 Combined Axial and Circumferential Cracking in WEXTEX Transitions l t The RPC testing results for the WEXTEX transition summarized in Table 4-1, indicate ! the presence of 4 tubas o.i which axial cracks were observed. The RPC isometrics for i each of these tubes was reviewed to determine whether any of them exhibited both axial j and circumferential cracks. None of these tubes had both orientations of cracking in the i WEXTEX transition. Thus thare are no occurrences of mi-ed mode cracking l (intersecting axial and circumferential cracks). ! i 4.3 Axial Cra:ks at Tube Cupport Plates f The existence of axially-oriented ID-initiated cracks in the dented . support plate ! t intersections on the hot leg portion of the tubes has been known since 1984. The . 4-5 ! [
?
denting responsible for this degradation occurred during the inillal cycle of operation at North Anna Unit 1; it is believed to have resulted from the inadvertent intrusion of lon exchange resin particles into the steam generators. The progression of the denting has been controlled if not completely arrested by the application of boric acid treatment as part of the secondary system chemistry control program beginning with Cycle 2 operations. Notwithstanding, the boric acid effect on denting, tube leakage attributed to ID cracking (axial PWSCC) in the dented support plate intersection; was observed during 1983 and during subsequent periods of operation. Inspection of the tubes with bobbin probes is effective to identify dented intersections with cracks extending beyond the support plate. Tubes which exhibit this behavior are then re examined with RPC probes to characterize the indication found by the bobbin probe. Additionally some of the indications detected with the 8 x 1 pancake probe may result from axial cracks, more likely from multiple axial indications (mal) than from single axial indications (sal). Usually axial indications found only by the 8 x 1 are confined within the dented TSP intersection, thus not visible to the bobbin probe. Each possible' indication (PI), whether axial or circumferential in orleatation, is , re examined with the RPC probe and so defined as axlal or circumferential. A summary of all indications found at TSP elevations is presented in Table 4 4, 4.3.1 Indications Outside the Tube Support Plate As can be seen from Table 4-4, 321 axial Indications (sal & MAI) were reported at the support plates; in each of these cases there is denting coincident with the TSP. Examples of a typical sal and a typical MAI are given in Figures 4-15 and 416. The portion of the cracks which extend beyond the support plate edge more properly beyond the dent, are visible to the bobbin probes. Typically cracks observed beyond the dent are 0.2 inches or more in length and tubes on which they are detected are plugged. 90,81, and 107 axial TSP indications were reported in S/G's A B and C respectively. 4.3.2 Axial indications inside the Tube Support Plates The tubes identified with axial cracking by RPC and which were not detected by the - bobbin coil, were detected by the 8 x 1 probe. For these cases, cracks are confined to within the dented TSP intersection. A relatively small fraction of the indications of axial cracking at the TSP's fell into tais category: 16 in S/G A.15 in S/G B and 12 in 4-6
i r l S/G C, a total of-43, representing only 13% of 'he total. Figure 4-17 presents an [
- ,- example of axial cracking characterized by RPC probing from which the confinement of -f the cracks within the dented TSP intersection can be appreciated.
l 4.3.3 Crack Length Distribution Outside Support Plates i Since cracks which extend beyond the confinement of the dented TSP Intersection may ! cause tube leakage, such tubes are plugged. Furthermore, growth of such cracks outside i
- the TSP which may go undetected due to the influence of the dents.must be understood so t
that unacceptable crack lengths will not develop during the ensuing operating cycle, , A histogram of the observed crack length derived from the bobbin probe is given in , Figure 4-18; from this it can be inferred that reliable crack detection extends 50.2 inch beyond the edge of the TSP. The average crack length observed 0.49 inches. ! suggests growth from previously undetected lengths (<0.2") of approximately 0.3 [ inches or 0.25" per year. Figure 419 compares the crack distributions from the [ 1991 outage with prior data from 1987 and 1989. It is seen that the largest indication ; lengths have been reduced since 1987 with similar overall distributions of lengths ; from all three outages. The number of indications in 1991 is comparable to 1987 and 7 higher than 1989. , t Figure 4 20 shows the distubution of indications from the three outages as a function [ of TSP elevation. The strong decrease with elevation and associated piimary temperature reduction can be seen in the figure. Very few indicatioris have been found I at the top (7th) TSP, l 4.4 Circumferential Cracks at Tube Support Plates ! Prior inspections with 8 x 1 probes had discovered degradation at dented TSP
- intersections which were thought to be " volumetric" in character. To understand the l
extent to which such degradation existed, the extensive 8 x 1 probe testing described in f
^
Section 4.1 was instituted. A total of 110 PI's identified by the 8 x 1 probe were '{ subsequently characterized by RPC testing as exhibiting circumferentially oriented degradation, including 62 col's (circumferentially oriented indications) thought to . t represent cracks and 48 indications regarded as " volumetric". Comparison of the RPC ! traces for the two groups of indications led to the conclusion that all should be regarded 4-7 i
+
l as representing circumferential cracking. Figures 4-21 and 4-22 provide typical examples from the two categories. To characterize the origin and nature of the cracking, a tube from S/G B (R11C14) with a COI at the #1 TSP was pulled for laboratory evaluation. Metallographic examination confirmed that the circumferential Indications resulted from OD initiated intergranular stress corrosion cracking at the support plate edges (See Section 3). An additional 433 TSP intersections were inspected with the RPC probe and no additional ciscumferential indications were found. This resi'll supports tne 8x1 detectability at TSP locations. 4.4.1 TSP Circumferential Crack Size Distribution The azimuthal are lengths associated with the TSP circumferential crack indications are summarized graphically in Figure 4-23 for the composite of the indications from all three S/G's. Only the crack angles associated with COls (an not MCis) are shown, since MCis are addressed separately, as discussed further in Section 5.0. The largest COI arc length reported is 212*, which is larger than any of the individual cracks seen in MCis. Ninety-five percent (95%) of the COI crack angles are less than 150*. . Occurrence of circumferentially oriented cracks in dented TSP's is a hot leg phenomenon which appears to be limited to the first five support plates. Figure 4 24 shows the axial distribution of the circumferential cracks as a function of plate number. It can be seen that more than 90% of the circumferential cracks reported were found on the first two support plate elevations. A fraction of these cracks were reported at the lower edge of the dented TSP, but 280% were reported at the upper edge of the plate. Figure 4 25 illustrates the composite locations of tubes which were found with circumferentially-oriented cracking at the support plates. It is seen that the TSP indications are essentially randomly located within the tube bundle. The nine tubes which had circumferential indications at the lower edge of the #1 TSP and their crack angles are as follows: 4-8
&GA &GB $GC Tube ID-Arc Tube ID-Arc Tube ID Arc R33028 112' R44CS2 118' R10C25 108' WEXTEX also R3C33 72' R9C47 129' (83') 94*
R23C35 - 97* R16C76 75' R31C53 - 84* R25C54 - 87* Only R10C25 in S/G C also exhibits degradation in the WEXTEX region at the top of the tubesheet; that degradation was reported in the field as multiple axial indications (MAI), distributed over a 126' azimuthal arc length. Additionally, this tube (R10C25) also is reported to have circumferential cracking at both the upper and lower edges of the #1 TSP. R9C47 exhibits multiple circumferential cracking at #1 TSP, v.ith reported arc lengths of 129' and 94* separated by an 83* ligament. 4.4.2 Combined Axial and Circumferential Cracking at Tube Support Plate Edges Because significant numbers of tubes had been reported with axial cracks at TSP's, those determined to have circumferential cracks at a TSP edge were reviewed for coincident occurrence. Table 4-5 lists the tube locations, crack angles, and separations between the axial and circumferential cracks found at TSP edges. Figures 4 26 to 30 show the RPC isometrics for the tubes listed. Table 4-5 shows that the separation angle between crack centers for circumferential cracks at the edges and the axial cracks extending outside the TSP is in the range of 80* - 120*. The pulled tube which confirmed the presence of circumferential cracking, R11C14, exhibited OD cracks whose centers were separated by 180* and located at the major diameter of the ovalized tube. A tube pulled in 1987 from North Anna #1, R9C58; exhibited axial ID cracks extending outside the TSP; its crack centers were found on the minor diameter of the ovalized tube. 4-9
Based on the RPC inspection data as well as the tube pull results, it appears that the locations of axial and circumferential cracking at the same elevation can be expected to be separated such that mixed modo cracking will not occur. Most of the circumferential cracks present in 1991 were shown by review of the previous inspection results (1989) with 8 x 1 probes to have been present at that time. Thus future circumferential cracks at the TSP's can be expected to be smaller than those found in 1991; it is unlikely that they will extend the 180 necessary to intersect with an axial indication. Figures 4-31 to 33 show the EC profilometry data for four of the tubes listed in Table 4-6, whose RPC isometric data are given in Figures 4-26, 27, 29 and 30. These tubes exhibit significant ovall:ation due to design, with differences between maximum and minimum diameters of about 30 mils. This association of the tubes with both circumferential and axial cracks outside the edges of the TSP's to highly ovalized tubes further supports the expectation that the axial and circumferential crack locations will be azimuthally separated. In summary, no intersecting mixed modes cracks have been found in the Ncrth Anna , Unit 1 S/G's to date at TSP's, and the a:Imuthat separation of the axial and circumferential cracks can be anticipated in the future, as supported by both the RPC , and profilometry data. Consequently there is no need to cons; der the coincidence of axial and circumferential cracks at common elevations in the structural evaluation in this report. 4.4.3 Circumferential Crack Depth Variation The characterization of Pls (possible indications) reported from 8x1 probe testing was accomplished using the RPC probe. Those Pls which were identified as COI (circumferentially oriented indications) or reported with percent depth calls exhibited the behavior which was assessed as circumferential cracking. Examination of the pulled tube (R11C14 from S/G B) determined that the observed degradation was in fact ODSCC circumferentially oriented, at the upper and lower edges of the dented tube support plate (#1). The RPC isometric plot for R11C14 1H describes an arc of 158' in azimuthal extent for which the maximum amplitude reported was 5.88 volts. Review of the metallographic data shows that 60% of the tube arc corresponding to this RPC trace had 4-10
1 nearly 100% through wall penetration by ODSCC. The 60% arc length transposed on
. the RPC isometric trace subtends that portico of the circumferential signal which exceeds the width of the peak at half the maximum amplitude. On this basis the portion of the observed cracks which should be assumed to be at maximum penetration can be Identified by measuring the width at half max. amplitude for those cracks whose reported are lengths exceed 120* to determine the potential for leakage and propagation.
This technique approximates the angular extent of maximum depth. It is again noted that the maximum depth may not be through wall. The ratio of maximum depth to total arc length provides guidance on modeling the cracks for structural integrity such as the potential for crack propagation due to vibration. The 120' arc length selected for review represents a lower bound through wall arc length for crack propagation considerations. The RPC arc lengths were reviewed for 95 tubes with circumferential cracks at the TSPs. Of these, 33 tubes had half-max, arc lengths exceeding 60% of the total arc length. These rescits indicate that maximum depths are typically less than 56%, with a standard deviation cf 10%, of the RPC arc length. 4.5 Detection Sensitivity for Circumferential Cracking 4.5.1 Detection Thresholds As discussed in 4.1 bobbin probes are used to detect axial and volumetric discontinuities. Thus though more severe circumferential cracks might be detectable with the bobbin, the presence of the transition signal and the dent signal serve to mask even the larger circumferential crack responses. All practical detection of circumferential cracks must be accomplished by NDE techniques which suppress or are insensitive to the effects of these disturbances. In the North Anna Unit 1 situation detection of circumferential cracks was accomplished using surface riding EC probes, such as the 8 x 1 and the RPC. These probes are optimized for omni directional crack detection because of the circular symmetry of the EC fields they induce. Further, the surface-riding aspect of their deployment causes the lift-off effect produced at transitions and detns to be reduced to manageable levels. 4-11
Circumferential cracks had been detected since 1987 in the WEXTEX transition. Prior inspections were conducted with the 8 x 1 probe for detection / screening with the RPC *; used to characterize the naure of any indications found by 8 x 1. Though a high rate of false calls was experienced in the earlier inspections,78 tubes had been plugged priot to Feb.1991 as a result of circumferential cracking. Experience at another plant in 1989 indicated that not all significant cracks detectable by RPC had been found with the 8 x 1 probe. Therefore all hot leg transition zones in North Anna Unit 1 were tested by Doth 8 x 1 and RPC probes. Table 4 7 provides a capsule description of the capabilities of the pancake coils used as sensors on the 8 x 1 and RPG probes. Based on a combination .of design features, laboratory evaluations and field experience, the thresholds for detection given in Table 4 7 are regarded as conservative in bracketing
- their expected performance. Figure 4-34 illustrates the geometrical aspects of the 8x1 coil and Table 4 8 relates the thresholds for detection with these probes to the technical bases which justify their use.
4.5.2 Circumferential Crack Detection Summary WEXTEX transition zone circumferential cracking at depths greater than 50% throughwall, with arc lengtns in excess of 50* were successfully detected and characterized with RPC probes. Shorter length cracks with deeper penetrations can also be detec;ed. Circumferential cracking at the edges of dented TSP's were successfully detected using the 8 x 1 probes for crack depths averaging 50% or more over arc lengths exceeding 75*. Again, deep cracks of much shorter azimuthal extent were also detected.
-4.5.3 - Comparison of Crack Detection between RPC and UT Tests (Other Plant)
The presence of !igaments in PWSCC circumferential cracks, as found in the North Anna pulled tubes, has also been demonstrated by UT examination of a large WEXTEX indication in another plant. Figure 4-36 shows an RPC trace showing a 322* crack angle. The RPC interpretation is a single crack since the amplitude does not return to the null point between amplitude peaks. Figure 4-37 shows the UT results gated to . emphasize ID crack discrimination. The result confirms the RPC angular extent and also shows the strongly-segmented crack morphology with relatively few high amplitude (dark red color in figure) locations. Figure 4-38 shows the OD gated UT 4-12 i l l
. .. . . - _ _- ._ . _ _ = _ . _ - _ . _ . . . . . _ . . - _ . . - _ . . . ._ . . _ . - . . _ . .
i.
)
signal which again shows the segmented crack, e large ligament (20 30') between the ! two significant cracks and an overall smaller crack angle, Further UT demonstration of .l the segmented crack morphology is shown in Figure 4 39 for a crack identified by RPC ! I as 255'. l i I 9 f 9 i i s t F 6 [ t t L h i h l
.i t
{ . i
?
9 i 4-13 T k
Table 4-1 .i North Anna Unit 1 *: RPC Inspection Results for WEXTEX Transition Zones February,1990 SLGA S/G B SLG..Q Inla! Total Tubes Plugged
- 52 56 107 216 for Degradation in WEXTEX Inspection Data:
Axial Cracks ODSCC 0 3 2 5 PWSCC. 0 0 0 0 4 Circumferential Cracks
-ODSCC 0 0 1 1 PWSCC 45 52 103 200 Indeterminate I 1 1 1!i Total Tubes with 52 56 113 221 WEXTEX Degradation Tube Plugging Attribution /Hierarachy
- 1. - Support Plate Circumferential/ Volumetric
- 2. WEXTEX, Axial or Circumferential .
3, Support Plate Axial -4. Above Tubesheet 4-14
. Table 4 2 - North Anna Unit 1 February 1991 Example of Multiple Crack Networks (Figures 4 3 to 4-8)
RPC Pictures E'evation gi U 25/51: C .11 81' 33'
.04 147' 93* Reported 321' angle
.06 235* 9'
? 86* 13*
17/33: C .05 68* 11'
.07 124' 26'
.09 94* 46' 14/61: . C .08 234* 30'
. 07 60' 36'
.13 108' 72* Reported 294* angle
.10 43" 40*
.09 58' 69*
10/76: C .09 137' 36'
.12 132' 56' 23/60: C .04 227'
.03 62* 24* Reported 262* angle
. .01 200* 74* 22/49: B .01 158' 18*
.04 188* 17* Reported 346* angle 4 15
Table 4-3 - North Anna Unit 1 February 1991 Multiple Circumferential Indications in WEXTEX Transitions n- W f12 Uc. 2 fu Steam Generator A R18C29 109' 35' 63' R19C34 71* 24' 82* R21C37 102' 9' 55' Steam Generator B R10C19 84' 24' 84* R20040 96' 11' 84' R22C4 158' 18' 188' R14C53 56* 54' 95' R9C69 86' 51' 75' RBC70 98* 8' 108* Steam Generater C R13C26 49' 7' 189' R19C26 96* 80* 48' R19C28 89' 82' 67 R22C28 68* 48' 54' R20C29 123* 69' 89' R20C30 120* 22' 33' R24C30 85' 96* 60' R22C32 96* 22' ' 87' R17C33-- 68' 11* 124' 26' 94' , R19C37 55' 9 133' On - Circumferential crack are length (degrees) Ug. - Ligament separating multiple cracks (degrees) 4 16
\
i I Table 4 3 (continued) North Anna Unit 1 - February 1991 ; Multiple Circumferential Indications in WEXTEX Transitions 5 f!1 L!L 1 22 L!1.2 id Steam Generator C .I R17C39 72* 28' 118' R19C40 135* 19' 97* R24C41 140* 25' 71' . R25C43 78* 55' 92* R24C44 90' 20' 60' R24C45 82* 14' 80* R25C51 81* 33' 147' Cracks at two 235' 9' 86* separable elevations R25C55 120' 72* 72* j R17C56 71* 80* 83' , R14C59 115' 38' 66* - R16C59 140* 61* 68' i R25C59 87* 72* 65^ R23C60 62* 24* 200* R14C61 234' 24* 60* (see R25C51 note) 43' 40* 58* 72* 108* . R22C66 92* 15* 168* Lab re. analysis R23C66 105* 55' 90' R10C76 137' 36* 132' . 1 R10C77 182* 22' 65*
?
' ? On - Circumferential crack arc length (degrees) Lig. - Ligament separating multiple cracks (degrees) I 4 17 ,
Table 4-4 North Anna Unit 1 February 1991 Summary of Tube Support Plate Indications Steam Generator 8 B. C .I21HI Axlal Within Support Plate 16 15 12 43 (Found by 8 x 1) Beyond Support Plate 90 81 107 278 (Found by Bobbin) Circumferential (Found by 8 x 1) , COI by RPC 18 18 26 62 Percent depth calls by RPC 13 3 32 48 Tubes Plucced for TSP Ind; cations: Axial 73 79 92 214 Outside TSP (bobbin calls) Within TSP (8x1 calls)
~
Circumferential 29 16 47 92 Total 101 95 139 336 , 4-18
Table 4-5 TUBES WITH AXIAL AND CIRCUHFERENTIAL CRACKS AT EDGE OF TSP ANGLE BETWEEN LIGAMENT ANGLE TUBE CIRC. CRACK ANGLh___ {IRC. & AXIAL CRACKS EE]MEEN E CIRC. & AXIAL CRACKJ
$/_(I [0 CATION 1090 530 A R6C54(5H) 840 '
1300 600 R10C59(1H) 1160 1170 590 B RSC24(IH)~ 1170 970 740(1) 950 120(4) 350 B20040(1H) . 820 270 C R2C37(1H) 920 1 NOTES;
- 1. ONE OF TWO AXIAL CRACKS DOMINANTLY JNSIDE TSP.
4-19 f L _ _ __ _
l l l Table 4-6 DETECTION THRESHOLD 8X1 PR0gg e COIL DIAMETER OF 0.3" SUBTENDS ~ 450 ARC COILS ESSENTIALLY FILL ID OF TURE o HINIMUM DETECTION THRESHOLD EXPECT DETECTION IF CRACK > 0.5 COIL DIA. OR 22-350 CRACK e ESTIMATED 100% DETECTION PRO 8 ABILITY
- CRACK > 1.5 COIL DIAltETERS OR 67-800 CRACK e APPLIED DETECTABILITY FOR CIRCUHFERENTIAL CRACKS
- 0% 9 350, 1001 0 750 RPC PROBE
~
e COIL DIAMETER OF 0.105" SUETENDs ~ 150 ARC e MAXIMUM PITCH = 0.050" OR ABOUT HALF THE COIL DIAMETER e PULLED TUBE DATA INDICATES DETECTABILITY OF 230 THROUGHWALL CRACK e SMALLEST REPORTED RPC CRACK ANGLE a 500
- SMALLER ANGLES CAN BE REPORTED AS AXIAL INDICATION e APPLIED DETECTABILITY FOR CIRCUMFERENTIAL CRACKS
- 0% < 230, 50% AT 230, 100% AT 500 4-20
.- o .. .
a i. Table 4 DETECTION THaESHOLD y BASIS 4 8 x 1 PaosE O TumouGHWALL CaACK ANGLE ~450 COIL SPACING AND CONSISTENCY WITH DETECTAsILITY OF 0% AT 350 8 x 1 To RPC ANGLE ConRELATION AND 100% AT.750 APPLIED TO ODSCC AT TSPs o DEPTH TuaESHOLD ~50% Fon'LAaGE PULLED TUBE R11C14 COMPARISON OF ME - ANGLES AND ACTUAL APPLIED To ODSCC AT TSPs RPC PaosE o TumouGHWALL CaACK ANGLE ~230 WOG WEXTEX'SUWGROUP STUDY FOR PULLED i i DETECTABILITY OF 50% AT 230 TusE. EXPERIENCE AND 100% AT 500 APPLIED TO PWSCC AT WEXTEX l-4-21
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Figure 411 , NORTH ANNAUNIT1 WEXTEXREGION REPORTED CIRCUMFERENTIAL ECI's 7% . . ...........O* + * (/) Z COMBINED LENGTH tiXCLUDING UGAMENTS O co- . . . . . . . . . . . . . . . . ... . . m b. g . . . . . . . . . . . .. . . . . . . . . LL O 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ... . . . . . . A U) m O .. .. ...o... .. . . 33 . . . . . . . . . . . . . . . . . . . . . . .. , D Z io .. . . . . . ...........P.... .. .. .. . . . . . . . . I I I O O . _, 00 120 , , 100 ,
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- NORTH ANNA UNIT 1 i
SUPPORT Pl ATE DEGRADATION PROGRESSION 250 I ' g Note: 1987 and 1989 data irviaded Di's, PCT and IM1 z 200- M ^ h cads. The INR gutm d the total is less inn -- O iss F QO 150- - - - i
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Figure 4-21 39 CH 3 V CH 7 V 5.11 499 Khz CH 1 50 a 4 I > ROW COL
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~ - - -
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Figure 4m
.R10C59 - 1 HL - A S G TUBING PROFILE 790 _g
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...........................~. Dist 3
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Figure 4-33 RCJ 'O COL 40 ElXL1HIM DL1 METERS 795 - - - - - - - - - - v.
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Figure 4-34 R2C37 - 1 HL - C S G TUBING PROFILE 785 -m-gg 88 JZI El DI<L 1 7gg. ....................-......-........-..........gfm p M; 'tD . . . . . . . . . . . . . . . . . . . . . . . . . . _ . - ,
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-1.9-1.7-1.5-1.3-1.1-0.9-0.7-0.5-0.3-0.10.1 0.30.50.7 0.9 1.1 1.3 1.5 1.71.9 1.OCitT10N FWJM CENTER OF TSP (IN.)
6/ 000. orc. CFcumterenalaf exfent surveyed by one cos (45 dog.) Guaranteed one col response.
/
/
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5.0 PROJECTED EOC CIRCUMFERENTIAL CRACK DISTRIBUTIONS This section develops statistical projections of the WEXTEX and TSP circum! re 3 crack distributions at the end of the current operating cycie in late 1992. The cre.ck angle distributions are used in later sections to assess the limiting crack angles and associated probability of occurrence for tube integrity evaluations. The crack angle distributions are developed using: EOC angle detection threshold + growth + EC uncertainty. This definition is consistent with R.G.1.121 guidelines. However, when the orry circumferential cracks lett in service are those below the detection threshold, the eddy current uncertainty is not required to estimate the EOC values. Inclusion of Inc uncertainty adds conservatism to the EOC projections, particularly for 8x1 probes for which the uncertainty is more significant thar' br RPC probes. The 8x1 uncertainty is applied for both TSP and WEXTEX projections since 8x1 growth rate data are used for the projections. The data utilized to develop the EOC distributions include: Detection threshold for through wall crack angle (Section 5.1) 8x1 probe for TSP indications RPC probe for WEXTEX Crack Growth Rate Distribution (Section 5.2) Developed from 8x1 data between '89 and '91 inspections EC Uncertainty Distribution (Section 5.3) Based on 8x1 uncertainty developed from cr.uparism with RPC crack angles Number of EOC Indications (Section 5.4) Projected from historical indications using tog normal projections The above distributions are combinod using Monte Carlo sampling analyses to determine the projected EOC circumferential crack distributions. The 1991 inspection results used for the crack growth rate and EC uncertainty distributions are based on single circumferential indications. The single indications have larger maximum crack angles than any one crack of multiple circumferential indications at the same location. The largest single crack tends to be most limiting against structural criteria since the 51
rnultiple cracks are shorter cracks separated by RPC detectable ligaments of about 35' o; larger (assuming through wall cracks) as discussed in Section 4.2. In addition, the ; multiple circumferential cracks in the WEXTEX transition are dominantly located below the top of the tubesheet where the tubesheet constraint prevents tube burst.
- To support the adequacy of the projection data base and methods, the 1989 inspection results have been used to project the 1991 crack angles which can be compared with the inspection results. This comparison of projections and inspection results is given in Section 5.5. The established methods are then used to project the 1992 EOC crack angle distributions for the WEXTEX transition and TSP locations in Sections 5.6 io 5.8. The distribution of the circumferential cracks by S/G and TSP is given in Section 5.9. A summary of the projection results is givm in Section 5.10.
5.1 Detection Thresholds The detections thresholds for 8x1 and bobbin coil probes were discusset in Section 4.5. Based on this review, the detection thresholds applied for the EOC projec!ons are: RPC: 50% detection at 23' crack angle (assumed 0% < 23*) 100% detection at 50* crack angle 8x1: 0% detection at 35* crack angle
'00% detection at 75* crack angle in the 1989 inspection, the 8x1 probe was used for detecting the pre. - < :
circumferential indications at both the WEXTEX and TSP locations. Thus ", N detection thresholds are used to predict the 1991 in.,pection results. For the 1991 inspection, the RPC probe was used to inspect 100% of the hot leg WEXTEX transition
- and the 8x1 probe was used at TSP locations. Thus the applicable detection thresholds for projecting the 1992 EOC distributions are RPC for WEXTEX and 8x1 for TSPs.
5.2 Growth Rate Distributions Crack growth rates are obtained between the 1989 and 19918x1 inspections. Based on the extensive 8x1 and RPC comparisons obtained in the 1991 'ispection, the 8x1 data analyses were reassessed to increase the sensitivity to calling indications with the 8x1 5-2
probo. Based on this increased detection sensitivity, the 1989 8x1 data woro roovaluated for each indication found in 1991. The results of this reevaluation are summarized in Tablo 5.1. It is soon that over 50% of the indications found in 1991 were found to have indications in 1989. As noted lator, this porcentago increases to over 60% when only the number of tubes inspected in 1989 is considered. Tablo 5.1 also providos the distribution of 1989 8x1 Indications relativo to the number of colls showing indications. This distribution of indications is later utilized to project the 1991 crack distribution. The porcentage distribution of indications left in servico is not significantly different betwoon all cracks including multiple col'S and that excluding multiplo col'S, Growth rate data based on the 8x1 data betwoon 1989 and 1991 is given in Tables 5.2 and 5.3. The growth rates are given in terms of the increase in the number of 8x1 coils with indications (number of hits) por operating cycle. The average growth in terms of 8x1 hits is shown in the right hand column of the tables and is typically about 1 hit (order of 30 to 45') per cycle. Comparing Tables 5.2 and 5.3, it is soon that the growth rato distribution for the Indications at TSPs is essentially the same with and without inclusion of multiplo col's, including multiple Col's increases the growth rate for WEXTEX indications by about 20% (order of 10') which is a relatively small difference. The similarities of the 1989 indication distributions and the growth ratos with and without mult'ple col's indicate that the 1992 EOC projections can be based on use of single crack data which is also desirable for structural considerations as noted above. The conversion of 8x1 hits to crack angles is developed below. Tablo 5.2 also shows the 1989 8x1 tabulation betwoon detected indications, no detectable degradation (NDD) and not inspected. For the tubos inspected in 1989,64% and 67% of the TSP and WEXTEX tubes inspected respectively, had circumferential crack indications. Assuming the same percentages apply to the uninspectoo tubos,64 of the 101 TSP indications in 1991 and 119 of the 178 WEXTEX PWSCC indications woro present in 1989. These values together with the distribution of indications left in service (Tablo 5.1) and the growth ratos of Tablo 5.2 are used in Section 5.5 to comparo the projection methods with the 1991 RPC inspection results. A best fit correlation between 1991 RPC crack angles and the number of 8x1 hits is used to assign crack ang;os to the 8x1 distributions. These correlations are shown in Figures 5-1 and 5 2 for WEXTEX and TSP indications, respectively. The WEXTEX anu TSP i 53
l l correlations show, respectively, intercepts of 39' and 35' and slopes of 37'/hlt and 29'/hlt. ..ie correlation coefficient, R, of 0,77 for the TSP and 0.76 for the WEXTEX indicates comparable accuracy for the 8x1 and TSPs compared to the WEXTEX transition. in general, improved 8x1 accuracy is expected at the TSPs as the larger 0.3 inch 8x1 coil diameter would be expected to have greater lift off effect at the WEXTEX transition than the 0.105 inch RPC coil. For the more smoothly varying diameter changes at the dented TSPs, the coil lift off effects are not expected to be as significant as at the expansion transition. The single 8x1 hit data was not included in the fits as inclusion of these data significantly reduces the slope and the correlation coefficient, it is believed that detection of the smaller cracks angles is sensitive to the 8x1 coil spac, q as also indicated by the 35' to 75' detection threshold estimated for the 8x1 probe. This detectability apparent'y resuits in one 8x1 hit where 2 hits might be expected. In developing the correlations of Figures 5-1 and 5-2, 8x1 data that deviated extensively from the data trend were reevaluated for the number of hits in 1991, in most cases, the initial calls underestimated the number of hits, The resulting data adjustments improved the correlations with resulting increases in tb slope and growth , rates with reductions in the EC uncertainty. These changes an improved the comparison betwen 1991 predictions and RPC measurements for the crack angles. . Two of the 101 TSP tube total mentioned above were not included in the 8xt-to RPC curve fit of Figures 5-1 and 5-2 and subsequent TSP figures: SG B RSC87 at H3 with an RPC angle of 106' and two 8x1 hits in 1991, and SG-B RSC9 at H2- with an RPC angis of 72* and two 8x1 hits in 1991. These tubes have a negligible effect on the 8x1 vs. RPC correlation. Since they were not inspected in 1989, they also do not enter into the growth rate distributions, and therefore are neglected. For WEXTEX tubes, two tubes also were not included: SG-A R15C50, an indeterminate indication, and SG C R32C49, and R10C25', with a 126* RPC indication. These are also judged to have a negligible effect on the RPC vs. 8x1 correlation. The correlations of Figures 51 and 5 2 are used to convert (see Section 5.5) 8x1 hit - distributions to crack angles such as for the indications left in service in Table 5.1. The slopes of these correlations are used to convert the 8x1 growth rates of Table 5.2 to crack angle growth. Figures 5-3 and 5-4 show the growth rate distributions as interval dit,tributions (% of indications with noted growth angles) and cumulative probability distribut;ons. 5-4
For defining crack growth models, it is necessary to recognize that the inspection growth - rates represent the total crack growth beyond the detection threshold. If the detection threshold is assumed to be 40 50% depth, the growth rate indicates the total growth in depth beyond 50%. It is conservatively assumed in this evaluation that the growth rates apply to through wall indications. Based on the ODSCC and PWSCC crack models developed from the pullsd tube examinations, as discussed in Section 3.3, the crack growth models applied in this study are shown in Figure 5 5. North Anna 1 pulled tube R11C14 with ODSCC indications indicated locally deep cracks with partial depth, multiple crack initiation around the remaining 360* of the tube circumference. The BOC growth model fur ODSCC is thus comprised of an assumed 50% deep crack with a
>50% deep crack at the detection threshold. At the EOC, it is assumed that the deep crack grows by the sum of crack growth plus 8x1 crack angle uncertainty as shown in Figure 5 5. The through wall crack length is estimated at up to 60% of the projected RPC crack angle.
The pulled tube results for WEXTEX PWSCC circumferential cracks show essentially continuous cracks with no partial depth cracking around the tube away from the principal crack or crackc. Thus the PWSCC cracks tend to grow by multiple initiation sites as a progression around the tube circumference. The BOC PWSCC model of Figure 5-5 therefore includes a segmented crack at the detection threshold with no crack present over the remaining tube circumference. The EOC model then adds the growth over the cycle to the segmented crack. As noted in Section 3.3, the multiple initiation sites for PWSCC circumferential cracks result in small ligaments remaining between crack initiation sites. A segmented crack model of short (~0.2") through wall cracks separated by ligaments is used to represent the PWSCC crack morphology for analytical evaluations rather than uniformly through wall cracks. Although the multiple initial sites and remaining ligaments have been found for the partial depth ODSCC indications in pulled tube R11C14, the through wall cracks indicated loss of remaining ligaments. Based on this pulled tube result and pending further evaluations of ODSCC circumferential crack morphology, the ligament model is not applied for ODSCC cracks in this study. 5.3 Eddy Current Uncertainties The eddy current uncertainties for the 8x1 probe can be developed from the correlations against the RPC probe given in Figures 5-1 and 5-2. The RPC crack angles are 5-5
s sufficiently accurate to use as the reference for estimating the 8x1 probe uncertainties. This process leads to inclusion of the increased RPC detectability into the 8x1 uncertainty. , The 8x1 uncertainty distributions are obtained from Figures 5-1 and 5-2 as the _ difference between the measured RPC angles and the best fit correlations. The resulting uncertainty distributions are given in Figures 5-6 and 5-7. As seen from a comparison of the two distributions, the 8x1 uncertainty for the WEXTEX transition is greater than for the TSP Indications. The WEXTEX uncertainty distribution does not follow a normal distribution and shows a low frequency occurrence of angular uncertainties up to 75*. These EC uncertainty distributions are used in the projections of the 1992 EOC crack angles. Noting that only cracks below the detection threshold were left in service at the 1991 inspection, the inclusion of these uncertainties in the projections is particularly conservative for the WEXTEX transitions which were inspected with an RPC probe. The 8x1 uncertainties are applied for the 1992 projections on the basis that 8x1 growth rates are used for the projections. 5.4 Estimated Number of 1992 EOC Indications The potential number of circum,'rential cracks at the 1992 EOC is needed to combine with the crack ang!e proj*u a obtain the potential number of tubes at a given crack angle.
~
Westinghouse experience in estimating tube plugging projections has shown that l ja.c, For the circumferential cracks at TSPs,110 indications were found in the 1991 inspection. As noted in Table E 1, 66% or 72 indications (including 66% of tubes not inspected in 1989) were found by review of the 1989 8x1 data to have been present in 1989. Figure 5-8 shows the log normal projection of the 1989 and 1991 cumulative number of indications to estimate about 154 cumulative or 44 new indications at the - 1992 EOC. Based on assessing the sensitivity of these projections, the upper bound on new indications by the EOC is estimated as 100 new indications. The upper bound - estimate is used in this evaluation to enhance the conservatism in estimating the number of tubes at a given crack angle. It can be noted that these estimates are based upon 8x1 detection thresholds. A more sensitive probe could potentially detect larger numbers of 5-6
indications but the greater numbers would be at smaller crack angles or less depth and - would not be expected to represent structurally limiting tubes. The WEXTEX circumferential crack estimate for 1989 includes 64% (Tabte 5.3) of the 216 1991 Indications or 139 indications found by 6x1 data review to have been present in 1989. In addition,73 WEXTEX indications were previously found in 1989 and prior yoars so that tho adjusted 1989 number of cumulative indications is 212 and the 1991 cumulativo indicatioes is 289, The log normal projection for WEXTEX circumferential The upper ndications yleids 382 cumulative indications or about 93 new indications. bound on 1992 new indications is estimated at 150 and this value is applied for the 1992 EOC projections. 5.5 Methods Evaluation Against 1991 Crack Distribution Results Projections of the crack indications left in service in 1989 to 1991 permits a comparison betwean the projections and the actual crack angle distributions found in the 1991 inspection. This comparison provides a verification step for the methods applied
- to make the 1992 EOC projections. This methods evaluation is described in this section.
As noted previously, the projections are based on single circumferential crack indications as found by the RPC probe. For the 178 WEXTEX singlo crack 1991 Indications, 67% or 119 were found to be present in 1989 and 59 were new indications. The now ludications are assumed to be present at the Bx1 detection threshoto (Sectioa 5.1) and are assigned a linear distribution from 0% at 35' to 100% at 15* The 1989 indications have the 8x1 distribution given in Table 5.1. The resulting distribution of the WEXTEX crack angles left in service in 1989 is shown in Figure 5 9. The WEXTEX growth distribution of Figure 5 3 and EC uncertainty of Figure 5 6 are applied to the 1989 indications left in service to obtain the 1991 projection. Monto Carlo sampling methods are applied to obtain the 1991 distribution shown in Figure 5-10. The projected distribution of crack angles is compared with the actual distribution in this figure. The goal of the projection methods is to adequately predict the largest cracn angles which could potentially be From structurally limiting and the probability of occurrence of the large crack angles. 4 Figure 5-10, it is seen that this goal is conserva'ively satisfied in that the largest , i actual crack angle of 245' is overestimated at 305* and the number of tubes at large l crack angles is slightly overestimated. The peak of the actual distribution near 65' to 5-7 ! 1
[ 115' is underestimated by the projections. This is due to the high growth rate tail of the ' Figure 5-3 distribution and the broad EC uncertainty distribution of Figure 5-6 applied to the projections. Overall, the agreement of Figure 5-10 confirms the adequacy of the projection methods and that the projections are conservative. "l For the 101 single crack indications in 1991 at TSPs, 64% or 65 indications were found in the 1989 and 36 were new indications. The new indications are again assumed to be distributed using the 8x1 detection threshold and the 1989 indications have the 8x1 distribution of Table 5.1. The resulting distribution for the 1989 crack angles at TSPs left in sulce is given in Figure 5-11. The TSP growth distribution of Figure 5 4 and EC uncertainty of Figure 5 7 are applied to- the 1989 indications left in - service. The Monte Carlo methods then yield the projected 1991 distribution shown in Figure 5-12. The agreement between actual and projected crack angles is excellent both with regard to the maximum prack angle of about 215' and the number of tubes at large crack angles. The predicted distribution is somewhat broader than the actual distribution, again in part due to the EC uncertainty distribution of Figure 5 7 applied for the projections. The good agreement between the actual and projected distribution as shown in Figures , 510 and 512 support the acceptability of the methods for projecting 1992 EOC crack angle distributions. This good agreement against 1991 RPC measured crack angles indirectly supports the adequacy of the 8x1 data for growth rates and indications left in service in 1989. Application of the methods and supporting data base can be expected to . yield moderately conservative 1991 projections. 5.6 EOC Circumferential Crack Estimates at 90% and 95% Probability Levels
-The ' distributions developed in Sections 5.1 to 5.4 are used in later sections to develop EOC circumferential crack angle distributions in order to permit assessment of the -
probability of limiting crack angles. To assess the magnitude of the crack angles and
~
. contributing factors without the complexities associated.with probability distributions, . the EOC crack estimates with all terms evaluated at the 90% and 95% cumulative probability level are developed in the section. The contributing factors are the detection - thresholds (Section 5.1), the growth rates of Figures 5 3 and 5-4 and the EC uncertainties of Figures 5 6 and 5-7. The results are given in Table 5.4 for both the 5-8
TSP and WEXTEX locations with each contributing factor at the 90% and 95% cumulative
+ probability level.
At the 90% probability level, projected crack angles at both locations are in the
-141* 146' range, At 95% probability, the estimated EOC crack angles are 165* for ODSCC at the TSPs and 169' for PWSCC at the WEXTEX transition. The 8x1 detection threshold 'q seen from Table 5.4' to be the largest contributor (73* of 165') to the projecte, JC angles for ODSCC at TSPs. For the WEXTEX transitions, growth rate is the principal contributor (73' of 169') to the estimated EOC angles at the 95% confidence -
levels. Growth rates, which are dependent on the length of the operating cycles, are seen to contribute about 40% of the projected EOC circumferential crack angles. 5.7 Projected EOC WEXTEX Crack Distributions Since all detectable circumferential cracks were removed from service at the last outage, any cracks Icft in service at the return to power would be below the detection thresholds. . For projecting the 1992 EOC crack distributions, it is conservatively assumed that the number of cracks projected by the EOC were all left in service at the
^
detection threshold crack distribution angles. The data base for projecting the EOC crack distributions, as developed in Sections 5.1 to 5.4, is summarized in Table 5.5. Applying the Monte. Carlo sampling methods to the noted distributions leads to the projected WEXTEX EOC crack distributions shown in Figure 5-13. The Monte Carlo methods yield
= the probability of a given crack angle times tha projected 150 indications. - For ease of interpretation, the results were rounded to the nearest whole number of indications.
The largest crack angle was assigned at the angle at which the tail of the distribution integrated to greater than one half ot an indication. For consistency _with the methods validation against 1991 data in Section 5.5, the acceptable Monte Carlo samples were required to project crack angles greater than 45' as typical of indications reported as circumferential cracks.
.. The results of Figure 513 indicate that the largest 1992 EOC WEXTEX single crack angle is expected'to be about 245*. Based on the overestimate in predicting the maximum crack angle for the 1991 crack angles (Figure 510), it is expected that.the I 245' projected crack angle is also overpredicted by 20' or more.
- '59 i
l I
5,8 Projected EOC TSP Crack Distributions The EOC projections for circumferential cracks were performed similarly to the WEXTEX projections using the data base of Table 5.5 The projected distribution for the upper bound 100 indications is shown in Figure 5-14, From this figure it is seen that the maximum 1992 EOC TSP single crack angle is expected to be about 185* with about 3 indications with crack angles above 155*. The TSP indications, based on 8x1 detection thresholds, peak at 85' while the WEXTEX indications of Figure 5-13, based on the smaller RPC detection threshold, peak at the lower cutoff angle of 45'. 5.9 Distribution of Indications at First TSP Elevation The prior sections developec the distributions for the total number of indications summed over all TSP elevations. The total indications are distributed to the top and bottom of the 1st TSP in this section. The key TSP elevation for evaluation of crack propagation by tube vibration is the lower edge of the 1st TSP. Circumferential Indications have followed the axial PWSCC indications in time and it is reasonable to expect that the axial indications are also a . predecessor by TSP elevation for the circumferential indications. Thus the 1991 axial indications are used to guide the estimate for circumferential indications at the 1st TSP for the 1992 outage. Table 5.6 summarizes the inspection results for axial and circumferential distributions by TSP elevation. Conservative estimates for 1992 EOC circumferentialindications at the upper and lower edges of the 1st TSP are also given in Table 5.6. Both the axial and circumferential distributions indicate a n6gligib!e likelihocd of circumferential indications at the top TSP, The applied values 1or TSP elevations are that 45% of the total S/G indications occur at the 1st TSP with 25% on these indications at the lower edge of the TSP with no or insignificant indications at the top TSP. 5.10 Summary of Probabilities for Exceeding Acceptable Through Wall Crack Angles Table 5.7 develops the distribution of the total number of circumferential indications to , the key locations within the S/G for tube vibration considerations. The results for TSP indications show 11 or 0.37% of the tubes at the lower edge of the 1st TSP. For WEXTEX indications,150 are expected in all S/Gs nith 149 (8.28% of tubes) in the 5-10
t contral sludge region and 1 (0.08%) in the periphoral region. It was shown in Section 4.2 that very few WEXTEX indications havo boon found outsido the expected sludge deposition region. This is expected to result from the highor tube wall temperatures in the central sludge region as a consequence of the insulating effects of the sludge deposits. Studge deposits of significant depth (>0,75") can increase the tube ID wall temperature at :ho WEXTEX transition at the top of the tubosheet from 600' - 60a'F typical of a free span location to the T-hot temperature of about G20 F. The higher temperatures in the sludge region can increase the initiation rato (small fraction of indications found in periphoral region) and the growth rato. The growth rato at the lower temperature would be expected to be about 50 60% of the 15*F higher temperature, The crack angle probabilities for the peripheral region are likely to be about [ la,c of the Figure 513 values for the contral sludge region, although this is ignored in the present statistical estimatos. The applications of the circumferemial crack distributions to obtain probabilities for events such as exceeding a crack angle for tubo burst or tubo vibration for a specified type of crack and location can be obtained from:
'C O
e m 5 11
~ Bsc For tube vibration, the number of tubes susceptible to craca propagation increases as the crack angle increases above the threshold angle (0) for the most limiting tube. In this case, the probability for crack propagation must be summed over the increasing threshold angles such as:
_ _ 8.C a where [
}b,c, For combined crack conditions such as circ *;mferential cracks at both the WEXTEX transition and bottom of the 1st TSP,-
_ a.c The limiting crack angles for tube burst are developed in later sections of this report. However to simplify the discussion in later sections, the probabilities associated with exceeding acceptable angles are developed in this section. The probabilities are i developed but d:scussion of the acceptable angles and interpretation of the results is j included in the sections developing the crack angles. 5-12 i
. . .. . ~. . . _ - . ~ - - . - . . = . . . . - - . _ - - .-
I Table 5,8 shows examples of calculated tube vibration propagation probabilities for , ODSCC at TSPs with a through wall crack angle model and with a through wall crack plus ( 50% deep crack over the remaining tube circumference. The latter model shows an [ example requiring summation over varying numbers of tubes subject to crack , propagation as a function of through wall crack angle. [ ' f I I i Ja,c, : i A summary of resulting probabilities for evaluations supporting tube burst and tube { vibration are summarized in Table 5.9. It is again noted that P po represents the number. i
. of tubes that at EOC conditions exceed the minimum acceptable crack angle for through l wall crack propagation. For Ppg<1, P po can be interpreted as the probability that one tube in the three S/Gs exceeds the through wall crack threshold angle. for crack !
propagation. ! [ P i
'{
i 1 I 5-13 i i
- s Table 5.1 Circumferential Indications Left in Service in 1989
'89 8x1 Indications in Service ,
No.11 No. '89 S.G 8x1 In6c's BXLL% 1 2 3 A 5 5 TSP - ALL CRACKS I~ A 31 9 7 2 0 0 0 0 B 21 14 8 4 1 1 0 0 C 58 32 25 4 2 1 0 0 Totals 110 55 40 10 3 2 0 0 Percent = 72.7 % 18.2 % 5.5% 3.6% 0.0% 0.0% TSD - WITHOUT YULTIPLE CtRC. CR ACKS A 31 9 7 2 0 0 0 0 B 17 10 6 3 1 0 0 0 C 53 30 24 4 1 1 0 0 Totals 101 49 37 9 2 1 0 0 Percent - 75.5 % 18.4 % 4.1% 2.0% 0.0% 0.0% WCXTEX - ALL CR ACKS A 40 19 9 6 0 2 1 1 B 28 20 12 5 2 1 0 0 C 76 51 28 15 7 1 0 0 Totais 144 90 49 26 9 4 1 1 Percent . 54.4 % 28.9 % 10.0 % 4.4% 1.1% 1.1 % WEXTEX WITHOUT MULTIPLE CIRC. CRACKS A 37 17 8 6 0 2 0 1 8 25 17 11 4 2 0 0 0 C 53 41 20 14 6 1 0 0 Totals 115 75 39 24 8 3 0 1 Percent - 52.0 % 32.0 % 10.7 % 4.0% 0.0% 1.3% , Rev.A , 5- 14
Table 5.2 8x1 Growth Rate Data for CircumferentialIndications- Single COls
'89 ~91 Growth:
increaw ;.; 8x1 Coils with Indications
'89 Data No. '91 No. '911 t Averace Not inso. 50 1 2 3 4 5 SG 8x1 Indic'ns RPC Indic*ns lodications UDD TUBE SUPPORT PLATE INDIC ATIONS 5 6 5 0 0 0 1.00 9 7 15 A 31 31tt 4 3 5 2 3 0 0 1.38 17 17tt 10 3 B
5 12 24 9 3 0 0 1.06 53tt 30 18 C 53 20 35 16 6 0 0 1.10 ( 101 49 (64%)t 28 24 Totals 101 PERCENT == 26.0 % 45.5% 20.8 % 7.8 % 0.0% 0.0% WEXTEX INDICATIONS 15 10 10 0 0 1 0.97 17 19 1 A 37 49tt 9 9 5 1 0 0 0.92 17 7 1 47tt i B 25 1 28 19 2 1 1 1 0.67 41 11 C 53 82tt 3 52 38 17 2 1 2 0.82 115 178ttt 75 (67%)t 37 Totals PERCENT = 46.4 % 33.9% 15.2 % 1.8% 0.9% 1.8% t Values in parentheses are percent of tubes inspected and found to have indications in 1989. tt Values are total WEXTEX circumferential cracks found by RPC. All RPC indcations at TSPs were found by the 8x1 probe.' Rev.A ttt Does not include the one ODSCC circumferential crack in SG-C. 5- 15
Table 5.3 8x1 Growth Rate Data for All CircumferentialIndications
'89 '91 Growth:
'89 Data - Increase in 8x1 Coils with Indications No. '91 No. '91tt Averace NDD _Not inso. s0 1 2 3 4 5 SG 8x1 Indic'ns RPC Indic*ns IDdical10Ds TUBE SUPPORT PLATE INDIC ATIONS 15 5 6 5 0 0 0 1.00 A 31 (31)tt 9 7 4 3 7 4 3 0 0 1.41 8 21 (21)tt 14 3 8 12 26 9 3 0 0 1.06 C 58 (58)tt 32 18 27 20 39 18 6 0 0 1.12 Totals 110 110 55 (66%)t 28 PERCENT = 24.1 % 47.0 % 21.7 % 7.2 % 0.0% 0.0 %
WEXTEX INDICATIONS l 17 10 11 0 0 1 0.95 A 40 52tt 19 20 1 7 11 9 5 2 0 0 0.93 8 28 53tt 20 1 2 30 11 7 2 1 1.07 C 76 110tt 51 23 23_ 4 58 42 27 9 2 2 1.01 Totals 144 215tti 90 (64%)t 50 PERCENT = 41.4 % 30.0% 19.3 % 6.4 % 1.4% 1.4 % t Values in parentheses are percent of tube inspected and found to have indications in 1989. { tt Values are total WEXTEX circumferential cracks found by RPC. All RPC indications at TSPs were found by the 8x1 probe. Rev.A ttt Does not include the one ODSCC circumferential crack in SG-C. 5- 16
Table 5.4 Circumferential Crack Estimates at 90% and 95% Probability Lovels EOC Crack Contribution at 90% Cumulative Probabl!ity levels EOC Crack Contribution at Noted Probability Levets Factor ODSCC nt TSPs PWSCC at WEXTEX 9.0%e 25 % 2.0.% 25 % Detection Threshold 8x1 Probe 71* 73*
. RPC Probe 45* 47' l
l Growth Based on 54' 67* 60* 73*
'89 '918x1 Data EC Uncertainty Based 21* 25* 36* 49' on 8x1 vs RDC ~
Correlation Total EOC Cire. 146* 165* 141* 169* Crack Angle 5-17 , t i
Table 5.5 Data Base for 1992 EOC Crack Angle Projections WEXTEX Transition TSP Edco Number of Indications 150 100 Detection Threshold 23* - 50 % 35* - 0% . o Assumed distribution of 50* - 100 % 75* - 100% cracks returned to service Growth Rate Distribution Figure 5-3 Figure 5-4 ' ^ EC Uncertainty Distribution Figure 5-6 Figure 5-7 l 0 5 18
i i i Tab!o 5.6 Distribution of TSP Indications by TSP Elevation . i r Tube Support 1991 Axlal 1991 Circumferential 1992 , Plate No. Indications at TSPs IndicatioDs at TSPs Aoolled Values 1 . Total 43.9 % 41.7 % 45% - Upper Edge 80% i bwer Edge 20% 25% f 1 i 2 Total 34.9 % 41.7 % ! 3 Total 11.6 % 10.2 % i 4 Total 5.4% 4.6% t 5 Total 3.0% 1.8% ; 6 Total ' O.9% 0%
-+
7 Total 0.3% 0%- i [ t b f 9 i
?
I 5-19 i i e 4
Table 5.7 Number and Probability of Circumferential *
- Cracks at Key Tube 1.ocations
. itgm - TSP E&e WEXTEX Traneition Total No. of Circ. Ind. 100 150
- in all 3 S/Gs Number by TSP Elevation Total at 1st TSP 45 (45%)(1) --
+
No, at Lower Edge of 1st TSP 11 (25% of 45)(1) -- Number of WEXTEX Indications 1(2) in Peripheral Region No. of Tube Susceptible to Circumferential Cracking
+ Sludge Pile Region for WEXTEX ---
1800
+
Peripheral Region for WEXTEX -- 1200
+ Total Unplugged Tubes for TSP 3000 --
Fraction of tube estimated to have Circumferential Cracks -
+ - WEXTEX Sludge Region ---
0.0828(3)
+
WEXTEX Peripheral Region -- 0.00083 Lower Edge of 1st TSP 0.0037(3) ... Notes:
- 1. Based on 1991 Inspection Results and TSP Elevation Dependence Found for PWSCC -
Axial Cracks
- 2. Based on North Anna and industry experience (Section 4.2) showing that <1% of ,
WEXTEX indications occur in peripheral region
- 3. It can be noted that probability of a WEXTEX and TSP lower edge crack on the same ,
tube is approximately 0.0828 x 0.0037 - 3x10 4 5-20
Table 5,8 Examples of Calculated Probabilit es Acceptable a,c LQCation TW Angle p 0 Po)
- b,C Lower Edge of 1st TSP 6 10 ' !
o Peripheral Region ' o Turbulence ! I o Through wall Crack ' r i Lower Edge of 1st TSP 0.0001 o Peripheral Region - 0.0020 o Turbulence 0,0054 o Through wall (TW) 0,0019 ,, Crack + 50% 0.0001 deep crack 0,00002 " 0.000001 0.
~ ~
0.0095 . 1 h Note 1. [
}b,c, t i
!O i I i 5 - E1 f r
l Table 5.9 Summary of Probabilities of Exceeding Acceptablo Crack Angles Minimum Acceptable ac o Evaluation Topic TW Crack,0 Ppg(1) Tube Burst b ,c ;
~
o WEXTEX Segmented:3aPNO ! 0.0036 o WEXTEX TW:3APNO 0.15
- APSLB -0
- o -- TSP TW + 50%: 3APNO -0 Vibration in 1st Sogn o WEXTEX: F.E., Central 0.001 F.E., Peripheral 0.0014 Turb., Central ~0
\
Turb., Peripheral 0.0038 o TSP: F.E., TW -0 F.E., TW + 50% -0 Turb., TW 6x10-7 Turb., TW + 50% 0.0095 o WEXTEX + TSP
- Turb., Periphera 6.9x10-6
- Turb., Central 1.3x10-7 Notes: ,
- 1. Probability that one tube in the plant exceeds through wall acceptable crack angle.
I - jb,c
- 2. Represents product of WEXTEX and TSP probabilities.
5-22
'l Figure 5-1 WEXTEX '91 Data: RPC Angle vs. Number of 8x1 Coil Hits 350 .1 1 I l,
~
~
I i I---- l- - Y = M0 + M1*X 300 _ J ~ ~- J ~ ~~~ - t Single amt liits MO 38.6
/
_ i M1 36.95 250
- - ~ ~ -- -
2 l~V T-- ~~~~ ~ J j ~Q- R o.758 e m ' o i G 9 i k 200 ---
---- i f g-f-Oo ' - -- 4 -~
0~ -- -- - - b ~ 9 ; y 150 ~
~~ i- -- ~~-- - - - ~ "---- ~ " ' I -- ~
~~-~~ --
~
o D. - n- 100 -'- ~ ~ -
~
noi,3: j + Sogle 841 Hts Not included g n Curve FA _. . _ g _ ._ -.. Muttpie Cots Exciuded 50 __ . . _ - . _ . . . . . .w . 4_-...._... __ _.._ i 1 . ! 0- 4 l ! i 0 '1 2 3 '4 5 6 7 8 wxix s e.inc . . n un ilumber of 8 x 1 Coll Hits S- 23 _,. . - - - . . . _ _ _ . , , . . . . _ - . . . . - , _ . . . . . _ .,-.,. _ - . . , . . , . - . , . _ - - . . _ ~ , .. , _ . . . _ . , _ ._ _ _ . . , , _ - , . .
Figure 5-2 North Anna 1 TSP '91 Data: RPC Angle.vs. Number of 8x1 Coil Hits 360- - I e 315 .
~T '
'l .
270 + -- _ ~ i 1 1 to j t_inear Curve Fit
$ $ .! Excluding Singlo 8x1 Hits.
E>
'J 225 I i t t- -fO r -
Single 8x1' liits } I O /i - 6 e 180'
$~
0 d 3 1 k - ( ' O
+
Y = M0 + M1*X - 135* c-- g _ i un 33,o 0-g .
- Mt 28.6 .
90 j + -- -- R 0.767 d , 4~ I ' 45 .
-E-e-) i 0
0 1 2 3 4 5 6 7 8 iseo.,w . o....n Number of 8x1 Coil Hits S * * *
. Figure 5-3
^
.WEXTEX Cumulative and interval Growth Distributions 100% I ... I ,, !-
m u -- m .
~ ~ ~~'-*- ~ ~ ~ ~
~ 80% ~ ~ ~ ~ ~ ~ ~~ - ~ ~ '
1-- ~ ~ - + - - - - b C O
---~ ~~ I-~ ~~'~~~ ~'"~-~ - - - --
a o 60% - O Percent E ' O Cumulaave Percent (%)
~
C - d
- 40% -' ~~~~
i * ~~~~'~~~~~' l~~~~~~~~i
~
t'-~
-~~ - - --
C - i e - l 0 ; l e Notes: l Q. - Muluple COls exduded
"~ ~'"'~~'~~~~ ~~ '*' "-"*-~~ ~ ~ ~ - '
20% ~ ~~~ ~ ~ ~ ~ ~ ' ~ ~ ~ ~ ~ - i l 0% ~ I
~ - - " - "
H P - l
-20* O 20 40 60* 80 100 120 140* 160 180* 200*
m ia _ _ u. Growth per Cycle (Degrees) 5- 25 l.
_ <gv - - - -
'S.
Figure 5-4 TSP Cumulative:and Interval Crack Growth Distribelions 100% 'I. 'l I r i r
,, 80%-
- ~~ T t--
r-:- c f S ! i I - g
$ i b a
~~ ^
*~~~~ ~~~~-~~'
~ ~ ' '
~ ~ ~ ~ ~ ~ ~
saunval orowm 5 60% T o comueve crowe n . i
;R - 2 _ j 5C 40% ~ ~ ~ ~~ ~~~
~ ~ ~ ~ ~ j ~~ ~ ~ ~ ~ ~
e U -
. ! l
- I
~ ~~~~~~'~'~ ' ~ ~ ~ ~ ' ' ~ ~ ~ ~ ~ ^I+~ T ~ ~~'-
20% ~~ ~ ~ ~ ~ ~ ~ ~ ' ~ I 1 0% -' l i '~l i
-20 0 20" 40 60" 80 100*
' Growth per Cycle (Degrees) 5- 26
.,.._...............?
t Figure 5 5 ;
. Crack Growth Models for ODSCC and PWSCC ;
I a,C ; ODSCC BOC : t, 435'-75' i
>50% - j
' ' MQ;. ;4 .
< Deteetability ;
/#' Threshold u-
,..< 7 ;;
- j. s, rf I j 2, - y 50% Depth
,i . j : c. PWSCC_ _ . BOC , 23' 50' :
/ V -
Detectability i
//
Threshold: 23* Throughwall 40' Segmented
- _L -A -,
- v/ u f -
l
~
[
= -.-
y 5- 27 ,
a FIGURE 5 6 WEXTEX 8x1 Probe EC Uncertainty Distribution I i 1 ! ' i 1 I l I t y 1 ) i f 1 14
! ! I l
j ' 0 ! ! ! I ! l
! I j '
i ; ; t - j
- - " - * -- - n,
, - "i- --- t-12 - * :--- - r, ,
i 5 ; i i , ! ! i i ! ! , Notes _ co , l ;
- I-- ~~~-' ' - ' ' ' ' ~ * ~ ~ * ' ' ' ' '-
g O f , ! ! i i l i ** $S' ewe " sat D**w nens "'o*n*e j . 4 i i re Enued
- j l ie! i i r ei i
- IMaine COis Ext-mio
- -+ - -- - - ----
o a .
= 1 l
j i l i i j i ; i' j- I {, ! t i . I * ! ! ! l < l 4 I
, l
^
[
- 6 i 7~-~t~-
i e e! ; i i ; i r m - 5 d i 4 q' , __.._.w___.._ . d l I j i i e3 i l z ! n n n g
.. . . ...-.-a.....
( 0 4
- !2- ! ,
i [i i 1 i
= ,-
4 i- R ii = 7P 4r .$r 45' -35' 4 IS' r 7 15' 25' 35' 45' $Y 65* 75' 83' 95* B:. ' midpoint (10 Degree Span) ri . ... Residual a Actual Calculated l WEXTEX Eddy Current Uncertainty
- w '
100% i E' i
!~ !
-v - ; O j ~
]
d 1 g ~#
> i a
= 8 0 % + ~ ~~ ' '~~" ~ ~ ~ ~ ~ ~'-~~~~0 ; , J
- g 3 '
!/ ;
i ! i 5 ! I i
~ '
c 3 60 % .
'd ,
) !
8 ! , I
!, i n
Q- 40% 7
*r i
e I ! ; 2 i + i To , : j i i . 3 , ! I i i i E "- ~ s 20% - i 1 0 > 4 i . i l t I l
- l. 1 1
j[ l i i . 0% - i i i i v0' -40' 20* O' 20' 40' 60' 80' 100' EC Uncertainty (Degeres) 5 28
Figure 5 7 TSP dx1 Probe EC Uncertainty Distribution 16 l , I *
=
14 4 12
~
10 . . 8 , 1 - 6, 4 , , 4 2 .- ,. [ 0-[ l l l l t t t
-45' 35' 25' 15' 5' 5' 15' 25' 35' 45' i I
EC Uncertainty (Degrees) l j
.I TSP _ Eddy Corrent Uncertainty f 100% -O !
{ J sos j_ t : l g 60% l i 40% i 3 .: D- 'h 20% . ~ = a 1 l os =" ! I k 5 29- ; i
- - _ . . . . _ _ . ..___..;___._ ...c.._..-,,__ ._._.......--.,___..m.... . _ - . . _-_...._-.,_.-._.r..,--.-.;_.-->
i Figuro 5J CircumferentialindicMl0n Projections at TSP
~
ah w I 6 i I i i gg t/rr+nluprobit2 ,-
/
/ 01 90 E
/
E / s 0 70 1 / s so yh' 5 50
% y, 40 #
Nai 30 #- +- g 20 g
'6 /
10 8
- s 4
U , / muALs 87.59939 __ 1991 7 l
/ -dr = 2.811811 __
0529
""~
eT r'2 J .9996278 ~~
')
l i i , l l 3
,i rvs=2A i i i 1
o i 10 103 1000 Tif1E (ETPY) S l 5- 30 l _
Figure 5-9 Distribution of WEXTEX Crack Angles Left in Service in 1989
^
100% i j A -[ l I i A
'~ ' - - ~
i i -- - 80% -
- '~~
7 C pO g , , , 4 -; _ f
~ ~ ~ ~ ~ ~
~~~' * - ~ ~
$ 60% - O f7 ^
o O- ii b 2 ~'~ 40% ~
~ ~ ~" ~
,'- ~^~i
-f ~ ~~~~ l ~ -W 3- O C -
E U 3 [ ; J
~^~"l ^*
- B*"'*"U "' -~~
~~ +~'~^
20% ~~ C - k.dications Below Detection Threshold
/
j e A!Isrdcabons
!^ i l ! ~
0% 0* 50* 100* 150* 200" 250* 300* m c,. . o.. Crack Angle (Degrees) 5- 31
J Figure 5-10 WEXTEX Circumferential Cracks: 1991 Actual vs. Projected Crack Distribution 35 -FH-F i i
-4 +1-t-H-FH-i-t--i ,
1---
,i I
30 ~
-~i ~~~ "- W"lj"-f+t- ~"t-=t ~
-i 1
i ~ j-1 1
; i; i l I i l
! ! i m 25 -~ ~
- i ~t'~ ~
* ~r~~~ "~*~ -'
i; ? + i-o I~~~ ir i. I i, Ja r i 3 it ,
; j
}- i
'-'~t-*
p " !T" ~'~ ~ ~J' ~ '~*~"'+' '~ g 20 -T l-- l u ? Il
! I G f i.
m s c i i
++- '~* i--
E 15 ~-+ i T g~+-' I s (
'-+-+t j
- ~ t- +--'
11 1998 Proj 3 z h! i ll '. s ll1 ta 1991 Act
++ 4 --
10 - -- j + '-- l ;
- " F "l --p * ~ '
_5 i2 s l i s .' s ; l
, .: l ,
9....l......,_..!._......,.__
< s ; <
_ t . _ i 9. . ... .! g i g g g s -
,_ i
. - e p '- r ,
) ,
, p h .~ h ~f 3 $ ,, I)
O~
' ' 2 ' ' i N M' "
l 'fl 3 bb-b--~-N-- t l S* 25 45 65* 85 105-125* 145 165 t85 205 225*245 265'285-305 325 345-
.i% Crack Angle (Degrees) 5- 32
r i t i
?
t i o ; i i o > I i @ ei - ; j e o 9 c i i I o o
- _ ___,.._& - _ . , . . _ . _ .- - 5 . . __ ;
I -
.g i
. I i
_ l t .. . - -- . .O b .u. .. _. . _oN f C i i'
- ; si ,
1 - n . vs
. i.
e a
, I I 8 >
y= il'< i . j m o CD
.._.__g
, vi
.2
..t . .
.-..-,7_
g n y 4 e e, , 6 g , g ; w s < ' : I i a e g _ . _ . . _ . .
$- _ _. l' .,, _...- _.e -_ _b < !
L s w .x w o o 9 .
.O \ M e s o i L O G, E :
W_ y -
._y....
N
. ._ _ o o
f f
- s-o i 1 -
N t i c j 's . :
.S 4
'\, o '.
~ .___.._._;. . - . . , . . . _ . _o i 8
l s lT i
.c !
a , m : b . . . . , . . . . o a N o
't b #
o
-o o
o o_ e e- , m (%) lusoJed eAll5lnWnQ g f 1 ! r s - L c
+
-i t
h r
o Figure 5-12 TSP Circumferential Cracks: 1991 Actual vs. Projected Crack Distribution 30 -t-l ! ! l l , I-i
^
i ^; !
! I -
3
~
~~'P' 25 ~- [-] Actual w/o ! - 93 Data Pomts (Each Data Set) l
- Sogle 8 1 Rts Eersnied
, l pg p,o g ;
c r j . 3 i l ! ! ! ! j~' "~ '
-f- %> 20 - -~'~~~~'-j~ ~ J ~~' ~ " 1 ~ i-
u [ t e i M . ! i j ! e j L - --
~-
tj -4 ! i- I 15 -- ~~- ' ~ 7 T - I i
~
3 ! jg ; g 2 1
$ 10 - -I'[-' - ~ ' ~ - '
f,, N Sj h i!! h 4 7-v i ~~ ' [
- , 9 x x - >
l h 7 9 ! 5- -
-g-" c i
-t ir-J )+ @ 3 If: C i ??
+- -' - - -
l
@M
! Wz s! M $ . ~ .g ',
1 --
+ o % -
3 ,~ . z' 3 ; v; % $ ,- - , ! l
- v .- >
b y y z : 4 8:
~
p 0 -' -k ' i $-i l- j $ ~ Ii I i i i k -kH S* 15* 25' 35* 45" 55* 65* 75* 85* 95*105115725'I35145"155'1651751851952C3215* RPC Crack Angle (Degrees) 5- 34
~
- o Figure 5-13 WEXTEX 1992 EOC Projections
~ 1 I I i i I C t i
- 25. . i. i .i i i i i i !
o 11 I i- = l i li-'l I i 1 l i! j l t j l ![-lgll!
! -1 i '
k i!' I
._{_ Hl- l53 -j l I I I i g _..._.j_.} p _ _ - _ _ .
_g*. T '-~ I l !l c i l. t 1i1 i *
.=
O i
~
Il I
- 1, l. ;
i l 3 li i ilile r t g l h [ill l l1!l l I l~ lI i i l-l l j - ' ~ ~
'-h-li
~~
-II+i :r'~'~I~t*T"'~iI-tit',ht-~ ~
ll O _. 3 l l I i i! o 1 i !l l! ! !1 ! !*
- -l l i I t t ! i '
-}y! . - .,-
1 l5 i! I! , I e 3 ,o -- 3 - !_n i l -y(3__ E 3
=
li
- i l
I s 'I i 1 il
*i= 5 i i e i Z I ll! ! l l' ! 2 i i g i j i
i i I 5~ ~*~4-~ _-'~illIl*l! 1~ { i l! i lli i i!!
*=
I i i
-t~i"*"~~
~'!!!
! i '!i ; i*
l-1l. s j
-- j
- ll r -
s I
+
!i F
; ! i 0
1
- ~ - ~ ' - ~ ~ ' ~ ~ ~ ~ ~ - -- -
~ ~I~ l 'I y 25* 45* 65* 85* 105* 125* 145* 165* 185* 205* 225* 245* 265* 285* 305* 325* 345' Crack Angle (Degrees) 5 35-
i o
~
Figure 5-14 ,
- -t TSP 1992 EOC' Projections i r
l-I 16 ! l } I-l-Fl+F - ' i i 1-F - l i I , i i 1
,ii 1-. :
i i l ' I!- i ! 4 g4 l .{ ..y p_ 3 _ _p . _ .. q.].._
, 4
. % ,._ c ,.._ p.7 _i _gl__j i__ p__
i 1 ! l-_ r ! -; i l 1 ! - l1l! I l!lr I! l I ! I ! i
-~~ t-~-
i ~!l l I, i . i ! I l c 12 - i i" l1 :
~i-"
~I ii.- i1 i-t t" i
- l
-t l- t~t^t-t + ~ "i !i t- ~It -
l*
=
i
- 5 : i i i r
o - . I i l! Iii 1 l l 2 1 .
= _i,_._ i ti t
,o ___, {_ _
_!pl4_;l_._pl+ . 4 :_p._.7_, I g* 1 l' [_ g . y !_ _ i 5 i f i i e i i i 'i i ii : i i i i S 8
-t-
*- J~ -'~~'~I~ M- ' ; -
-+
O
+- ! ;)
il i ji ~ ; , 5i '1 ll
! !ll
} .
O ^ r- -+- 1 , i !'1I* .
!l I
6~ "'I -}-i- --+i ,- --m , -i +1 r -
'I --
0 I
- t!
i i i z [ l _ i !li ij i ijj ji i }! i ; i _I_;l. 1 l [ i i i i l l{-.__.i.__--j._j._!_y;_.:j i .. {.! 4. !_l . 3 . ;4 . __ i __ ' 4__ 1i Ii! ' i ! ! " _. I ^ j i, j l j ! jt;-- i ! ! ! l . 2 i l' i l-
-" + +-'-+ i it
-! j-* * ; 1 1,--
- i. -
r .
- . t .
t 0
- - - - - - - -~- - - - -
; ' i- H - b N - M-L ! !!- .
1 5* 25' 45' 65* 85* 105* 125*145*165*185*205'225*345*265*285*305*325*345' 3 i Crack Angle (Degrees) : 5- 36 ; t
. , , - _e ..-- - - - - - - - - - - - ~ - - - - - - - - - - - - - - - - - - - - - - " - - ' - - - - - - - - - - - - ' - ~ ^
6.0 TUBE INTEGRlTY ASSESSMENT: TUBE BURST AND LBB M The NRC Regulatory Guide," Bases for Plugging Degraded PWR Steam Generator Tubes", issued for comment, addresses tubes with through wall cracking. Any through wall crack morphology that is projected to result in a condition such that the limiting crack morphology is exceeded during an operating interval when a corrosion growth allowance f for continued degradation and an eddy current uncertainty are considered, is unacceptable for continued operation. The Regulatory Guide utilizes safety factors on loads for tube burst and collapse that are contistent with Section ill of the ASME Code, j By also involvit.g 'an allorance for continued degradation and for nondestructive e examination detection accuracy, its use establishes a reactor coolant pressure boundary , that should have an extremely low probability of abnormal leakage, or rapidly propagating failure, and of gross rupture. The required confirmation that the leakage. rate limit being used is.less than the leakage rate of the largest permissible crack (leak before-break) completes the " defense in depth" approach of Reg. Guide 1.121. [ I t Per paragraph C.3.d (1) of Reg. Guide 1.121, the analytical and loading criterla 'I > applicable to tubes with through wall cracks in thinned and unthinned tubes are: j i i
~
1.. Through wall cracks in minimum thickness tubes should not propagate and f result in tube rupture under combined accident conditions. L
- 2. The maximum permissible crack tength of the largest sing:3 crack should be such that the burst pressure is at least 3.0 times the normal operating )
pressure differential, f 3, The leakage rate limit under normal operation set forth in the plarit technical j specifications should be less than the leakage limit determined for the largest f permissible crack. lt [ The projected end of cycle (EOC) condition of the steam generator tubes is compared to ; these criteria and is addressed below. I i i 61 j i I
6.1 Burst Strength of Tubing with Circumferential Cracks 9 The burst strength'of tubing with circumferential cracks is described in the following paragraphs. Burst tests were conducted on 0.875 inch diameter tubing with a wall - thickness of 0.050 inches. Since mill annealed Alloy 600 tubing has a high toughness and ductility the fracture mode of tubing w!!h cracks is typically one of plastic collapse. Hence the burst properties are controlled by the plastic flow strength of the tubing rather than the fracture toughness. This makes it possible to use narrow EDM slits to simulate the behavior of sharp cracks. ) i i Burst tests were conducted as follows. [ l l l \ l 1 4 Ja,c, ; Figurc 61 shows plots of burst pressure versus total through wall circumferential EDM slit angle. The lower curve is for specimens without any lateral support. The asymmetric cross section at the slit location leads to pressure generated bending >
. stresses. Consequentially as plastic col! apse of the net section occurs the test specimen plastically bends a substantial (45' or more) extent.
- In a steam generator, the presence of tube support plates, limits jatcral motion of tubing to a negligible level. The total diametral gap is approximately 0.016 inches. When lateral motion of the circumferential cracked test specimens is restricted by a simulated - ,
tube support plate. the burst pressure increases dramatically. Lateral support . . I , ja.c. The high burst pressures depicted in Figure 6-1 for the case of restricted 6-2 4
-- n --,.,L,-. ----L,.n -.n_.-,,.. ,,, .~, ,w.n-.w, , , . ,,a... , - . . . ~ , . , , . , ,,,.w, c.,,.e,,,.n_. , ,..m.,,,e
l i i i i lateral motion are in fact somewhat conservative, in most cases the crack opening at the l indicated burst pressure was sufficient to release the sealing bladder but crack tearing { was minimal or absent. l i At very high angles, in the case of restricted lateral motion, the axial force generated by f the and cap pressure must becomo sufficient to fracturo the net section in direct tension. l This is indicated by the final straight line segment of the middle burst curve. It was obtalnod by calculating the axlat pressure needed to generate a nominal not section stress , equal to the ultimate strength of the tubing. I I j i I t i l
)b.c, ;
I i' ( i t
]b,c. The tensile proportlos for the !
tubing used in these tests is as follows, yield strength,47.8 ksi, and ultimate strength, f' 105 ksi. h in order to examine the consequence of additional tubing degradation in the presence of through wall circumferential cracks, a series of tests were conducted using a complex l f crack geometry. [ t i
*' I
. 't I
]b,c, }
t 63 I
, _ . - ._ _u,_._..., . . _ . . _ _ _ _ _ _ _ _ __;._ _ - _
With undented tube support conditions, the presence of additional partial through wall , cracking did not affect the tube burst strength until the net remaining section failed in direct tension. This is illustrated by the lower straight line segment in Figure 6 2. , Here the nominal net section axial stress due to the end cap pressure force is equal to the tensile strength of the tubing. At smatter crack angles a bending mode of fracture is operative. As noted cartier plastic rotation at the cracked section is not substantial until other uncracked sections of the tube plastically deform. Hence the wall loss due to additional partial through wall cracks did not have a significant effect on the burst pressure. A weakening effect is only observed when the mode of failure is axial separation. The burst test data for the complex circumferential crack geometries is listed in Table 6.2 along with the tensile properties of the test tubing. To cover the case where circumferential cracks may occur at both the top of the tubesheet and at the first support plate location or if a crack develops between two adjacent support plates, several tests were conducted with lateral motion restricted by a simple rest stop,- that is the hole in the support plate was not simulated and rotation at the support plate location was totally unrestricted. As shown in Figure 6 3 this had no , significant effect on the burst pressure of circumferential through wall slits. 6.2 Burst Strength for Axial Cracks: Seating Bladder Considerations In testing of tubes with through wall slits [
}a,c,
, However, recent tests in Belgium have demonstrated that for axial cracks foil reinforcement of seals leads to the same burst pressures as in tests with very high *
- capacity pumps and no seal whatsoever. This is illustrated in Figure 6-4 as taken from the EPRI expansion zone PWSCC document. The Westinghouse and British curves do not use any seal reinforcement and lie below ths Belgian burst curve. The burst data without any sealant system falls along the burst curves where seat reinforcement was 64
i t i t i used. Therefore it is appropriate to remove the conservatism of the Westinghouse axial !
. crack burst curve and use the Belglan burst curve.
6.3 Leakage Model j An analytical model has been developed for predicting primary to secondary leak rates j through cracked steam gonorator tubes. Tho model is based on [ ( i f i Ja.c. The model predictions are shown to exhibit acceptable agreement with laboratory and l pulled tube test data. ! 6.3.1 Model Description t Overview: The crack leakage model assumes [
.I
( I I I t i r
'I je.c.
t i 5 65 i l t i
Governino Ecuations: The governing equation for flow through an axial crack is [
}b,c
~
_aC . Gulu im
,a.C f
a t
?
?
k 7 h
'L 6-6 t
_ _ _ _ _ _ _ . _ . . . _ _ _ _ _ _ _ . . _ _ _ . _ - - . . _ _ . _ . ~ . . . _ . _ _ _ _ . _ _ _ . _ . _ . . _ - _ . _ --_ a.c
) . i P
i t t i 1 i 1 .
- l. j i
i h F s t i 9 h
-l t
Critient Flow Model' For crack leakage flows, the possibility of critical flow occuring exists. ! When critical flow occurs, the rate of dischargo of the two phase mixture is limited to a maximum value dependent on [
]D 0 Due to its importance in accident analysis, two p'haso critical flows have been [
k studied extensively both theoretically and experimentally. The theoretical evaluations can bo '
,- divided into two general categories: [ Ja.c 67 >
I
,m, mA _sA,m%_,_kns.n.L-a.. .ns&m & W-,E-L-~L4ML3s &'tesM b4 de- mhm44~MA J & $ k- 3-- b- SB 43X4M M 3h e $5M -4 3 ~MA.'enA6Ab- m .h4" 4h-Enb*n,%1ask--e4 are~Mai-.*R*LM-a I
l i b 5,C 4 i q l 1
-0 P
I. s T
<a .
}
l 4-I M
-eas 68
__..2,.-....._. _ _._,. _ . . . - _ . . __.,. ... .
"&m.. -%4ho-"'
,{
8,C 4 i 0 0
+
l ., + M m 69
a,c
~
?
6.3.2 Computer Code Description l The equations presented in earlier sections woro programmed into a Fortran code called CRACKFLO. The program is executable on the CRAY and IBM PC computers. A brief description of key logic features follows, i [ 4: I i
}a,c, 6,3.3 Crack Leakage Comparison With Experimental Results The crack data base consists of field (pulled tubes) and Westinghouse laboratory forrned stress corrosion cracks.- Crack model results are compared with the test data in Figures 6 7 and 6 8 for normal plant operation' and steam line break conditions, respectively, The crack i
l- 6 10 L
, ,__ --- . - . . . - . - . - - _= . - - _ - - - - - model results are also compared to other experimental data (labeled crack data base in the figures). As indicated in the figures, good agrooment between prediction and measurement is shown. Data scatter is attributed to crack geometry parameters which are difficult to define with any degree of precision. Crack geometry parameters which affect flow are [ ] la c, j l l An error analysis of the measuremer'is veisus predictions was performed to obtain ths l standard deviation for uncertainty analysis. The standard deviation of the prediction is [ ]b c for normal operation and [ ]b,c for steam line break for the log log comparisons of Figures 6 7 and 6 8. For an uncertainty factor, t , the predicted leak rate l would be factored by [ ]b,c for normal operation and by [ ]b,c l for steam line break, i ! 6.3.4 References 8,C l l I i; i F f k 6.4 Tube Burst Capability 1.
'f The burst test data have been compared with analytical predictions. The good agreement f
of burst test data and calculations for tubes with through wall cracks provides for high [ reliability of extrapolation of test data to tube service conditions. This is accomplished [ 6 11 I I . . . _ _ _ .- ._ i
I I by adjustment of average flow strength to the temperature of interest and the use of ! lower tolerance limit (LTL) values. -f i 6.4.1 Tubes with Axlal Cracks -I Burst capability for tubes with axial cracking at TSP intersections has been I demonstrated by test to be a function of exposed cracn length. WCAP 12349. In Section i 6.2, it is demonstrated that the Belglan burst equation is the best representation of j burst capability of tubes with axial cracking ava!!ctle. For use at operating conditions, ! t the Belglan equation is applied to the inconel Alloy 600 tubing (7/8 x 0.050 inch, mill I annealed) with a 650*F flow stress of 63.0 kal per WCAP 12522. The resulting burst l pressure of tubes with single axial cracks is plotted in Figure 6 7. The single axial ! I crack burst curve also applies to multiple, parallel axial cracks for the longest crack in j the array (see EPRI NP 6864 L). The exposed axial crack length resulting in burst at j three times normal operating pressure differential (3AP) is [ ]b,c and .he l length for burst at steamline break (SLB) pressure differential (APSLB) is [
]b,c, !
I
'The presence of smallligaraents in the axlai crack networks has been demonstrated by ,
examination of pulled tubes at North Anna 1. WCAP 12349. The Individual cracks have aspect ratios (length / wall thickness depth) ranging from [ f
]b,c'In length. Tubes with colinear, axial, thru wall EDM slots have been tested and results provided in Table 3 3 of WCAP 8429. Figure 6 8 shows excellent j agreement between calculated and measu:ed burst capability when the burst capability is calculated with the equation :
t a,c . l: ! s where [ [
]b,c [
_,a,c - L 4 6 12 [ L- i
where ac [ i The segmented crack burst capability for three aspect ratios is plotted on Figure 6 7 showing the improvement in burst strength. [- t l I [ f I t f I 4 l i e
]b,c, l i
6.4.2 Tubes with Circumferential Cracks' < i The burst capability of circumferential cracks at the top of the tubesheet (TTS) or at the wge of the TSP's is based on [ ; I ja.c, , e f The equations utilized for unrestrained tube stress are: ; 6 13 f i e l
a'. Without ligament se aCs b) With ligament _ ,8,C
~
where
~ a.C 4
These equations are [ ja,c, Above about [ ]b,c of arc at room temperature (RT), the mode of failure changes from bending to axial pull separation. The equation for this portion of the curve is
.,a,c The application of the burst equations to SG operating conditions requires the use of LTL l:
strength properties, SF = 63 ksi, from WCAP 12522. L-I ., Figures 610 and 611 show the tube burst capability for uniform through wall and . segmented cracks, respectively. For uniform through wall cracks, the crack angles of' [. ]b,c are'obtalaed for 3aP and APSLB ,. respectively. With segmented
- cracks, the respective crack angtes increase to [ )b,c,
-. 6 - 14
.~ .. _ _ _ _ . _ . .._ _ ._ _ _ ._. _ _ _ . _ _ . - _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ .
l t i ! For consideration of the potential for a 50% deep crack to be undetected by ECT, {
~
additional burst tests were performed as described in Section 6.1 to assess the impact of. f this condition on the thru wall burst data. Figure 612 shows this RT test data plotted l versus the single crack LTL strength (at 6500F) curvo developed above. Sinco the tuoos are dented in the NA 1 TSP intersections, the data with full axial restraint may be indicative of actual strength, but these data are ignored duo to the difficulty of justifying j prototypicallity of the applied restraint. Considoring the 50% wall loss data with no axial restraint, the offect on burst capability was assessed to be negligible up to about
.[ ]b,C of arc where the reduced wall thickness affects the axial separation mode of failure. The 50% axial separation curve at LTL properties lies appropriately below the i data in the range of [ }b,c of arc. To be consistent with the prior RT data f with no wall loss on the romaining circumference, these data were scaled by the flow I i
stress ratio of.the test sarnplos for this comparison. Figure 613 and Figuro 614 J show the burst pressure capability of tubos with single and segmented circumferential f cracks for the inconel Alley 600,7/8 x 0.050 inch mill annealed tubing at 650*F. The f single crack allowablo are lengths with 50% wallloss are [ ]b,c for 3AP and [
]b,c for APSLB. The segmented crack case utilized has olastic ligaments of
[ f 0.05 inch that provido approximately normal operation AP additional strength as in the
- axial crack case. The individual crack lengths are taken at an average length of 0.25 j inch (5/1 aspect ratio) based on tubo examinations reported in WCAP 12349. The h i
a!!owablo arc lengths are [ ]b,c for 3aP and [ ]b,c for APSLB. ! 6.5 Loak Before Break (LBB) Capability ! t f The leak-beforo break rationale ls to limit the maximum allowable primary to
~
l secondary leak rato during normal operation such that the associated crack length is less l
. than the critical crack length corresponding to tube burst during a postulated steam line break . event. Thus, on the basis of leakago monitoring during normal operation, it is !
t assumed that unstable crack growth loading to tube burst would not occur in the unlikely j event of a limiting accident. Leak before break provides protection against a coupled f steam line break and tube rupture event which is outside of the design basis of the plant. ! The administrative leak rate limit for normal operation is 50 gpd per steam generator. l t The leak before break rationale for a single through wall crack, growing in an orderly ! fashion is straightforward. However, typical cracking patterns are more complex. '[ i 6-15 f 1 r--+,--- r c yv,__..m_.m.,
Crack networks rather than single isolated cracks appear most often. Pulled tube ' examination results and laboratory cracked samples provide examples of crack H networks. As discussed above, ligaments of material between through wall cracks in a crack network increase tube strength relative to that of a single through wall crack of - the same total length. While ligaments increase strength, they also decrease the leak rate relative to the same length for a single through wall crack. F .i.1 Tubes with Axlal Cracks The leakage model for tubes with axial cracks described in Section 6.3 was applied using NA 1 SG conditions to obtain the following predicted leak rates for normal operation and SLB as a function of crack length: Crack Leak Rates Length ILQ ILO S],8 (Irtch) (gpm) (gpd) (gpm) a Consideration has been given to the error of the predictions of leak rate compared to measured leak rates. In Section 6.3, the error is given as [ ja,c, 6 16
i I I I During normal operation, the leak rate limit ic 50 gpd for NA 1. Referring to the [t
- tabulations above and the mean curve of Figure 615 of leak rate versus axial crack }
length, the allowable leak rate of 50 opd will limit the through wall axial crack length to f [ ]b.0 Since the through wall 3AP crack length (Section 6.4.1) is [ f
]b,c, LBB is assured on a mean leak rate basis. Considering the predict; ors error for
=95% certainty, the axial crack length will be no longer than [ ]b,c. At this j length, AP SLB si satisfied with margin since [ ]b,c is acceptable.
i For segmented cracks, the conservative assumption is that the ligaments are elastic at { normal operation AP and the leak is therefore a minimum equal to the sum of the leakage ( of the Individual cracks. The macrocrack length is therefore maximized and strength ! minimized on this basis. The leak rates for segmented cracks are plotted in Figure 616 [ with the points for the typical 5/1 aspect ratio crack shown as the 0.25" segments line. [ For this typical aspect ratio, on a mean predicted leak rate basis, the 50 gpd limit f corresponds to a macrocrack containing l ]b,c, [ From Sect lon 6.4.1, the macrocrack length needed for 3aP burst capability is less than ! or equal to [ ]b,c, which is less than the leak limit pccmitted, [ ]b,c, f 1 However, LDB is assured for segmented cracks for APSLB since a [ ]D.C !
, macrocrack would have APSLB burst capability. [
6.5.2 Tubes with Circumferential Cracks The leak rate model of Section 6.3 for tubes with circumferential cracking results in the ! following predictions of leak rate for NA 1 SG normal operating and SLB conditions as a ! function of arc length: I Arc Leak Rates ! Leriath LLQ SLE ELS i (cegrees) (gpd) (gpd) (gpm) r
- - b,c {
t
- i
[ l I I 6-t7 ! (_._. - . _ - . , . . - . , . . - , , , , ,.,..m. . _ . . _ . , _ . _ . . . , _ . . . . . . . . _ . . . _ _ . . . . _ . , . . . _ . . _ . . . _ _ . . _ _ . _ . _ . . .
With the 50 gpd leak rate limit, the corresponding arc length of a single circumferential crack is slightly less than [ )b,c. Considerable margin is available versus the single - crack that lowers tube burst capability to 3aP, [ }b,c, and to APSLB'I }b,c (Section 6.4.2). 6.6 Leak Rate Under SLB Conditions This section develops estimated leak rates under SLB conditions based on tr.aximum leakage from the limiting crack leaking at 50 gpd under nnrmal operation or from the projected EOC crack distributions. Use of the EOC crack distributions is highly conservative in that the projected distribution would be expected to leak at wellin excess of 50 gpd based on application of the leak rate and crack modols used in this report. The potential for leakage from a single crack during SLB Is limited by the 50 gpd normal operation leak rate limit. Assuming operation at the 50 gpd limit, it is 95% certain that a single through wall axial crack (a conservative assumption) is shorter than [
]b,c (Figure 615). It is also a 95% certainty that a [0.5 inch]b,c axial crack would leak less than [ }b.c based on nominal SLB leakage of [
lb,c for 95% certainty. Applying the same rationale to the limiting circumferential crack of a [ }D C mauocn.ck, as developed in Section 6.5.2. leads to an expected leak rate tess than [ j
}b,c. Thus, the [ ]b,c SLB leak rate for the limiting axial crack would be the estimated SLB teak rate recognizing the 50 gpd leak limit.
In order to bound potential led age in the unlikely event that through wall or near through wall cracks that do not leak during normal operating conditions but may teak during a postulated SLB exist, a SLB leak rate can be obtained based on calculaied leakage for the projected EOC crack distributions. This calculation utilizes nominal leak rates , for the entire distribution. The EOC distributions utilized are the 1991 axial crack distribution of Figure 418 which is assumed to apply to the 1992 EOC and the , projected 1992 EOC circumferential crack distributions of Figures 5-13 and 5-14. The segmented crack modelis applied for the projected PWSCC axial and WEXTEX crack distributions. For ODSCC circumferential cracks, it is assumed that the through wall 6-18
crack length is [ )b,c of the projected RPC crack angles of Figure 514, The resulting total SLB leak rate is es'imated at about [ )b,c including contributions of about [ ]b,c from WEXTEX cracks, [ ]b,c from circumferential cracks at TSPs and s')out [ }b,c from axial cracks at TSPs. 6.7 Conclusions The test and analysis basis for burst capability and leakage of tubes with axial or circumferential cracklag provides a justification 'of structural integrity and LBB per Reg. Guide 1.121. Margin is demonstrated for burst capability during accident pressure loadings (APSLB) as single and segmented cracks are limited by the normal operation - leak rate limit of 50 gpd. A 34P burst capabill'y is assured in all cases for tubes with circumferential cracking and for single axia: cracks. Segmented cracks up to [
.]b,c in length are assured of 3AP burst capability but segmented craps lirrmu cnly by the 50 gpd normal operation leak rate limit may be longer. The strength of segrnented cracks has conservatively been minimized consistent with a minimum laakage model (elastic _Ilgaments).
The EOC crack length distributions projected for 1992 (Section 5.6) provide another aspect of the margin available for operation within Reg. Guide 1.121. The maximum EOC circumferential crack arc length is projected to be [ ]b,c at the TSP and to be [; ]D.C at the TTS. Historically, there is small or no leakage associated with these cracks and therefore they are anticipated to be segmented cracks or partial depth cracks. The allowable arc lengths for 34P capability are [ ]b,c for a segmented crack with elastic ligaments (minimum leakage model) and without or with 50% wall loss, respectively. For exposed lengths of axial cracks above or below the TSP edges, the maximum projected length is [ }b,c. On a segmented crack basis this is well
- below the [ ]b,c apSt.B capability length, it is short enough that a burst capability of about [ ]b.c would be expected versus a 3AP burst capability of - .[ ]b,c, 4
The potential SLB leak rate is expected to be limited by the 50 gpd operating leak limit to approximately [ ' )b,c for a single through wall axial crack, in the unlikely event that tubes with through wall or near through wall non leaking cracks during normal 6 10
operating con (ibns are present,2nd leak rates are calculated from the entire EOC c. rack distributions, the bounding leak rate would be approximately [ )b,c, . P
=
4 d - S. 3 9 w., 17 6-20 1 i
TABLE 6-1 BURST TEST RESULTS b,c M 4M 1
***O 875 INCH DIAMETER BY 0.050 INCH WALL MILL ANNEALED ALLOY 600 TUBIliG O.2% YIELD STRENGTH = 47.8 KSI, ULTIMATE STRENGTH = 105 KSI 6-21 l 1
TABLE 6-2 BURST PRESSURE FOR COMPLEX CIRCUMFERENTIAL CRACKS THROUGH WALL BURST CRACK ANGLE PRESSURE (PSI) RESTRAINT l'50 -
~ b,c 180 210 LATERAL 240 ONLY 270 300 150 180 210 LATERAL 240 PLUS 270 AXIAL 300 ALLOY 600 HEAT 1759 '
YIELD STRENGTH = 45.8 KSI ULTIMATE STRENGTH = 103.1 KSI _ _. a,c t em 4 G-22
h TABLE 6-3 BURST PRESSURE OF CIRCUMFERENTIAL THROUGH WALL CRACKS WITH RESTRAINTS SIMULATING CRACKING
- AT TUBESHEET AND FIRST SUPPORT PLATE LOCATIONS THROUGH WALL BURST SLIT ANGLE PRESSURE (PSI)
- - b ,c 1gg 240 300 - _
ALLOY 600 HEAT 2742 YIELD STRENGTH = 50.7 KSI ULTIMATE STRENGTH = 107.9 KSI e 6 e 6-23 l 1
a b,c e Figure 61, Burst Pressuro vs. Total Through Wall EDM Slit Angle 6-24
b,C w Figure 6-2. Burst Pressure vs. Total Through Wall EDM Slit Angle - Specimens With Through Wall EDM Slits and 50% Deep Partial Through Wall Circumferential S!its 6-25
l D.C Figure 6 3. Burst Pressure vs. Total Crack Angle - Lateral Restraint via TSP Hole vs. Lateral Restriant via Rect Stop 6-26
G-t
~
b,c I (- Figure 6-4. Comparison of Several Tube Burst Test Correlations Along With Lower Bound Tube Rupture Equation 6 27
Westinghouse Proprietary Class 2 Figure 6-5 Measured vs. Predicted Leak Rate (Normal Plant Operation) b,c "E" m 6-28
- ~ . . . .
Westinghous3 Proprietary Class 2 l e l Figure 6-6 Measured vs. Predicted Leak Rate _._ (Steam Une Break Conditions) b,c l i h 6-29 -
- . , . . . . . . . - . . . ,. ~ , . . . . , . . - . . . . - . .,,.- , . . - . . , . ,. .. - . . . - , . . . , . . - ,.
l t Figure 6 7 b,c e
)
h
- , asas 4
6-30
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N 4 qi d %e s D'C l
/
/
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\
+ \
\
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N
./ / / / ///'/'
- - - - -- - ~ '"
4 Figure 6 9 b,c 4 9 M w 6-32
i 4 Figure 6-10 b,C m I I e 9 k 6 4 I
+
p ( e t s T e e. h I I
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I F G-33 c 9 k
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4 _~ . x a _a. .- . . .. ..u..az.- , =.x =, . ~ . ~ , -
?
h 4 ii I i
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i t e !
.l Figure 6-12 ;
b,c !- M a
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b P
?
i t l t t 5 b i
?
L e - t 0
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M 7 F r I t
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f r v M
)
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- ,e
a 4 - 4 Figura 613 b,c l l o O: i i l I me s* === i-i l l 1 i r 6-36 l l
S S Figure 614 b,c M
- g :--
emme m e 6 37
1 Figure 615 b,C ei - esmum S e 6-38
, ~ . -y
I Figure 6-16 b,c S unuse O O 6-39 1 i 1
Figure 617 ' b,c 4 9 meus "
.9 i
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1 6-40 l 1 1 i 1 i
t i 7,0 TUBE VIBRATION ASSESSMENTS The purpose of this section is to provide descriptions of the methods, applications, bases j for relevant paramete. values, input information, and results from the subject j straight-leg tube vibration assessments. The objective of the assessments is to provide both the limiting circumferential crack angles for potential crack propagation associated with tube vibration and the associated crack leak rate time history. i For the analyses reported in this'section, it is assumed, consistent with the fiald data i review of Section 4.0, that (1) all steam generators (S/Gs) are in a similar condition, ! (2) there is denting at the tube / tube support plate intersections, (3) ODSCC can t potentially exist at the top- and/or bottom edge of each TSP with - the ! non-through-wall-cracked region of the tube cross sectiori at each of these cracks j having 50% wall thinning (see Figure 5-5), and that (4) PWSCC through wall cracks f can exist at the top of the tubesheet WEXTEX transitions. Thus, the scope of work [ reported in this section is to present for normal operation conditions (see Section 8.0 ! for steam line break (SL3) conditions evaluations), both fluidelastic and turbulence ! response evaluations to determine a vibration related limiting crack angle result f t associated with each of three crack conditions (top, bottom, and top and bottom of any j span) at various postulated crack angle conditions for each of the seven straight leg - spans on each of 3388 tubes in any S/G. For perspective, this amounts to considerably ; more than three million evaluations. An overview of the conceptual models used to -j complete 'this scope, as well as that of insuring that the leak before break concept is ! applicable to the S/G tube straight-leg evaluations, is provided in Table 7.1. l Included for each of the models (in Table 7.1) are purpose, assumptions, fluidelastic. j evaluation minimum reduced velocity (U/ID) stability criteria, damping inputs, and 'I description of intended use for results. More details associated with these models and j the basis for the leak before-break position are provided in Section 7.1. i Methodologies used to address both the fluidelastic and turbulence mechanisms of f t crossflow induced tube vibration are also described in the following section. It is noted ; that the mechanisms of axial flew turbulence and vortex shedding are not considered ! 1 vlable as major causative mechanisms based on field experiences and, hence, are not j addressed further herein. 7-1 i
7.1 Overview of Analysis. Methods and Models Evaluations for each -of' the mechanisms considered here for tube straight. leg fluidelastic and turbulence assessments are accomplished by starting with the same two Westinghouse proprietary computer programs which provide for the generation of a finite element model of the appropriate tube and tube support systems. [
']a,c. Mnre details associated with the separate fluidelastic and turbulence evaluations are addressed in Sections 7.1.1 and 7.1.2, respectively, in general, flow induced vibration-(FIV) evaluations are performed to determine the acceptability of tubes potendally subjected to both the fluidelastic and turbulence mechanisms (separately or combined) and with various boundary conditions, and crack sizes. These evaluations include effects of normal operation and load follow considerations as they apply to interim operation. The following paragraphs d!.scuss
~
specific tasks typically addressed for these evaluations. I
}a,c, Table 7.1 provides some salient details relative to the conceptual models implemented .
for these North Anna 1 specific evaluations. The first of these models is defined to be the " base case." The solution set obtained from this base case is defined to be set 1. The . set 1 solution is taken as the base or reference case and is comprised of a limiting angle and its associated mechanism for each tube. The set 1 co!ution completes the scopo defined above (see Section 7.0) using a given (reference) set of inputs (see Table 7.1). 7-2
.It is notably conservative to assume through wall cracks for these base case analyses since pulled tube results show short cracks separated by ligaments actually exist in the field. These ligaments would increase tube stiffness and decrease the likelihood of fluidelastic vibrations, and magnitude of turbulence vibration, compared to the case of a through wall single crack of large circumferential extent considered herein. Further, the-present evaluations are based on ignoring the considerable (relative to clamped conditions) damping associated with the tube cracks.
Tube damping plays the very important role of establishing- the tube vibration (and stress) magnitude for both the fluidelastic and turbulence mechanisms once all other parameters are established. Test programs have been completed at the Westinghouse S&T Center to measure mechanical tube damping for the relevant clamped tube conditions. One program dealt with the WEXTEX joint at the tubesheet while a second dealt with a representa;ive of more firmly clamped tubeshe.it conditions. This more firmly clamped representative condition is referred to in the following as the "hard" condition. These programs and the interpretation of the results from them are addressed in Sections 7.2 and 7.3, respectively. Recommended nominal tubo damping
^
for both the WEXTEX and hard clamped tubesheet conditions and for dented (clamped) conditions at the tube support plates (TSPs), in the form of explicit equations, are provided in Section 7.3. The base case model uses these damping representations. Solution set 1 from the base case model provides the conservatively obtained limiting fluideiastic and turbulence results. These results are made conservative by ignoring both the stiffness effects of the ligamerits and the damping effects from the cracks. It is judged that they are applicable to crack angles of at least 180 degrees, even though the data of Section 7.2 shows decreasing damping with increasing crack angle.
- Arguments supporting a leek-before. break position follow from the remaining three models on Table 7.1. This ligament case provides for the addition of the ligaments stiffoning effects to the base case and, thus, produces a more expected (and somewhat less conservative) result for the limiting tubes. #
The third conceptual model, the pre propagation case, is also turbulence limited and includes the effects of the additional damping introduced by the cracks. This additional damping due to the cracks themselves is quantified by test results in Section 7.2, 7-3
below. This concept is only used qualitatively in defining the expected sequence for propogation of a corrosion crack by vibration induced crack propagation. The final conceptual model, the propagation case, addresses the scenario of events after a relatively large rate of crack growth is initiated and long time periods of dynamic equilibrium, defined below, are no longer attainable. By taking these last three conceptual models in a time-line fashion and considering the . physical parameters and events associated with each we construct the following crack propagation scenario. First, the ligament model is limited by the turbulence mechanism and is stable in the short term. SCC growth is the major cause of low rate but continuous crack angle changes. The damping is essentially constant and small because of the steady state response. At some point in time an increment of crack propagation is expected. This will not only increase the crack angle but will also increase the damping in accordance with the test results for newly formed cracks reported in Section 7.2. This 1ew condition represents the pre-propagation state (model).' It too is limited in amplitude by turbulence. Based on the damping test results presented in Section 7.2 of this report, it is expected that, as time proceeds, the damping will decrease due to the wearing away of the newly opened crack surface anomalies and also due to some loss of the WEXTEX joint damping because of the fact that the joint is now not worked due to the
" hinge" developing at the crack (which changes the mode shape). This reduced damping will lead to higher amplitudes and more stress. The higher stresses over time will tend to increase the crack size again. The new crack surface anomalies would again introduce higher. damping which would, once again, stall the crack growth. Thus, the cycle is -
repeated and a condition of dynamic pseudo-equlibrium is reached whereby the crack growth rate can be expected to be quite low. If, at any: time, the mechanism establishing the limiting crack angle switches to . fluidelastic due to the combined effects of crack angle and damping, then the tube condition would be described by the propagation model, the fourth conceptual model on. . Table 7.1. However, for these evaluations, the fluidelastic limitirig angle is quite high relative to the turbulence limiting angle ([ ]b,c), 7-4
- __ . -_- ~ . .. ._- - . . . -
)
This ensures that a long time span _will exist wherein the tube is operating within the j
*- parameters and conditions associated with conceptual models 2 and/or 3. ;
- - Under these conditions, in the low probability event of a large tube crack developing and I progressing, as in going from the expected (ligament) case to a condition in which the ligaments are no longer existent, the tube will not undergo a rapid increase in I turbu!Once induced vibration amplitude that could potentially lead to a tube rupture in a !
time period less than that required for the normal and orderly shutdown of the plant. f That is, the loss of ligaments would create the conditions where there would be ample : i
- time for the crack to leak as it develops, such that the leakage monitoring system would i recognize the leak condition and cause the plant to be shut down. Thus, the ,
leak before-break scenario is appropriate for the North Anna 1 conditions. Normal Operation Evaluations : Fluideiastic Mechanism i P Using the ATHOS thermal-hydraulic computer program generated normal operating condition results, including appropriate secondary flow velocity and density distributims along the axis of the tube, and clamped support conditions, an evaluation is , performed to' determine if fluidelastic vibration is predicted for each of the spans ( > i -- i f I r I JaA Tubes : experiencing fluideiastic excitations vibrate at one or more of their fundamental frequencies with what may be large amplitude displacements. Large crack tip stress intensities could conceptually develop and this could result in rapid crack propagstion f leading to a severed tube if the condition were not detected in time to shut down prior to tube severing. $ Fluidelastic stability assessments are performed using the [ : Ja,c. Inputs to this code are [ ja,c l t 7-5 l r r
-[
ja.c, [
]b,c. An 3 valuation is completed for all tube spans and for all conditions of interest. These results can be presented in many different forms. Generally, it is instructive to produce maps showing the worst case boundary condition result at each tube location for each crack angle, span and tube. For the North Anna 1 WEXTEX applications, these maps show either the stability ratio (if it is greater than or equal to unity (1)) or the turbulence rms (root mean square) vibration amplitude (if the tube is stabla). Stability ratios greater than unity are considered unacceptable and, thus, this defines a limiting angle due to the fluidelastic mechanism for each span and tube.
Normal Operation Evaluations : Turbulence Mechanism For fluidelastically stable tubes, an evaluation is also performed to determine the independent turbulence response of the various spanc on each tube for the same (clamped) boundary conditions and crack sizes. [ Ja,c . 7-6
(
..... ja,c,
~*
"*eam Line Break Evaluations The method used to evaluate the effects of SLB is different from that used to determine the normal operation effects on the tubes and is comprised of a time history structural analysis of the tube and support system. These SLB assessments are addressed in
' Section 8.0.
7.1.1 Fluidelastic Analysis Methods The fluideiastic evaluations performed for the degraded tubes are to determine two relevant quantitles, stability ratio-(SR) and tube stress at the cruc.k locations (which correspond to the most highly stressed tube locations). SR (see below for mathematical definition) is a ratio of cffective velocity to the initiating velocity' associated with .j fluidelastic : vibration.
-)
Stress as a function of time with both mechanisms active msty also be determined. These stress time-histories are a result of the fluidelastic forces and drag forces simultaneously acting on the tube. [
]a,c, Typically, stability ratios are determined for specific configurations (flow and
' boundary conditions) of tubes. These stab;lity ratios represent,a measure of the potential for tube vibration due to fluidelastic forces during service. Conservatively, values greater than unity (1.0) are to be avoided since they imply some nonzero vibration amplitude. If the stability ratio is less than 1.0, then the tube would not experience fluideiastic excitation.
The evaluation methodology is comprised of determining values of the parameters contained in the following equations [ ja,c: 7-7
F .aldelastic stabitity ratio (SR) equals Uen/Uc for modo n, where Uc (initiating velocity) and Ugn (effectiva velocity) are determined by:
- a.c I
a s l . 4 ! 3 As can be seen from the above equations, the important input parameters are the stability constant (beta) and the damping values (: eta cr (J. The bases for these are presented below and in Section 7.3, respectively. 7-8
i
!! Is very important to note also that it has been recently recognized that the existing test data shows a lower limit to the instability velocity associated with a given tube bundle. This can be interpreted as a lower limit on the reduced critical velocity. Below this limiting reduced critical volccity value tubes remain stable, even for very low values of damping which would put the associated point in reduced velocity vs.
mass damping parameter space on the upper side of the extended line representing the following equation 7 3 (which represents equation 71 reformulated with the reduced -i velocity pararneter on the left hand side) as:
- ac (73)
More 'on this is provided in Section 7.2. The North Anna 1 fluidelastic evaluation results are based on a lower limit value for the reduced velocity parameter of 1.1 and are, therefore, consistent with the recommendations made in this regard. Fluidelastic Evaluation Parameters and Their Bases The beta and damping values used for these analyses are based on an extensive data base comprised of experimental results obtained at both Westinghouse and from other experimenters. In addition, previous field experiences are considered. The values used are provided wi h the results. The bases for the values used in these evaluations are provided next. Basis for the Fluidelastia Constam Beta The most complete summary of information relevant to establishing a beta value for use with fled issues is given in the [ ]c.c; f Based on the data contained in [ , Ja.c relevant to all Series 51 S/Gs. I ja.c, 7-9
.. ._ , . . _ - _ ._.~___.___._2___._..-. . _ _ _ _ _ _ _ _ . _
l I .
. i i
l i L ja ..c Thus, the value judged appropriate for the North Anna 1 WEXTEX , b ' cracked tube evaluations is { l,c, i Basis for Tube Damping Tube damping is generally dependent on frequency and support conditions as a minimum. The c!amped condition mechanical damping has been shown by tests to also depend on amplitude of vibration. For these evaluations, damping values obtained and/or
- developed in Section 7.3 are used.
l 7.1.2 Turbulence Analysis Methods The turbulence evaluations performed for the degraded tube cases are to determine tube !' stress at the crack locations which correspond _to the most -highly stressed tube locations. Stress as a function of time with both mechanisms active may elso be determined. These stress time histories are a result of the fluidelastic forces and drag forces simultaneously reting on:the tube. Comparing the values of stress to the values
. required for crack propagation will determine if the crack would be expected to grow 7-10 L
t
..--,e m , . . , y- - ,m , , r.; cy.,-y,_..,_,y',,..,_ . , , , - , ,. --
.-.%_ .,,, , , , - . . m--,,,,_,_,. . . . . . _ _ , , , _4 , _ , - - - - , .
l during, for example, a postulated SLB event. In this case, however, since the boundary
- conditions are all clamped due to either accombly (WEXTEX) or denting, there are no ambiguities such as changing boundary conditions caused by gap openings and closings and, therefore, no need for the combined mechar, ism evaluations.
Typically, stress,es are deterinined for specific configurations (flow and boundary conditions) of tubes. These stresses are input to crack growth evaluations for determination of acceptability. The evaluation methodology is comprised of determining values of the parameters contained in the following equations [ Ja,C. In particular, the value of the root mean square (rms) mcdal vibration amplitude is sought as:
., J . C O
7-11
i Turbulence Evaluation Parameters and Their Bases For the North Anna 1 evaluations, both Cy { ]b,c and s [ )D.C values were taken appropriate to cross flow turbulence force measurements. Knowledge of these constants implicitly establishes the appropriate turbulence driving forces which are consistent with the test measured force Power Spectral Densities. As with the fluidelastic mechanism, tube damping is generally dependent on frequency and support conditions as a minimum. The clamped condition mechanical damp!ng has been shown by tests to also depend on amplitude of vibration. For these evaluations damping values obtained andor developed in Section 7.3 are used. l The remaining physical and thermal hydraulic parameters are based on the Series 51 S/G design paramotors and the ATHOS code output. [ ja,c, 7.2 Damping for Dented Tube Conditions Tube mechanical damping is generally dependent on frequency and support conditions as a minimum. The fixed condition damping has been shown to also depend on amplitude of vibration. This section addresses the tests which form the bas 6s for most of the damping used in the straight-leg tube evaluat%ns. : 1 7.2.1 Damping Tests on Tubes with WEXTEX Tubesheet Joints 1 Mechanica! damping data are used to assess the fluidelastic vibration potential for tubes with circumferential cracks near the top of the tube sheet. In the WEXTEX plants, explosively expanded tube-to tube sheet joints are present at the top of the tube ,. sheet. In order to perform the evaluations, the tubes are assumed to be dented at the I first TSP (tube support-plate)! hence,- nominally fixed boundary conditions exist at q l-both ends of the span. A test apparatus was designed and built to obtain damping data for
- a single span tube with a simulated WEXTEX joint at the lower end, and a simulated dented TSP at the upper end. The apparatus is also used to obtain the damping of tubes 7-12 u__ ._ _ __ ___. ___ _ __ _ _
i f i with stress corrosion cracks, and through wall fatigue cracks generated at the SCC l
. location. l i . Test Aoparatus !
l The conceptual design of the test apparatus is shown in Figure 71, and a photograph of i the setup shown in Figure 7 2. [ l l i I i jb,c, f f i The damping can be measured by two differe.it methods: ; i (1) Pluck Test Method (logarithmic decrement method, rate of vibration amplitude } decay is a measure of the damping) : F i (2) Shaker Test Bandwidth Method (tube is driven at a constant vibration amplitude l rand the change in drive force as the drive frequency is varied is a rneasure of the i dampirig). L i The pluck test is best suited for linear or near linear conditions, for which the i damping is not a strong function of vibration amplitude. The shaker test is suitable for f linear conditions, and also for nonlinear conditions under which the damping may vary f significantly with vibration amplitude, Since the dampirg of a tube with a through wall [ t crack is nonlinear, the shaker test method gives the preferred results for this case. However, pluck tests are also run. They provide a useful check and confirmation of the f shaker test results. [ f i 7 13 : I _ . ~ _ . _ ~ _ _ . _ . , i
i- , i
" luck Test Method In the pluck test the damping of the tube vibration is measured using the logarithmic decrement method. Horo, [
b J,c, _ a,c 1 l l i 4-For ideal viscous damping, the logarithmic decrement is independent of vibration amplitude, but for practical cases, some variation of c'amping with vibration amplitude is present. This variation can be examined by measuring the damping at several different vibration amplitude levels along a test record. The damping generally i increases.' with vibration amp *, de for the simulated tube sheet joint and TSP l-7-14
conditions tested. For design calcelations, the appropriato damping for evaluating the potential for fluidolastic vibration can be associated with a vibration amplitude that is either expected to occur in any event duo to turbulenco, or is determined to be benign with respect to increasing tubo degradation. Shnker Test Bandwidth Method For this method the tube is driven by an [
]b,c. The porcent critical damping is calculated by
- 0iC e Test Soecimens WEXTEX Tube Sheet Simulation i
The WEXTEX tube sheet simulation shown in Figure 7 5, is [
}b,c 7-15
t t I f
]b,c, ,j Dented TSP Simulation .
'The simulated der.ted tube samples were made at the Westinghouse Science and Technology Center. Figure 7 6 shows a simulated 3/4 inch thick TSP specimen. The specimens are made of A 285 Grade C plate. The conter hole (0.898 inch diameter) in the simulated tube support plate is [
i 1
)b,c, Cracked Tube Simulation The generation of stress corrosion cracks in the specimens and the use of these
- specimens in the test program are the subject of Section 7.2.2.
Test Resultr Tubes Without Stress Corrosion Cracks The test matrix shown in Table 7.2 Identifies the main tests used to evaluate the damping. Other tests also run, but not shown in Table 7.2, involved additional checks and evaluations of the test apparatus to verify that test rig features did not introduce spurious damping due of the relatively low damping involved. 7 16 l, l. l
l l l Idealized Clamped Conditions - Reference Damping Tests 1-4 in Table 7 2 provide reference damping values for the apparatus. [ Ja,c, Two different span lengths between the top of the simulated tube sheet and the bottom of the simulated tube support plate (TSP) were considered,49.7 inches and 60.5 inches. I ja,b,c, Figure 7 7 shows the damping measured for the 60.5 inch tube length. [
]a,c. Data are shown for three cases, the original reference tube for which damping was measured in only the N (normal) directions, a second ,
reference tube for which damping was measured in both the N and T (transverse direction, 90' from the N direction) directions, and the spliced tube. The reference damping is quite low and is reasonably similar in the two perpendicular directions. No strong dependence of damping on vibration amplitude is evident. The average damping is equal to [ ]b,c, The spliced tube tests were conducted to verify that [ e ja.c, 7-17
] Figure 7 8 shows damping data for the 49.7 inch tube with ideally clamped conditions at both ends. The damping is relatively low and there is no evident trend of damping versus vibration amplitude. The average damping is [ ]b,c, WEXTEX Joint with an idealized Clamp at the TSP Location: 60.5 Inch tube. f l b.c Figure 7 9 shows the damping versus vlbration amplitude for an inillal deflection in the N direction (no test was conducted for the T Cirection) for W(1) (WEXTEX sample 1). The configuration corresponds to Test 5 in Table 7.2. [ ja.c. It was obsarved that the nattaal frequencies of the tubes with the WEXTEX Joints are slightly lower than the natural frequencies for ideally clamped tubes since the cylinder , used for the WEXTEX joints is more flexible than the support for the ideally clamped tubes. Figure 7-1.0 shows the damping versus vibration amplitude for WEXTEX sample W(3), for both the T and N directions. The configuration corresponds to Test 7 in Table 7.2. [ ja .. c 49.7-Inch Tube. I IiDO , Figure 711- shows damping versus vibration amplitude for an initial deflection in the - N direction for WEXTEX sample W(1). The configuratiran corresponds to Test 6 in Table 7.2. The damping tends to increase with vibration amplitude. The damping is 7-18 s- ,- % ~
, _ , . _ . . , ,. ,.,_..,y.,,_.-._.m, , , . r---q s,.,, % ,,,,.,,-y ,,,w _- -
l higher than the referenco caso values, and also generally higher than the damping for the 60.5. inch tube at a given vibration amplitudo. Figuro 7-12 shows damping data for WEXTEX samplo W(3). The configuration corresponds to Test 8 in Tablo 7.2. [
)b,c, WEXTEX Joint with a Donted TSP:
60.5-inch tube. I ]D . C-Figure 713 shows the damping versus vibration amplitudo for WEXTEX samplo W(3) with D(1) (donted TSP sample 1) for both the N and T directions. The configuration corresponds to Test 9 in Table 7.2. [
)b.c, in Figure 7-14 the damping in the N direction for WEXTEX sample W(3) with dented support D(1) is compared with the damping of WEXTEX samplo W(3) with an idealized clamp at the upper end. The results are quito similar, in Figure 7-15 a similar comparison is made for the T direction. [
]b,c, Figuro 7-16 shows the damping versus vibration amplitudo for WEXTEX sample W(1) with D(3) for both the N and T directions. The configuration corresponds to Tos.11 in Table 7.2. The samplo D(3) has 0.00 mils of diametral denting.
Figuie 7-17 summarizes the damping data from Figures 713 and 7-16 for the 60.5. inch, 42 H tube with a WEXTEX joint at the bottom and a dented TSP at the top.
]b,c, 7-19
49.7 Inch tube. I ta C Figure 718 shows the damping versus vibration amplitude for WEXTEX samole W(3) with D(1) for both the N and T directions. The configuration corresponds to Yest 10 in - Table 7.2. [
. )b,c. In Figures 719 and 7 20 the damping for WEXTEX sample W(3) with D;1) is compared with damping measured for WEXTEX sample W(3) and a clamped condition for the T and N directions respectively. [
jb.c, Figure.7 21- shows the damping versus vibration amplitude for WEXTEX sample W(1) and D(3) for both the N and T directions. The configuration corresponds to Test 12 in Table 7.2. ' The sample D(3) has 0.90 mits diametral denting. Figure 7 22 summarizes the damping data from Figures 7-18 and 7 21 for the 49.7 inch, 63 Hz tube with a WEXTEX joint at the top and a dented TSP as the bottom. [ jb,c, 7.2.2. Tests on Tubes with Stress Corrosion Cracks Mechanical damping measurements were made on three tubes with stress corrosion cracks (SCCs). The SCOs were generated on the tube specimens in high temperature autoclaves. The first specimen testad, NLS 12, whs obtained from a previous study of the stress corrosion response of- Alloy 600 tubing with explosive expansion tra isitions. The specimen was produced by [ 4 l ja.c 7-20
F =A 3A-542.4MS., 44 -*==.JL A4h meMmL&-'E#&5'-J . M-+J -nf 4 4 h T' b 7 I s i [ I e t t 1 5 i e F 6 t l. i h i I i e i 4 b s i s e
.t t
i b e F T h a f i t t 5 4 1
'I i
f P 5 5 4 f 3 f I Y r
' Y C
I i i s i I 4 [S L e f i 4 l i a
?
I ._
]a,C [
e i 7 21' ! ie
?
re,*.--9 ,etv*--'--r-- +-ee-+-*v =--rt- -*we' ' v w - N e em'
- v '* w w w n m ' v' vv v == v-wwt'*=M*-"' d +-e*w*ww-'-*-* '
- - ' ' - * * * - -++*-**C=*
* = '- - = ' ** - - ~r--'+*me'"e*=*-m*----r'-r~N*"r-
l ja,c, Overat! Summarv The following general observations are made on the basis of the completed test program:
- 1. The damping of a tune with an SCC crack (and an idealized tube sheet simulant and a dented TSP) is low, in the order of [ ]b,c, and is not approclably greator than the
[ ]b,c damping of a tube without a crack.
- 2. When the through wall crack exists, the damping bocomes [- ,
.)b,c, When the through wall crack is present in an operating steam ,
Osnorator tube, leakage flow through the crack will be detected by the leak rate monitor.
- 3. When the tube with a through wall crack is vibrated at a constant amplitudo, the damping falls off progressively with time for two of the three SCC samples tested. *;
I I
)b,c,
- 4. Sample SPI 22, for which the damping reduced [
jb,c,
- 5. The results indicate that the through wall crack damping can be an important consideration in analyzing the leak before break characteristics of tubes, it will bo 7-22
most significant for the [ c ]b,c,
- 6. The tube modal effectivo damping associated with the WEXTEX joint (or hard-rolled joint) is expected to (
n jb,c, Discutslon of Tott Resuht The Test Matrix for the cracked tubo damping tests is glvon in Table 7.3. Calibration Considorations Tests Tosts were conducted using a tube with no cracks in order to establish a basis for Interpreting the data for the crackod tubos. The damping for an uncracked tube with an idealized tube sheet and a dented TSP was measured in Tests 1, 2. The results are shown in Figure 7 23. The damping determined from pluck tests is nom;nally [ lb,c, The damping dotormlned from the shaker tests was higher, nominally [ ]bc. Accordingly for the reference tests (Tosts 817) described in the present section, and for the tests involving-Sample NLS 12 (Tests 3 7) the measured shaker test damping was reduced by [ ]b,c. Subsequent tests on SCO samplos SPl.L and SPI-22 showed good agreement betwoon damping determined by pluck tests and damping determined by shaker tests. Accordingly, no corrections are made in Tests 18-28. The correction applied in the Tosts 317 is believed to provide conservative results. No attempt was made to reproduce Tosts 1 and 2 to determine the source of the difference between the damping measured in the pluck tests and the shaker 7 23
~
. - . - - . - - . . . - - . - - . = . _ - - . _ . _ - _ = - _ _ -
tests. [ n ja.c, .: The damping was measured in Tests 8,9 for an uncracked tube with a WEXTEX tube shoot and a dented TSP. The damping determined using the bandwidth method Is about l ]b,c ta a 10 mil reak to peak amplitude as shown in Figuro 7 24. The damping measured in the pluck tests is higher, Figure 7 24 corresponds to the original WEXTEX configuration. Figure 7 25 shows data from Tests 14,15 that correspond to a moddication (see Result 3 below, the modification involven a hellarc wold repair to the tube) of the original configuration using the same WEXTEX joint. The modification Is not bolloved to have changed the relevant tube proporties, however it did involve removing the tube from the apparatus and reinsorting it back into the apparatus. Since there is some small clearance between the WEXTEX joint and the tube, some differences in damping values are expected when the setup is changed, since the tube can star 1 a different fitup in the WEXTEX jolnt. As shown in Figure 7 25 thero is good mont in the damping measured in the pluck tests and in the shaker tests. The og at 10
- mils peak to peak amplitude is about [ ]b,c, Tests were run on an uncracked tube (
Ja c. -The following results were obtained:
- 1. The damping at 10 mils p-p amplitude for a tube with a 90 degree slot (Tests 10,
- 11) is nominally [ ]b,c as determined in the shaker test. See Figure 7 26. The' shaded array in Figure 7 26 envelopes the damping for the uncracked tube as given in Figures 7-24 and 7-25. The 90 degree slot [
)b,c, t-7-24
- 2. The damping for a tube with a 180 degree slot (Tests 12,13), is nominally
,, [ jb.c (as compared with the nominal damping of [ )b,c for the idealized tubo shoot). Soo Figure 7 26. That is, the 100 degree slot essentially
, [ ja,c, The large chango in damp'ng between [
]b c. Soo Figuro 7 26.
- 4. [
]b,c. Those results can be used to estimate the combined damping for a WEXTEX joint with a cracked tube.
Test Results for Tubes With Stress Corrosion Cracks Damping infctmation was obtained for throo tubes with SCCs. The 49.7 inch tube is supported by the idealized tube shoot simulant at the lower end and a dented TSP at the upper ond. Samofo NLS-12 The first tests were run on Sample NLS 12 with a limiting peak to peak vibration amplitude of 10 mils for the shaker tests. The following results were obtainod:
- 1. The damping for tube NLS 12 with a SSC crack of [
)b,c 7-25
r i I
}bc, ,
- 2. Sample NLS 12 was removed from the shaker test apparatus and a through wall ,
crack was generated by conducting a fatigue test. [
]b,c. See Figure 7 28.
- 3. The damping for NLS 12 was measured over a period of time during which the tube ,
was vlbrated at 10 mils peak to peak amplitude (or pernaps as much as 15 mils ; peak to-peak during some intervals due to a malfunction of the controlling accelerometer). See Figure 7 29. [ ; f d
]b,c. A pluck test was then (1) The relationship between estimated crack angle and frequency as used here is preliminary. The intent is to illustrate trends.
7-26 r v r [^ 4 .r. , .,.....-.~r__,..
.r.._,_m ,, . . . ,,-_md._--,,,-- ,,-,.%...,..-- , , . . ~ , , - ,
I i conducted with ! jac. The resulting natural frequoney was ! L b ! l,c, i Samoto SPI-24 ' The second series of tests woro run on Sample SPI 24. It was specified that onco a
?
through wall crack is generated the vibration amplitude be limited to 5 mils peak to peak, sinco the 10 mil poak to-peak amplitudos (or perhaps largor) used for Samplo NLS 12 resultod in increased crack growth. The following results woro [ obtained:
- 1. The damping associated with a SCC crack (155 degroos, as determined from a plastic replication of the tube ID) was measured in Tests 18,19. Figure 7 30 shows the results. The pluck tests and the shaker tests give similar values for damping, hence no correction to the shaker test data is mado. The nominal damping is about [ ]b.c, as compared with the [ l,cb value reference value for no crack. [ -
)b.c,
- 2. A through wall crack with an angle of 164 degrees was generated. The damping was
- measured in Tests 20, 21. The damping at 5 mils peak to peak amplitude is ;
[ ]b,c, as shown in. Figuro 7-31 for timo equal to zero. Tho [ ' jb,c, e
- 3. The damping was measured in Test 22 over a period of timo during which the tubo was vibrated at a 5-mil peak to peak amplitude. The damping varlod as shown in Figure 7 31. The solid circles, plotted at the same timo as the open circles represent the damping measured at startup each day. Tho open circlos reprosent the damping i measured at the end of the previous day just beforo shutdown. While there is no oontinuous decrease in damping with timo evident, as was the caso for Sample NLS 12 (see Figure 7 29), .I does appear that the damping has started to fall off at 24 hours. 5 I
7-27 i r- ~
,,v- -
- + , . . , - ,e - , , - - . , - - . - . , - -- , , . _ . . - - - - - - , - - _ - - - . - , ~
r 4, A pluck test run following the 24 hour cycling period gave a damping value that is in good agreement with the damping obtained by the shaker test. See Figure 7 31. , Samolo SPl.22 - The last tests were run on Sample SPI 22. The following results were obtained:
- 1. The damping with the SCC (188 degrees, as determined from a plastic replication of the tube ID) was determined by pluck tests and also by shaker tests. Figure 7-32 shows the results. The shaker test damping was about [ ]b,c, which is similar to the lower values from the pluck tests. [
)b,c,
- 2. A through wall crack with an angle of 157 degrees was generated. The damping measured in Test 26 at a 5-mil peak to peak amplitude is about [ ]b,c for time equal zero, as shown in Figure 7 33, The pluck test damping, as measured in Test 25, is about [
)b,c, *
- 3. -The damping was monitored in Test 27 over a time interval during which the tube was driven at a 5 mil peak to peak vibration amplitude. The damping varied as shown in Figure 7-33. The solid circles, plotted at the same running time as the open circles, represent the damping measured at the startup each day, while the open circles represent the damping measured at shutdown on the previous day. There is a continuous
[
]b,c. After 24 hours the damping had [ ]b,c, suggesting that little if any through wall, crack related damping is being generated. This suggests that the asperities in the crack region have been worn away or permanently defolmed to the point where metal-to-metal contact is either not occurring, or is very slight.
- 4. The vibration amplitude was then increased to eight mils peak to-peak in Test 28.
The damping [
]b,c, 7-28 l
t
- 5. The observations in lloms 3 and 4 above suggest that as the crack rolated damping [
. l l r
}b.c, 7.3 Interprotations and Rocommended Tubo Damping The purpose of this Section is to provide the basis for the values of damping (( [zota), j
% of critical) used to datormine the limiting acceptablo crack angles associated with tho i I
two vibration mechanisms and the clampod support conditions for the North Anna Unit 1 F i WEXTEX tubos in addition, interpretations of cracked tubo damping data relevant to ; estab!!sning expected adequate timo to doloct a loaking tubo condition and shut down the plant before the leak becomes largo aro provided based on the damping test databases ; presented in Soction 7.2. b The damping values derivod horo for use in thoso evaluations are based on the test results from Section 7.2 which compriso the basis for mechanical damping for thoso clampod boundary conditions, with and without tubo cracks. Specific de Ting tests woro porformed for- the WEXTEX supported (1st tubo span) casos, wv..o ( Ja c is used to establish the viscous damping componont for clampod boundary conditions. Note that similar data form the basis for the damping recommended fpr uso in the evaluations without WEXTEX joints, such as is appropriato to North Anna 1 for the second span of the tubos above the tuboshoot All those damping recommendations represent good faith efforts to address a subject which, by its very nature and the nature of the methods used for measurements, is complex at bost, and is subject to the it'fluonce of statistica:ly distributed paramotors and measutomont errors. The damping equations developed are judged to bo consistent with engineering experiences as well as tho data of l Ja,c in addition to data from many other Westinghouse test programs and from the literature. The method developed for the purpose of interproting the test data amounts to obtaining the mean damping as a function of vibration amplitudo and frequency using linear regression techniques. Thoso developed regrossion models for clamped tubo damping provido the basis for thoso VRA ovaluations to put potential dented tubo field issues into bottor perspectivo. 7-29 -
For the purpose of providing these damping recommendations, and consistent with [ ja,c, damping is considered to be attributable to two sources. These are - referred to as the bulk fluid viscous and, in the case of pinned boundary conditions, the , squeeze film term. In the case of clamped (fixed) boundary conditions (appropriate to 4 : thethe North Anna 1 evaluations \ the squeeze film term becomes the mechanical { damping component. Total clamped bound, y conditions damping is taken as the sum of l the two components, viscous and mechanicat The recommendation for viscous damping for clamped boundary conditions is based on [ ja.c as the first term of equation 29 of that paper. Evaluated exp!!citly for Series 51 conditions, the recommendation, in terms of percent of critical damping, [ is: _ 5.C t 7.3.1 Mechanical Damping for First Tube Span (WEXTEX) Conditions e" The recommended WEXTEX mechanical damping expressions, based on all the foregoing considerations and the data in Section 7.2, are depicted in Figures 7-35 and 7 36. . i Based on the data It is observed that: a b.c _1 Explicit recommendations, in the form of two equations, giving all information necessary for WEXTEX clamped condition mechanical damping predictions for the mean - as a function of amplitude and frequency, and consistent with the above observations, for use in the North Anna 1 evaluations consistent with the models of Table 7.1 follow. 7-30
= . . .. . _ . . - - - _ _ . _ . _ - _ . _ -. _. . _. . _ - - . -
IO a.c
~
7.3.2 Mechanical Damping for Upper Tube Spans with Dented Support (Non WEXTEX) Conditions Tha recommended dented supports condition mechanical damping expressions, based on all tht foregoing considerations, are depicted in Figures 7 37 and 7-38. The data on which this recommendation is made were obtained from tubes with a lighter mechanical fit and considerably less (conservative) damping than that associated with the WEXTEX conditions. Thus, it is judged appropriate that the damping projections associated with these data . be applied to the non WEXTEX conditions in the North Anna 1 steam generators.11 is noted that the regression fits obtained from working with this specific test database tend to have a slight
- belly" in the dampir g vs. amplitude curves at lower amplitudes and frequencies (see Figure 7 37). Based on experience, judgement and
! context, this " belly", believed to be physically anomalous, was removed - with I mathematical techniques which replace it with a minimum length straight line, which has a very small but positive slope, for the North Anna 1 evaluations. 7-31
. .. . - _. ~. . - = . - = . _ . -. _ __ .. - - - _ . - . _ Pased on tF 'Jures it is observed that: the actual data ranged from 3.8 to 98.5 mils pk pk. .; the F*T9 regressio'n modelis recommended throughout the range of the data for predictions and extrapolations. . Explicit recomntendations, in the form of a single equation giving all information - necessary for clamped (dented) condition mechanical damping predictions for the mean as a function of amplitudo and frequoney, and consistent with the above observations, for use in the North Anna 1 evaluations consistent with the models of Table 7.1 follow. The definitions of parameters olven in Section 7.3.1, abovo, are entirely applicable here too. a,C 7.4~ Limiting U//nD for Predicting the Initiation of Fluidelastic Vibration in Tube Arrays in Water ~ Fluidolastic vibration is characterized by s threshold velocity, below which the vibration amplitude is caused primarily by turbulence and is relatively small, aad above which the amplitude increases rapidly. The threshold or critical velocity Uc is commonly given by equation 71 that was first developed at Westinghouse [ ja,c,
. Figure 7 39 shows a typleal stability ' diagram [ ]a,c. Instability is predicted when the effective velocity, Uen, is greator than the critical velocity. The effective velocity h wenly given by[ la,c equation 7-2.
As noted by [ ja,c, ' 7-32 __ _ _ _ ~ . _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ . . . _ _ _ _ _
A first estimate can be made [ ja.c of the range over which the
, quasi-steady method can bo expected to be most applicable. The quasi steady approach essentially [
jac a,c (7-7) I ja.c, Subsequently, other investigators, starting with [ n jac 7-33 ,
._ _ , _ _ _ _ . _ , . _ _ _ . . ._2.._..
1 { l ja.c. _. a.c (78) I i T i h 4 Ja,c.- [ In a very recent paper, [ v ja.c.
.[ .
I i-ja.c 7 34 n. I
i i l l ! l i t t i i i 6 t 1 t
-Ja,b,c, i
- ,i
- Table 7.4 summarizes some pertinent correlations that have been recommended in the literature for predicting the _ onset of fluideiastic vibration for tube arrays in the liquid l region. The results are plotted in Figure 7 42. [ Ja,c 7 35
)- ja,c. For completeness, some of the individual correlations gWen in Table 7.4 will be discussed in more detall. The very recent: paper by [ _
.g.
Y a% ~ A la,c , 7-36
- ,; 'O l
!l l r ja.c, I.
.. ja.c,
- There is also other informaFan in the literature concerned with the fluidelastic vibration of a single t.'exible tube among rigid ns ors. Although this type of configuration introduces other interpretation considerations, recent results reported by [
ja,c, Further work is needed to better characterize the fluidelastic vibration of tube arrays in water. However, there is large body of information provided by active investigators in the field of flow-induced vibration of heat exchanger tubes that supports the increa::ed stability of tube arrays in the liquid region, over that predicted using equation 7-1, which is similar to the t Ja.c design guide. The overall review of the available data is believed to support [ Ja,b,c, 7-37
7.5 : Results from Normal Operation Straight Leg Tube Vibration Evaluations The'inillal step in the normal operation straight-leg evaluations was to exercise the Base Case model (see Table 7.1) using a reference set of input parameter values, [ - ja,c, These tube bundle maps present the stability ratio result if a particular tube was determined to be unstable and the rms turbulence amplitude if the tube was stable, in accordance with the scope defined in Section 7.0, maps were obtained for sach of 3 crack conditions (a single crack near the bottom of the span either at the tubesheet or the lower TSP, a single crack near the top of the span, and, thirdly, a crack of equal size i . both the top and bottom of the same span) for each of 7 straight-leg spans on each of 3388 tubes, These preliminary results were used to determine which regions of the tube bundle were potentially susceptible to flow induced vibration concerns and which we: " obviously not susceptible. Specifically, these runs demonstrated the following:
- a,c Based on the above results the final evaluations, which were performed with the appropriate parameter values, were limited to the first and second tube spans above the tubesheet.
i 7-38
1 i Several examples of these output tube bundle maps from those final evaluations are
, provided by Tables 7.5 through 7,7. These tables are all for the condition of a single crack at the bottom of the first span and, of course, deal with the hot leg Each reports an appropriate allowable turbulence induced strose with and without the SCC crack effect. Each shows results for each tube in the quadrant defined by columns 1 through 47 for specific rows (namely, rows 1, 6,11,16, 21, 26, 31, 36, 41 and 46). This, of course, is a mirror image of the other half of the tube bundle (as noted). An *R*
indicates there is no tube at that location. Numbers indicate turbulence stress levels. Numbers with stars (*) Indicate unacceptable turbulence stress levels. A series of stars (*"""*) indicates a flubelastically unstable tube. The purposes for presenting these particular maps as thu examples follow. Table 7.5 was generated for results that consider a total crack angle of I
]b,c, Table 7.6 was generated for a total crack angle of [ ]b,c. This represents the smallest angle (again, for these specific conditions) with tubes which are unacceptable based on the fact that they were evaluated as being potentially fluidelastically unstable. It is again noted that the tubes with stress sufficient for crack propagation are at the periphery of the tube bundle.
Table 7.7 was generated for a total crack angle of [ ]LA These results show where in the tube bundle the tubes tend to be limited by each of the two mechanisms (instability or turbulence). The high stresses are again at the periphery of the tube
~
bundle. It is very important to note that the results on the tables just discussed show that the tubes susceptible to either of the vibration mechanisms are at the periphery of the bundle. This condition holds independent of the value of the crack angle being 7-39 I
. .= - .__
considered.- it is further observed that since these tubes are at the periphery, they are outside of the tube bundle region with the sludge pile. Based on the data in Section 4.0, - since these tubes are outside the sludge pile,it would be an extremely unlikely event that these peripheral tubes would have cracks at the tubesheet. - Turbulence Limiting Crack Angles Umiting crack angles were determined based on the methodology describec 4 :ove in this Section, The important results are pesented below in Table 7.8. Fluidelastic Umiting Crack Angles The subject fluidelastic evaluations, performed with a beta value (see Section 7.1.1) of [ la,c and a reduced velocity (U//D) value of [ la,c, provide the limiting angles (largest angles without instability) given in Table 7,9 . 7.6 Influence of Ligaments on T*/be Vibration and Crack Propagation The results of Section 7.5 above as given in Tables 7 8 and 7 9 apply to unllorm
~
through wall circumferential cracks, in corrosion degradation, either neatly through wall ligaments or wall thickness ligaments are expected to be present in the EOC cracks. The segmented crack model is applicable to PWSCC cracks in the WEXTEX transluin as described in Section- 3 based on pulled tube exams. This model represents the macrocrack as [ jac, Ugament Effects on Fatigue Crack Propagation The threshold through wall circumferential crack size for the onset of fatigue crack propagation due to turbulence loading is approximately [ Ja.c as shown in , Table 7-8. For partial wall through circumferential cracks, a [ Jac crack is needed and it must be over [ ]a,c of the wall thickness in depth. In fact, elastic calculations show that the remaining wall thickness needed to prevent the onset of 7-40 1
h fatiguo crack propagation due to turbulence loading is [ i
. )b,c, If a ligament exists betwoon coplanar through wcil circumferontial cracks the appiled stress intensity at the far ends of the crack array is much [ !
I t
?
i
)b,c, I
f
]b,c, Ligament Effects on Fluidolastic Instability of Tubes with Circumferential Cracks The through wall circumferential crack sizo nooded for tho onset of fluidolastic instability is about [ ja,c (Tablo 7 9). A bounding analysis was performed on ligaments offects on the bending stiffness of tubos with circumferential cracks. The ' -;
presence of oven a single ligament has a large offect on the bonding stiffness, compared ! to a tubo with a single through wall crack of the shmo total longth as the segmented crack. It was found that- ligament configurations which romain essentially [ j I b Jc ; 7-41 !
I
]b,c, The above results Indicate that corrosion cracks with ligaments typical of the segmented crack model and or with wall thickness ligaments on the order of [ ]b,c would not result in crack propagation due to turbulence or fluidelastic vibration. Thus the potential for vibration induced propagation of corrosion cracks is limited to essentially
[ ]b,c, , 7.7 Crack Growth Rates Associated with Vibration Induced Crack Propagation This Section looks at the subject crack growth in order to obtain a better view onto ~ leak-before-break and to establish cxpected times to detect a leaking tube condition and shut down the plant before the leak bocomes large. From Tables 7.8 and 7.9, the limiting turbulence and fluidelastic angles are obtained. The intent here is to use these angles along with the conceptual models of Table 7.1 to define conditions for computing conservative (small) time periods associated with the process of the crack growing through a transition from the start of the turbulence driven crack propagation case to the start of the fluidolastic driven propagation case and/or from the start to the finish of the fluidelastic propagation case. 2 Here, the raethods are based on the application of crack growth cdculations consistent with linear fracture mechanics, in addition, for the fluidelastic case, the tube vibration amplitude is assumed to be dependent on_ the iluidelastic stability ratio to some known power (itself a function of amplitude). i: is important to note that these assumptions , were used in the resolution of the original North Anna R9C51 U-bend high cycle, limited displacemt nt fatigue failure. In these evaluations however, the stability ratio , is based on the minimum [ ]a,b,c consistent with Section 7.4. 7-42
l I It is notably conservative that, for the turbulence case, as well as in defining the
, minimum initiation fluidolastic amplitude, [
ja.c, I jb.c, 7.7.1 Peripheral Tube with a Crack at the WEXTEX Transition For the limiting tube with a crack at the WEXTEX transition we start at the initiation of 4 propagation with the conditions of a stable tubo driven by turbulence and a crack angle of[ }a,c (see Table 7.8), The end of the turbulence driven propagation 4 occurs when we reach the crack angle associated with fluidolastic instability, which is, from Table 7.9, equal to [ ]bc. Application of the Iterative procedures described above gives [ ]b,c as the limiting total growth time during the turbulence propagation phase. The crack angle vs time plot showing these results is provided on Figure 7-44. Once the crack has reached the [ ]b,c a fluidelastic propagation caso is initiated. Applying the above techniques to these conditions, where crack growth is driven by fluideiastic instability, leads to an additional fluidelastic driven growth time to tube rupture of [ )b,c (also shown on Figure 7 44). If it is assumed that crack propagation is initiated only by fiuidelastic tube vibration, without prior turbulence driven crack propagation, the time to grow from a [
)b,c crack to a [ ]b,c crack is about [ ]b,c. Figure 7-45 shows the crack angle versus time for this case.
7-43
- _ - _ _ - _ _ - _ _ _ - _ - _ - _ _ _ _ - _ _ - - - - - - - _ - - _ - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - ~
o 7.7.2 Peripheral Tube with a Crack at the Top of the First Span , Similar evaluations were made for the case with a crack just below the first TSP and with 50% wall thinning. Here, the turbulence limiting angle is [ ]b,c (see . Table _7.8). The associated fluideiastic limiting angle is greater than [ ]b,c, The total crack growth time associated with this turbulence propagation phass was determhed to be greater than [ ]b,c Figure 7 46 shows the crack angle versus time for this case. 7.8 Leak Rate Versus Time for Vibration Induced Crack Propagation To assess operational times avalla,ble to respond to vibration induced crack growth, the crack growth rates obtained in Section 7.7 are used to define leak rate versus time in this Section. The general trends for leakage are discussed based on considerations of the ligaments present in the cracks. The specific times for vibration amplitudes to braak the remaining ligaments and the associated leak rates are not currently predictable and would be strongly dependent on the details of the ligament structure. The leak rate versus time analyses are therefore performed for the time frame after the ligaments ' are assumed to be lost by inillal vibration or corrosion so as to result in a uniform through wall crack that centinues to propagate as a result of tube vibrailon. Based on gensial considerations of leak rates for ligament models and the influence of loss of !igaments on tube damping, the following qualitative sequence of crack propagation and leakage can be postulated:
- 1) Crack angle,0, developed by corroslor.
- 2) Continued corrosion thins the ligaments to the point wherecrack tip stress from vibration amplitudes exceed the threshold for crack propagation in the ligaments. The stresses are expected to be largest in the ligaments due to local ;, tress concentration -
such that ligaments would crack before the total crack angle increases due to vibration. Based on the above analysis aesulls, initial vibration amplitudes to break the ligaments wofd likely be due to turbulence induced vibration.
- 3) As a few ligaments are lost, leakage wou!J increase as the uniformly through wall crack length incretases. The leakage trend would likely be a series of small step 7-44
increases as individual ligaments are lost. With vibration amplitudes sufficient to break ligaments in the corrosion crack, even normally tight corrosion cracks would be expected to have leak rates typical of fatugue cracks which have been used to develop leak rate versus crack angle correlations. Loss of only a few ligaments would result in 60 - 90 degree through wall cracks with leak rates greater than 50 gpd and lead to initiation of plant shutdown.
- 4) As ligaments are lost, tu'oe darrping increases as shown by the test results of Section 7.2. The tests show that after fatigue loadings are app!!ed to break the ligaments, tube damping increases significantly due to added damping from the crack face irregularities.
This added damping obtained as ligaments are lost would tend to reduce vibration amplitudes and slow the progression of the crack through the remaining ligaments. The tests also show that continued vibration leads to a loss of the crack damping over a period of about 24 hours. Thus, loss of ligaments should occur over a period of time which could vary from many hours to days, depending upon stress amplitude. The time from initial loss of ligaments with resulting leakage to initiate plant shutdown and the time that all ligaments are lost can reasonably be expected to be adequate to shutdown the plant before propagation of the corrosion crack angle,0, occurs.
- 5) After the ligaments are lost, further crack propagation would lead to increases in the initial corrosion crack angle, O. At the point of propagation of the corrosion crack angle, leak rates would be high and plant shutdown initiated or completed.
Based on the above discussion, it is conservative to evaluate the time available to shutdown the plant for assumed vibration induced tube crack propagation based on the crack growth rate for the through wall crack angle required to initiate crack propagaton as developed in Section 7.7, above. Leak rate versus time is calculated for the limiting crack angle of [ }b,c for turbulence induced crack growth and [ ]b,c for fluidelastic vibration Induced crack growth. It can be noted that the known cases of fatigue crack propagation (N)rth Anna 1,1987; Mihama 2,1991; Indian Point 3,1988) have been totally fatigue induced between crack initiation and the final termination crack angle at plant shutdown. There are no known occurrences of a corrosion crack of significant extent being propagated by fatigue to assess leak rate trends under these conditions for comparisons with the analytical models. 7-45
Figure 7 47 shows the WEXTEX leak rate versus time for the case with turbulence driven crack propagation initiated at [ la,c degrees and fluidelsstic driven propa0ation after a [ la,c degree crack angle is reached. The Initial leak rate for a [ la,c through wall crack with no ligaments is expected to be about [ -
}a,c. The leak rate increases slowly with time [
la,c such that ample time exists for leakage detection and plant shutdown. A leak rate of about [ ja c wou!d be expected in about [ ]a,c. With the North Anna 1 leakage monitoring procedures, plant shutdown would be completed before the [ ]a,c leak rate would be attained.
' The assumption of only fluidolastic vibration causing crack propagation represents a somewhat more limiting challengt e the leakage monitoring system than turbulence induced vibration, ' Figure 7-48 shows the leak rate versus time for fluidelastic vibration induced crack propagation initiated at [ l b,c is expected to be about [ ]b,c, such that plant shutdown would likely be initiated by the 50 gpd administrative limit prior to reaching this through wall crack angle. The leak rate would increase to about [ l b ,c. Initially, the rate of leakage rate increase is about[ }b,c and increases to about '
[ ]b,c. Plant shutdown should be completed within
-about 7 hours to prevent a 360 degree tube upture. For the North Anna 1 leakage monitoring procedures, plant shutdown would be achieved in about a 3 hour period after exceeding the administrative leak rate limit of 50 gpd. Thus, shutdown would be achieved prior to a 360 degree tube rupture even if fluideiastic instability causes the crack propagation.
l 7.9 Summary and Conclusions This Section provides the detailed methods and test and analysis results from a comolete - straight-leg normal operation FlV evaluation of the North Anna 1 steam generator tube bundle in its present condition. It considers all relevant tube and support conditions and potentially active vibration mechanisms, s L The detailed analyses were performed for uniform through wall crack angles to estimate the crack angle potential 4 resulting in crack propagation. Comparative analyses of tube stiffness and crack tip stress intensities between cracks with ligaments and through wall cracks indicate that [ ]b,c 7-46
propagation even for macrocracks up to 360*, The segmented crack model with elastic , ligaments as used to represent PWSCC cracks and applied for tube burst modeling in Section 6 would permit 360' macrocracks without crack propagation. For ODSCC, it is expected that remaining wall thickness ligaments would be beneficial in reducing the potential for crack propagation. Limiting through wall angles for crack propagation were determined for every straight leg span and tube in the steam generator. For WEXTEX transitloon circumferential cracks, the limiting through wall crack angles are [ ]5.C for turbulence and [ )b,c for fluidelastic excitation and are located in the peripheral tube region of the steam generator. From Table 5-9, it is seen that the associated probabilities for these through wall crack angles are about [ }b,c, respectively. For WEXTEX cracks in the central sludge region, the limitir.g crack angles increase to [
]b,c for fluidelastic vibration and [ }b,c for turbulence excitation. The limiting crack angles are reduced for combined circumferential cracks at the WEXTEX transition and at the bottom of the 1st TSP, However, the probability for occurrence of these combined crack c onditions is
[ ]bc, For circumferential cracks at TSP edges, the limiting through wall crack angles for propagation are [ ]b,c for turbulence and [
}b,c for fluidelastic vibration. No cases were found that resulted in fluidelastic instability for TSP circumferential cracks up to the largest angle of
[ ]b,c used in the analyses. This results from the fact that crossflow velocities in the inlet pass are higt near the tubesheet and low at the bottom edge of the 1st TSP. Crack propagation rate analyses were performed for the limiting through wall crack angles to obtain crack angle and leak rate versus time. These results provide a basis for evaluating the North Anna leakage monitoring system and procedures. Turbulent driven crack propagation proceeds slowly and can readily be detected by leakage monitoring in ample time to implement plant shutdown if the 50 gpd limit is exceeded. For
~
fluidelastic driven crack propagation initiated at a [ ]b,c crack angle, it takes about [ }b c to proceed to a tube rupture. Since North Anna provides continuous leakage monitoring and shutdown capability within 2 hours of exceeding the 50 gpd admini trative leak limit, the fluidelastic driven crack propagation would also be detected and the plant shutdown prior to a tube rupture. 7-47
]
Overall, the vibration analysis results indicate that remaining ligaments in corrosion cracks are likely to prevent crack propagation due to tube vibration. This is . particularly true for WEXTEX PWSCC cracks for which ligaments are found between the microcracks in pulled tube examinations. Thus WEXTEX crack propagation is unlikely - due to crack morphology considerations. Fluidelastic instability is not expected for circumferential cracks at TSPs due to flow distribution considerations, if large through wall cracks and associated crack propagation should occur (< 1% probability of 1 tube in 3 steam generators at EOC), the most likely condition would be turbulence driven propagation of a circumferential crack at the bottom of the 1st TSP. This would result in a slowly increasing leak rate and plant shutdown without a large leakage event. The North Anna leakage monitoring system is capable of detecting leakage and shutting down the p; ant prior to a tube rupture even assuming fluidelastic driven crack propagation. I 7.10 References _ aC l t l l l m M 7-48
- - ~ - - - - - , - - - - , - - - - , , , _ , _ , , _ _ _
E M _ BeC . 1 h f
- 4 6
k 9 h r
')
l m l l t i y 8 t n E h 5 e L t' e F t G t a e f m same 1 7-49
9 __
% mM g E
' 8, C f
e w 9 8 4 h eeen 7 50
- e - + -
3 1 m TABLE 7.1 : VRA OPERATING CYCLE ST-LEG VIBRATION EVALUATIONS OVERVIEW h4ODEL PURf0SE ' CRITERIA / INPUT DAMPING HESULTS full 0 ELASTIC TUR0LENCE i _= a, C :' t I m m I. N b i emus. ____. . - _ . . _ ~ , - - - - - - - - - - -
> - - - + - - - mum am
' = - - - - - - - - - k 7-51
+ . 3 1 ' . , $ w w ~ .+< x
Table 7.2 Tube Damping Tests a,c
~
p;l D(1)' = 3.80 mils diametral denting D(3) = 0.90 mils diametral denting C 9 1 7-52
. _- _ __._________________________-_L-__-_---_---
o . Table ' 7.3 - Test Matrix for Investigating Damping of Tubes with Stress Corrosion Cracks Test 1 2 3 4 5 6 7 8 9
- - - a,c Configuration Tube Ssmple Tube Sheet ._
, TSP Crack Angle (SCC) Crack Angle (Through-wall) - Pluck Test Shaker Test mils n-p is Hz 24 Hour Test mils p-p _ NC = No Crack a. 180* total SCC indication, 84* significant depth indication SCC = Stress Corrosion Crack b. Total fatigue crack 138*, ligament at about 120*; hence, crack SCCP = SCC with Fatigue Crack not continuous NCS = No Crack, Slot Added c. Machined slot I( ) = Idealized Tube Sheet d. Test not run W( ) = WEITEX Tube Sheet e. Weld repair D( ) = Dented .ISP 7-53
- -e . .n
{.
- . - . . . - .. ..- ~ .-- .
Table 7.3 - Test Matrix for Investigating Damping of Tubes with Stresa Corrosion Cracks (Cont'd) d Test 10 11 12 13 14 15 .16 17 18 19
- a,c Configuration Tube Sample Tube Sheet TSP Crack Angle (SCC) _
Crack Angle (Through-wall) _ Pluck Test _
, Shaker Test mils p-p fy Ils ,
24 Ilour Test mils p-p _ NC = No Crack a. 180* total SCC indication, 84* significant depth indication SCC = Stress Carrosion Crack b. Total fatigue crack 138*, ligament at about 120*; hence, crack SCCF = SCC with Fatigue Crack not continuous i NCS = No Crack, Slot Added c. Machined slot
- d. Test not run I( ) = Idealized Tube Sheet e. Weld repair W( ) = WEXTEX Tube Sheet D( ) = Dented TSP 7-54
Table 7.3 -- Test Matrix for Investigating Damping of Tubes with Stress Corrosion Cracks (Cont'd)l Test 20 21 22. 23 24 25 26 27 28 a ,c Configuration Tube Sample Tube Sheet TSP Crack Angle (SCC) Crack Angle (Through-wall) Pluck Test Shaker Test mils p-p is Hs 24 Hour Test mils p-p NC = No crack a.180* total SCC indication, 84* significant depth indication SCC = Stress Corrosion' Crack b. Total fatigue crack 138*, ligsment at about 120*; hence, crack SCCF = 300 with Fatigue Crack not continuous NCS = No Crack, Slot Added c. Machined slot I( ) = Idealized Tube Sheet d. Test not run
- c. Weld repair W( ) = WEXTEX Tube Sheet D( ) = Dented TSP 7-55
Table 7.4 Summary of Recommendations in the Literature for Predicting the Initiation of Fluidelastic Vibration in Square-Pitch Tube Arrays in Water - a ,c .. O D l l 7-56 l-l
+:
- Table 7.5 a,b,C
' \ e m #AAA4 O 9 7-57
Table 7.6 . mm.m a,b.c G w n (- 4pene
'4 r
4 7-58
A t. P 3 i Table 7.7 9
'$ y b
Y
. a,b c .,
- i
~
b i i i
?
e
- I P
?
t 9
?
'h h
e e
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T F
'h a
h
-1 i
4 6 I
- 5 1
F I
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)
p i 7 59 ;
}
?
s
Table 7.8 Limiting Throughwall Crack for Turbulence Vibration Crack Propagation . Limiting Turbulence Total Span Crack Crack Angle (degrees) of Cracks Number Location (s) no thinnino 50'4 thinnina b,c 1 1 WEXTEX (Bottom) 1 1 WEXTEX (Central Region) 1 1 Below TSP 1 (Top) 1 2 WEXTEX and Below TSP 1 2 1 Above TSP 1 (Bottom) 2 1 Below TSP 2 (Top) 2 2 Above TSP 1 (Bottom) and Below TSP 2 (Top) - - f h e Q 7-Eo
Westinghouse Proprietary Class 2 Table 7.9 Straight-Leg Normal Operation Fluidelastic Limiting Throughwall Angles for Fluideiastic Vibration Crack Propagation Limiting Turbulence Total Span Crack Crack Angle (degrees) of Cracks Number Location (s) no thinnina 50% thinnina b,c 1 1 WEXTEX (bottom) 1 1 WEXTEX (Central Region) 1 1 cielow 1st TSP (top) 1 2 WEXTEX and 1st TSP
- 2 1 or 2 Any of Above l -.
^ e: .6 O
7-61
J U e 40 '
- l. ;
s M V W k e6 Q. Q. M 4 m G G C C . . .
..t.4
$1
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r0 9 B N 6
- t4
.to
.e Q
~
l N G h 3
.t4
.o 0
4 I I 7 62 4
8C e
, #*L Figure 7-2 -
Photograph of damping test apparatus 7-63
a.c ($s
~
;0 t
r> Figure 7 Split-ring clamp used te provide ideally-clamped end conditions 7 G4
_ = - _ . . . . _ . - - . - ~ . . . - --- . - - - _ _ _ __- --.-. .- . _ _. i e f
~~
y - i ; e.imen . i 1 q _._ _ ' T . I 1
, - i i i , t,
- i ___
(a) Initial deflection in the H direction. Relatively i i little vibration buildup in the T direction (tube with
~
clamped end conditions) 1 y _.
..m..
' i Hi;tlH h{Mhbh$NNNNNMM$hi5${ E,l ;: "'" I i ,,.;,..; won i o n h g ggg!gi j;113nuNdbo" ._ _ ;
'Sih ._ . ._ . _ _
t
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. Figure 7-4 - Typical oscillograph records showing maplitude decay i traces in both the normal, N, and transverse, T, directions ,
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a bec ! p - n ' P i l e l-Figure 7-7 - Damping versus vibration amplitude for the 60.5-inch, [
]Vtube ideally-clam Data are shown for two reference tubes, ped and at each also for end.
one of the reference tubes that was cut and then spliced (welded) together again. 7-68
i i a,b.c i
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I Figure 7-8 - Dampingversusvibrationamplitudeforthe49.7-inch.[
't bu e ideally-clamped at each end. Damping shown is fo]r the N direction i 7-69 i
L
-y- ,,m, -w,,. y v. - - - . - w,-e, , ,
a.b.c M Figure 7-9 -- Damping versus vibration amplitude for the 60.5-inch,[
'f%ube with WEXTKI joint W(1) at the locer ond, and an ideally-clasped condition at the upper end. Also shown are daaping acasurements obtained after water was drawn -
into the-. joint by capillary action. Da:nping shown is for the N direction 7 70
]
[ a,b.c 4 i L t i i F { t uma Figure 7 Das ing versus vibration amplitude for the 60.5-inch,{ ube with wmn joint W(3) at the lower end, and an s, i eally-clamped condition at the upper end. Damping is shown for both the N and T directions > 7-71
a,b.c Figure 7-11
-Dampingversusvibrationamplitudeforthe49.7-inch,[
Nube with WEXTEX joint W(1) at the lower end, and an ideally-clasped condition at the upper end. Damping shown is for the N direction 7-72
a,b.C f 1 b l 9 l t 1 W
~
Pigure 7-12 -Dagpingversusvibrationamplitudeforthe49.7-inch,{
]tubewithWEXTEXjointW(3)atthelowerend,andan ideally-clamped condition at the upper end 7 73
- i [
8,b,C s Figure 7 Damping versus vibration amplitude for the 60.5-inch, [ ]M,ube with EXTEI joint W(3) at the lower end, and dented TSP sample D(1) at the upper end. Damping is shown for both the N and T directions 7- 74
e,b c 1 1 i i L T I i F . Figure 7 Comparison of damping in the N direction for the 60.5-inch tube, with WEXTEX joint W(3) at the bottom and t the dented TSP rample D(1) at the top (see Figure ; 7-13 ), with a tube with %TIITJ joint W(3) at the : bottom, and an ideally-clamped condition at the top (see ; Figure 7-10 ) 7-75 , I
~ , , _ _ , _ . _ - . . . , ,- .,
a,b.c l l l l Figure 7 Comparison of damping in the T direction for the 60.5-inch tube, with WEXTEX joint W(3) at the bottom and - the dented TSP sample D(1). at the top (see Figure 7-13 ), with a tube with WEXTEX joint W(3) at the bottom, and an ideally clamped condition at the top (see Figure 7-10 -) 7-76
a,b.c I6 Figure 7 Damping versus vibration amplitude for the 60.5-inch tube with TEITHI joint f(1) at the bottom, and dented TSP sample D(3) at the top 7 77
a,b.c e
!W Figure 7 Summary of the damping data for the 60.5-inch, d Hs tube. with a "EITEI joint at the botton, and a dented TSP at the top 7 78
- - .,_._.- .- i
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i r t t Figure 7-18 - Damping versus vibration amplitude for the 49.7-inch tube with WEXTEX joint W(3) at the bottom, and dented TSP sample D(1) at the top 7 79
a,b c
~
i Figure 7 Comparison of damping in the N direction for the l 49.7-inch tube, with WHXTEX joint W(3) at the bottom and l the dented TSP sample D(1) at the top (see Figure l 7-18 ), with the tube with WEXTEX joint W(3) at the l bottom, and ideally-clamped conditions at the top (see Figure 7-12 ) L 7 80 l
~ _
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- a,b.c
( l Figure 7-20 < - Comparison of damping in the T direction for the 49.7-inch tube, with EXTEX joint W(3) at the bottom r.ud the dented TSP sample D(1) at the top (see Figure 7-18), with the tube with EITEX W(3) at the bottom, and ideally-clamped conditions at the top (see Figure 7-12) 7-81
i a b.c r - i r b i 4 4 a Figuxe 7 Damping versus vibration amplitude for the 49.7-inch tube with WEITEX joint W(1) at the bottom, and dented
- TSP sample D(3) at the top i
7-82
t i
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Figure 7-22 - Summary of the damping data for the 49.7-inch, [ ]"'" ! tube with a WEXTEX joint at the bottom, and a dented TSP ; at the top 7-83
. ---y-yw y -,_y- , - --
a,b.c 4 c , t Fig, 7-23 Weasund tube damping with an idealized tube sheet at the lower end and a dented TSP D(4) at the upper end. Tests 1, 2 7 84 5s
a,b.c ' t
~
1 1 l l 4 t I l 0 Fig. 7-24 Measured tube damping with WEXTEX tube sheet simulant W(3) and dented TSP D(1) . Tests 8, 9 7-85
I a b,C s I
=
Fig. 7-25, Measured tube damping with WEXTEX tube sheet simulant W(3) and dented TSP D(1) . A 180 degree slot previously machined in the tube has been filled-in by beliarc welding. Tests 14, 15 7-86
1 i
~" a ,b.c -
l l G,' Fig. 7-26 Measured _ tube damping with WEXTEX tube sheet simulant W(3) and dented TSP D(1) . Circumferential slots are machined into the tube'to simulate the decreased stiffness that results when through-wall cracks occur. The ef fect of slot (crack) angle on the damping available from the WEXTEX tube sheet simulant is shown. Tests 8-17 7-87
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a,b.c (1 Fig. 7-28 Measured tube damping for SCC Saple NLS-12. The tube is supported at the bottom by an idealized tube sheet simulant and at the top by a dented TSP. Results are shown for the SCC only, and also f or the SCC with a f atigue crack. Tests 3-6 7-89
a b.c l l ~ Fig. 7-29 Variation of tube damping and crack angle (nr estimated from -- the tube resonant frequency) with vibration time for SCO . Sample NLS-12. Tube vibration amplitude is 10 mils peak-to-peak. The tube has an initial through-wall crack of 120/138 degrees. The total visual crack angle is 138 degrees, however, a 120 degree extent is separated from the rest of the crack by a short uncracked length. Tube has an idealized tube sheet simulant at the lower end and a dented TSD at the upper end. Test 7 7 90
a.D,c Fig.7-30 > Measured tube damping for SCC Sample SPI-24 with only the SCC. Tube has an idealized tube sheet simulant at the lower end and a dented TSP at the upper end. Tests 18, 19 7-91
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4 1 < l i e h Fig. 7-31. Measured tube damping ior SCO Sample F1I-N af ter a through- -' wall, 164 degree fatigue crack is formed. Tube has an ide lized tub + sheet simulant at the lower end and a dented TS2' at the v ,:- end. The variation of damping with time is shown. The aoe vibratinn amplitude is 5 mile peak-to-peak and the reronant frequency is api,roximately C 37' Open circles, plotted at the same time t as for the r 'id c.ircles, correspond to damping measured at shutdown at the end of the day. Solid circler, correspond to damping maasured when the test is restarted. Testr 20-22 7-92
)
_ _ __ _ _ _ _ _ _ _ _ _j
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i Fig. 7-32 Measured tube damping for SCC Sample SPI-22 with only a SCO. Tube has an idealized tube sheet simulant at the lower end and a dented TSP at the upper end. Tests 23, 24 . 7-93 ,
1 a,b,c l i i i Fig. 7-33 Measured- tube damping for SCC Sample SPI-22 af ter a through-wall, 157 degree fatigue crack is formed. Tute has an 4 idealized tube sheet simulani at the lower end and a dented ! TSP at the upper end. The variation of damping with time is thown. The tube vibration amplitude is 5 mils peak-to-peak, , and the resonant frequency is approximately [ }?'Open 1 circles, plotted at the same time t as for the solid circles, correspond to damping measured at shutdown at the end of the l day. Solid circles correspond to damping measured when the test is restarted. Tests 25-27 7-94 >
a ,b,( l* t Fig. 7-34 Weasured tube damping for SCC Sample b?I-22 with a 157 degree
- through-wall crack. The vibration amplitude is 8 mils peak-to-peak. Damping measurements were made following approximately 26 hours testing at 5 mils peak-to-peak. Test 28 7-95
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8.0 EVALUATION OF POTENTIAL CRACK PROPAGATION UNDER SLB FLOW CONDITIONS Response of a tube to normal opara' ion loadings is addressed in Section 7.0. That evaluation requires steady state thermal / hydraulic conditions which are obtainerd 'm the ATHOS code. A Steam Line Break (SLB) would subject steam generator tubes to a transient type loading where the fluid velocity and density profiles are comprised of the pressure difference time-history generated during the transient. These potentially cause Flow-Induced Vibration (FIV) response of the tubes due to the associated secondary side fluid flow velocities and densities along the tubes. The FIV evaluations consider the effects cf these crossflow velocities during the SLB and they are performed to determine the effects of a postulated SLB on partially degraded tubes as part of the straight leg evaluations. This section provides descriptions of the methcds and analyses used to perform the SLB ovaluations. 8.1 Steam Line Brmk Loau Analysis This section datermines the dynamic loads to tubes and ;abe support plates during steam line break (SLB) events. These loads provide the SLB transient flow conditions for flow vibration analyses of degraded tubes at the tubesheet area and tube support plates. 8.1.1 initiat Conditions of slo Blowdown Flow As reported in NUREG/CR 4407, " Pipe Break Frequency Estimation for Nuclear Power Plants," published in May 1987, there were 19 breaks over 810 plant y- 3 of US PWR and BWR plants. All of the are staall breaks in terms of leak rate ancor pipe size. None of them is related to the steam line. The steam generator design specification assumes one steam line break event ever 40 years. Bayesian analysis of thasc plant pipe break events suggests the frequercy of the steam line break to be about one over 400 years for each plant, which is much smaller than the design basis. Table 8.1 lists the 19 pipe breaks by operational mode. Norma! operation mode experienced 13 breaks, shutdown mode had 4 breaks, and starting up mode shared 2 breaks. Nono of the breaks were due to transients; this is understandable because the duration of the transients is negligibly short when compared to other modes of operation. As given in Table 8.1, there is only 0.6% of a year being spent in the transient mode. l 8-1
Normal power operation contributes 60% shutdown mode Ol ves 21% and startup modo shares 11% n if we infer from the above, should a steam line break event occur it could be expected to take place during normal power operation. However, one can not rule out that it could
- take place during shutdown or startup. Based on previous blowdown analyses of steam - line breaks, the hot shutdown at zero load will yield the largest thermal and hydraulic loading to intemal components among all of operational modes. The startup mode has its conditions near the hot standby mode.
Based on the above evaluation, the initial conditions are specified as the hot standby with zero power and a normal water level. For the 51 Series steam generator, the normal water level is 506" above the top of the tubesheet. 8.12 - Flow Calculations bv the TRANFLO Code The TRANFLO computer code is { Ja,c, North Ant ',it 1 utill:es three steam generators of typical 51 Series deslan. A computationt ,odel using TRANF'_O code was developed for the North Anna steam generators. ' Figures 81 shows location calculation nodes of the 51 Series steam generator. Figures 8-2 and 8 3 present the detailed computational model with nodes and flow connectors for the secondary' side fluid. Figure 8-4 shows the primary side fluid nodes, heet transfer nodes of tubes and fluid or heat connectors. The computatio: model consists of the following elements: f a,c 8-2
l l _ a,c l . As can bo Foon in the figures, the model dividos the region between the tuboshoot and the first (lowest) tubo support plate into [ lac. The SLB loads to tubes in the tubesheet vary radlally; the loads increase toward the periphoral tubes. Not all of the tubes in the tubosheet area are subject to degradation. In general, the circumferentially cracked tubes are confinod to the interior tubos where the sludge pilo forms. Thorofore, [ ja,c, The following subsections prosent pressure drops througt, the tube support plates and thormal and hydraulics in the U-bond, and thermal and hydraulics at the tubesheet. 3.1.3 Pressure Droos throuah Tube Succort Plates The pressuro drop through tube support plates (TSP) for large and small steam line breaks were calculated. The !;rge steam line break represents a break upstream of the flow limiter, such as just outside the steam nozzle. The small steam line break simulates a break downstream of the flow limitor. The large break has a blowdown flow area of 4.6 ft2, and the small break has an area of 1.4 ft2 , The first (1st) tube support plate is near the tubesheet where the flow is highly three dimensional with s*rong crossflow across the tubesheet. The pressure drop through th) 1st TSP shows a radial oistribution. The tube rows are divided into five radial zones. Each radial zone corresponds to a secondary sido fluid node, as shown in Figure 8 3. The innermost zono has the highest pressure drop among the five zones. The uppermost (i.e., 7th) TSP has the highest peak of averago pressure drop over the whole plato. For the large break, the peak is [ la,c and it takes place at timo equal to about [ Ja,c seconds after break For the small break, this peak is 8-3
[ ]ac and it occurs at tima equal to about [ ]a,c seconds. After peaking, the pressure drop quickly decreases to a small value. 1 8.1.4 Flow Conditions at the First Soan Above the Tubesheet . The crossflow velocity between the tubesheet and the 1st TSP can be significant whlle the remaining span of the straigh* leg of the U tube experiences negligible crossflow velocity. Significant crossflow velocities may result in flow inouced tube vibration. Tube vibration in turn could cause the further growth of cracking of the circumferentially degraded tube. Fluid density, velocity, vold fraction and local pressure were calculated for five representative tube rows during e large steam line break and a small steam line break. l l The five representative rows are Row Nos. 8,17,27,36 and 46. They are considered to l represent the channel between tube Columns 47 and 48. For other tube rows, linear interpolation can be used to determine the density, velocity, vold fraction and pressure for .Other tube rows. For tube columns other than 47 and 48, the principle of ax' symmetry applies. For exampie, a circle can be drawr, using the radius of the tube
- Row 36; results for tube Row 36 are also applicable to any tube ateng the circle.
Over the wrapper opening span (i.e., 0" to 14" above the tubesheet), the crossflow velocity is such that it increases as tube location moves radially outward. Tubes at the outermost tube row (i.e. Row 46) experience the highest crossflow velocity. The peak velocity takes place at about [ ]a,c second after break. For the span from 14" to 50" (1.0.,1st TSP) above the tubesheet, the crosstlow velocity increases from Row 8 toward Row 36, and it then decreases to zero from Row 36 to Raw 46. As far as tube vibration is concerned Row 36 is judged to experience the highest energy input. Row 46 has the highest peak velocity over the 14" span abeve the tubesheet, but zero over tne next 36" span; its overall energy input to the tube should be less than Row 36. ! Tubes which are exposed to the sludge pile on the top of the tubesheet may be subject to tube degradation. The potential sludge pile zone can be determined by fluid flow conditions at the tubesheet during normst power operation. Based on three dimensional
~
thermal and hydraulic flow field calculadons using the ATHOS computer code 'or the 51 Series steam ger ear, Figure 8-5 shows contours of crossflow tube gap velocity at the 8-4
L: 7 tuboshoot. According to laboratory tests and field inspection, the tube zone with a tubo -
, gap velocity of less than [ ja.c ft/ soc is potentially subject to- sludge accumulation at the tubosheet. in addition, circumferentially cracked tubos at or just above the ,- tubeshool wore found essr QTj within the potential sludge zone, as discussed in Section 4.
As discussed earlier, velocity increases toward the porkhoral tubes. The Row 31 Column 33 (R31C33) tubo represents the outermost tube among tubos located within the potential sludge zone. This R31C33 tubo can bo chosen to yield a bounding evaluation of tubo vibration due to SLB for degraded tubes in the sludge pile. Thermal and hydraulic conditions for the R31C33 tube during large and small steam line breaks woro datormlned. The conditions consist of tube gap crossflow velocity, fluid density, vold fraction and pressure as a function of time. These conditions apply for the tube span from the tubesheet to the first tube support. The peak velocity is about [ la,c and the corresponding fluid density is about [ la,c. During normal full power operation, this tube is subject to a velocity of [ Ja,c and a density of [ ja.c. So, the peak dynamic pressure is abvut 0.8 psi during the steam -line break, and about [ ja c during full power operation. 8.2 Vibration Analysis Methods for Steam Lino 5 reak Conditions Fluid velocity and density distribution time-historios are gonorated for the SLB conditions using the Westinghouso propriotary computer code TRAUFLO. The time history responses of the tube to any given set of loadings rnpresenting potential limiting flow and response conditions arts- determined through a [ ja c, it is important to note that tube damping is [ n e ja.c, 85
This ' capability- Is embodied in the [ ja c, These SLB time-history responses are compared to the threshold strets for crack propagation response obtained during normal operation (steady state) condition evaluations. Should the tube be acceptable under the limiting SLB transiem, then the tube will be acceptable under transients which produce smaller flow velocides, but which extend over a longer period of time. 8.3 Straight Lcg Large and Small Steam Line Break Vibration Evaluation -It has been postulated that various velocity and density time-histories could occur during SLB. [ ja,c, A matrix of the SLB time history cases evaluated for the North Anna 1 assessments is - given in Table 8-2. Note that both a large and a small steam line break are included for completeness. [ jac, For tubes near the periphery only the case with a crack at the first tube support plate j was considered because the information in Section 4.0 shows there are negligible circumferential cracks expected at the tubesheet outside of the sludge pile (interior) zone. l 8-6 1 l
I From the analyses which form the basis for the SLB evaluations it is concluded that the
, major tube response 1c to fluidelastic forces and that this response occurs e.pproximately [ Ja.c into the transient. The fiuldelantic response does r.nt , build up much because rf the limited time during which the tube is unstable.
Longer term turbulence forces would not be limitin0 because of the very low SLB secondary flow velocities at longer times into the transient. The SLB displacement results of Table 8-2 can be evaluated in terms of absolute magnitudes of tube stress. I Ja,c, On the basis of the analyses end results reported in this section it is concluded that neither the large nor small SLB events represent limiting conditions for the VRA cracked tube evaluations. 8.4 Conclusions TRANFLO SLB analyses for small and large breaks were performed to obtain crossflow conditions in the first pass to assess the potential for crack propagation under ELB conditions. These analyses included radial and axial distributions of the flow conc'itions to characterize the flow magnitudes and distributions of the flow distributicas for vibration assessments. The analysis results show very low vibration amplitt des for through wall circumferential cracks as I arge as [ Ja,c, such that r egligible potential exists for crack propagation in the straight leg of the tube under SLB
- onditions.
8-7
l: Table 8-1 Number of Pipe Breaks Categorized by Operational Mode -i Average % Thermal tio. of of Year % of Operational Mode Power, % F::llures Spent in Mode Contribution Starting Up s5 2 2.5 11 Normal Power Operation >5 13 62.5 68 Shutting Down 15 0 --- - Shutdown 0 4 33.6 21 1ransient - -- 0 8 Transients / year - All Modes --- 19 99.4 100 Note: Percen' of mode is based on actual operating exportertoe of US PWR and BWR plants in 1983. t 8-8
. . _ . . _ ~ ~ . . . _ ... ~ - . .-. _. , TABLE 8 2 VRA SLB EVALUATIONS MATRIX AND RESULTS Tube Bundio Location I la,C inie rio__r Perlohe rv IIS US & TSP 1 TSP 1 ISP_1 6
Disolacement Results (inches x 10 ) Small SLB [ ]a,c Large SLB [ }a,c t Essentially same result obtained for both Row 36 and Row 46 peripheral tubes. P d e O a 89
'~
8,C-
.l l
I
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( J 4 4 l Figure 8-1 Schematic of 51 Series Steam Generator and Secondary Fiow Nodes for TRANFLO Calculation 8-10
~ B,C Figure 8-2 Secondary Side Fluid Nodes and Flow Connectors for 51 Series Steam Generator 8-11
- - a,c Figure 8-3 Secondary Side Fluid Nodes between Tubesheet and 1st Tube Support Plate e
1 8-12
_ a,e k 1' i* A a-Figure 8-4 Primary and Secondary Side Fluid Nodes, Tube Metal Nodes, and Heat and Primary Flow Connectors 8 13
a.c
~\
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~
Figure 8 Crossflow Velocity' Zones at Tubesheet of Typical 51 Series Steam Generator (Symmetry along the line between Tube Columns 47 and 48 are assumed) o l l l 8-14 [
9.0 COMBINEDACCIDENTANALYSIS This section describes SSE and LOCA analyses performed to assess the potential for crack propagation under these individual and combined events. The potential for tube deformation r. tutting from TSP loads is also evaluated. 9.1 SSE Analysis Seismic (SSE) loads are developed as a result of the motion of the ground during an earthquake. A seismic analysis specific to Series 51 steam generators has been completed. Response spectra that umbrella a number of plants with Series 51 steam generators, including the North Anna Plant, have been used to obtain tube support plate (TSP) loads and the displacement time history response of the tube bundle. A nonlinear j time-history analysis is used to account for the effects of radial gaps between the secondary shell and the TSP's, and between the wrapper and shell. l The seismic excitation defined for the steam generators is in the form of acceleration response spectra at the steam generator supports. In order to perform the non-linear
-{
time history analysis, it is necessary to convert the response spectrum input into f acceleration time history' input. Acceleration time histories for the nonlinear analysis l are synthesized from El Centro Earthquake motions, using a frequency suppression / raising technique, such that each resulting spectrum closely envelopes the corresponding specified spectrum. The three orthogonal components of the earthquake are then applied simultaneously at each support to perform the analysis. l The seismic analysis is performed using the WECAN computer program. The mathematical model consists of- three-dimensional lumped mass, beam, and pipe elements as well as general matrix input to provide a plant specific representation of the steam generator and reactor coolant piping stiffnesses. In the nonlinear analysis, the TSP /shell, and wrapper /shell interactions are represented by a concentric spring-gap dynamic element, using impact damping to account for energy dissipation at
- these locations.
The mathematical model which is used is shown in Figure 9-1. The tube bundle straight leg region on both the hot-leg side and cold-leg side is modeled by [ ja,c 9-1
Ja,c, A summary of the resulting TSP forces are summarized in Table 9-1. The TSP forces are provided for the initial 3 seconds of the transient when the LOCA event is in progress. The highest TSP force, [ la,c, occur late in the transient (> 15 soc), after the LOCA forces have diminished. Plots showing the in plane and out of-plane displacement time histories for the tubesheet and for the top TSP (TSP
#7) are shown in Figures 9-2 through 9-5, respectively. Note that the results for TSP 7 are typical of the other TSP locations.
In order to evaluate the effects of ci"cumferential cracks on the seismically induced tube stresses, an analysis was performed [ ' Ja.c. Two separate tube passes were analyzed, the first pass, and tube pass number seven. In reviewing the displacement time history _ response for the seven TSP's from the overall model, it was concluded that the responses are quite simliar, with TSP 7 having the highest magnitude response. On that basis. [ ja.c, Four cases were considered for each of the two tube passes. An initial reference case, with no cracks was analyzed, followed by the following three cracked tube cases, a crack o at the bottom of the tube, a crack at the top of the tube, and a crack at the top and bottom of the tube, in each case a crack extent of [ Ja.c was considered. The crack was introduced into the model by modifying [ ja.c, 9-2
The analysis was performed using a series of special purpose programs to develop load y : steps from the overall model-solution, perform the dynamic analysis calculating the displacement response of the tube, and then calculating the corresponding bending stresses in the tube, The programs were verified by comparing the response to a time t history solution using a general purpose finite element program for a short time period ! of the transient, i i in each case, the presence of the crack dramatically reduced the bending stresses l adjacent to the ' cracked region. For the reference case, the maximum tube bending stress is approximately [ ja.c. With the crack present, the maximum tube ! stress is on the order of [ ja.c, with the majority of the time history bending ._ stresses being less than [ Ja,c. A plot of the reference stress time history for f the first tube pass adjacent to the tubesheet is shown in Figure 9-6. The corresponding plot with the crack present is shown in Figure 9-7. A summary of the maximum and minimum tube bending stresses for each of the cases analyzed is summarized in Table 9-2. ' Although the crack analyzed was somewhat larger than the maximum crack size now expected, the analysis results should not be significantly affected. A table summarizing the equivalent moment of inertia for the tube cross-section as a function of crack angle ; [ is provided in Table 9 3. This shows that the section is essentially a pin joint for both
}
crack angles. ! 9.2 LOCA Analysis f i LOCA loads are developed as a result of transient flow, and temperature and pressure fluctuations following a postulated coolant pipe break. As a result of a LOCA event, the
-steam generator tubing is subjected to the following loads:
- 1) Primary fluid rarefaction wave loads.
2
- 2) Steam generator shaking loads due to the coolant loop motion.
- 3) External hydrostatic pressure loads as the primary side blows down to almospheric pressure.
9-3
g
- 4) Bending stresses resulting from bow of the tubesheet due to the secondary to primary pressure drop.
- 5) Bending of the tube due to differential thermal expansion between the tubesheet and first tube support plate following the drop in primary fluid temperature.
- 6) Axially induced loads resulting from differential thermal expansion between the tubes and tie rods / spacers due to the tube being tight in the <
first TSP, and the reduction in primary fluid temperature. (Based on available data, the majority of intersections are considered to be tight. ,
' Because the majority of the intersections are tight, the TSP will respond with the tubes, and the resulting loads on the tubes are judged to be small.).
The rarefaction wave which passes through the tube results primarily.in bending stresses in the tube U-bends at the top TSP. The tube is mentially unaffected at the top of the tubesheet, and at TSP 1-6. During the time the rarefaction wave is passing
. though the bundle, the primary to secondary pressure differential never exceeds the '
normal operation value.
. The LOCA shaking condition results in bending loads due to the shaking of the steam generator caused by the break hydraulics and reactor coolant loop motion. To obtain the LOCA induced hydraulic forcing functions, a dynamic blowdown analysis is performed to obtain the system hydraulic forcing functions assuming an instantaneous (1.0 msee break opening time) double-ended guillotine break at the steam generator outlet nozzle.
The hydraulic forcing functions are then applied, along with the displacement time-history of the reactor pressure vessel (obtained from a separate reactor vessel 1 blowdown analysis), to a system structural model, which includes the steam generator, i the reactor coolant pump and the primary piping. This analysis yields the time history displacements of the steam generator at its upper lateral and lower support nodes. f These time history displacements formulate the forcing functions for obtaining the tube stresses due to LOCA shaking of the steam generator.
.c
- - Loading mechanisms (3) through (5) above are not an issue since they are a non cyclic j loading condition and will not result in crack growth, and/or result in a compressive 9-4 I
~
membrane loading on the tube that is beneficial in terms of negating cyclic bending'
, stresses that could result in crack growth.
9.3 - SSE + LOCA Effect on Umiting Tube Condition 9.3.1 Tube Deformation For the combined SSE + LOCA loading condition, the potential exists for yielding of the tube support plate in the vicinity of the wedge groups, (see Figure 9-8), accompanied by deformation of tubes and subsequent loss of flow area and a postulated in-leakage. Tube deformation alone, although it impacts the steam generator cooling capability .; following a LOCA, is small and the increase in PCT is acceptable. Consequent in leakage, however, may occur if OD circumferential cracks are present and propagate through wall as tube deformation occurs. This deformation may also lead to opening of pre-existing tight. through wall cracks, resulting in primary to secondary leakage during the SSE + LOCA event, with consequent in leakage following the event, in-leakage is a potential concern, as a small amount of in-leakage may cause an
. unacceptable increase in the core PCT.
For a LOCA event, the tubes are subject to both a rarefaction pressure wave that travels through the tube bundle, and to loads resulting from shaking of the overall steam _ generator. The rarefaction wave results in a hot. leg to cold bg pressure differential that causes a lateral load to be imposed on the tube U-bend. The lateral load is reacted by the TSP through wedges that bear against the wrapper and shell wall. The lateral load varies from row to row both in amplitude and period due to the different bend radiL Integrating this load over the entire bundle ma/ result in a significant load on the TSP. The LOCA shaking and the SSE event also result in lateral motions of the bundle and subsequent TSP loads. Typically, the LOCA shaking loads are small compared to LOCA rarefaction, but the SSE loads may be significant, although generally less than the LOCA rarefaction loads for a major pipe break. TSP loads are ' calculated for each of the
?-
loading , achanisms through independent time history analyses. To get an overall TSP
. load, the LOCA rarefaction and LOCA shaking loads are first combined on an absc!ute basis (which is very conservative), and then the total LOCA load is conservatively combined with the seismic load using square root of the sum of the squares. For the LOCA rarefaction loading, the top TSP reacts the majority of the load. Seismic and LOCA 9-5
7 shaking loads affect all of the TSP's to a varying degree, depending on the frequency content of the loading versus the system response characteristics. 4 9 The pressure differentials that occur dtring a LOCA are a function of the pipe break location. Because North Anna has qualified leak before break for primary piping, major pipe breaks need not be considered. Analysis of two minor pipe breaks, the Pressurizar Surge Line, and the Accumulator Line, has shown that the Accumulator Line break results in the highest TSP loads. Except for the bottom TSP, the wedge groups for each of the TSP's are located at the same angular location as for the top TSP (see Figure 9 8). Thus, if TSP yloiding occurs at the lower plates, the same tubes are affected as for the top TSP. For the top TSP, however, the wedge groups have a [ }a,c width, compared to a [ ja,c width for the other plates. This larger wedge group width distributes the load over a larger portion of the plate, resulting in less plate and tube deformation for a given load , level. For the bottom TSP, the wedge group width is [ la,c, and the wedge groups are rotated [ ja,c relative to the other TSP's. Thus, separ,'Its calculations to determine the number of deformed tubes are performed for TSP 1, TSP 2 6, and for TSP 7. The limiting conditions, in terms of TSP wedge loads, are found to be SSE + LOCA for TSP's 1 and 7, and seismic alone for TSP 2 6 (due to the higher seismic load late in the seismic event). Utilizing crush test data recently obtained for Series 51 tube support plates, analysis results show that there will not be any tubes that undergo significant permanent deformation (a change in diameter of >0.025 inch). Based on these results, it is judged that there will not be any significant in leakage under the post LOCA secondary to primary pressure drop. 9.3.2 Crack Gromh The purpose of this analysis is to estimate the crack growth under LOCA + SSE loads for C North Anna (Serier 51) tubes in the WEXTEX transition zone, and at TSP locations. The applicable LOCA shaking stresses are [ jb,c, 9-6
These stresses are for the reference geometry and do not account for the presence of cracks in the tube. Based on the results for the selsmic analysis, where a significant reduction in bending stress occurs when the crack is present, it is anticipated that the LOCA shaking stresses would also be reduced. Thus, these stresses are concluded to form a conservative basis for the crack growth calculations. The applicable seismic stresses, accounting for the presence of a crack in the tube of [ la,c, are shown in Figure 9-7. This plot shows the maximum stress range to be on the order of [ l b ,c. Combining the selsmic and LOCA stresses by the square root of the sum of the squares gives a cyclic stress amplitude of [ ]b.c for the limiting condition. For the remaining seismic response, the vast majority of the stresses are less than [ jb,c, Thus this stress level will be assumed to combine with the LOCA shaking stress of [ ]b.c, to give a combined stress of [ ]b,c, For the LOCA event, after a time period of approximately 6 seconds, the secondary to primary AP results in a compressive axial stress that offsets the bending stress, and no further crack growth will occur. Assuming the inillal 3 seconds of the selsmic avent to o be typical of the subsequent 3 seconds, one additional stress range of [ )b,c will be experienced during the next 3 seconds. Thus two cycles of [ )b,c wjti be assumed for the crack growth calculations. The applicable number of cycles for the [ )b,c stress range is 240. The crack growth calculations assume an initial crack of[ -)b,c, or a half angle of [ ]b c. The formulation for calculating the number of cycles to give a growth in the half angle of 1* is: Da Dn - CGC x (Del Kl)CGE CGC - Crack Growth Constant CGE - Crack Growth Exponent 97 I r
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i
The crack growth calculations show that for 2 cycles, negligible crack growth will occur for a stress of amplitude of [ )b,e and an initial crack _.f angle of - [ )b,c. For an RMS stress of [ ]b,c, a small amount of crack growth occurs, such that the final crack length has a half angle of ( )b,c. Thus, the final ovorall crack length is [ ]b,c, and the tube will maintain its structural intecrity under combined LOCA + SSE loading. 9.4 SSE + SLB Effect on Limiting Tube Condition The steam line break event is, typically, about 2 seconds in duration for loadings in the generator (with a double ended main steam line pipe break - 4.6 ft2 break area). During this time period the peak velocity across the tubesheet in the sludge pile region is approximately 4 times the normal operation velocity. Even with this velocity increase, the tubes in the centra! hot leg region are not susceptible to significant fluidolastic vibration and therefore, no growth of existing circumferential cracks less than or equal to [ ]b,c in extent would be expected to occur. O 9-8
T l. t
.- Table 9.1 -
Summary of Seismic TSP Forces s I t a,C , F e I { h e I t s Muur Y 5
?
m I h D I i t 9
-I e
i t= i V h i s I' r i 9-9
Table 9.2 - Summary of Soismic Stresses Tubes with Circumferential Cracks a,c l l l 4 Omsuu 9 10
l Table 9.3 Tube Moment of inertia as a Function of_ Crack Angle 8,C-T I , a
.4 P
2.
~
9 11
a,c-s
.i l
k p s f Figure 9-1 Finite Element Model Cverall Seismic Analysis l 9-12 4
- 6. !
1
. . . - . . . . . . - . - .. - - - . - ... . - . . . - . . . . . . = . - .~
B,C l l-t .N l i I
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I ~ Figure 9-2 Seismic Displacement Time History Top of the Tubesheet i UX (In-Plane of U-Bend) Degree of Freedom l i 9-13
8,C i J e
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Figure 9 3 Seismic Displacement Time History Top of the Tubesheet UZ (Out-of-P!ane of U Bend) Degree of Freedom 9-14
;I l
I 8,C 4 e I b i-3 -. Figure 9-4 - Seismic Displacement Time History - TSP Number 7 Ux (In Plane of U. Bond) Degree of Freedom 9-15
4 8,C W 1 ) i - b' t i I e
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Figure 9-5 - . Seismic Displacement Time History TSP Number 7 a UZ (Out of Plane of U-Bend) Degree of Freedom l 9-16
a,c
}
i a :
- 1 Figure 9-6 - '
- i Tube Stress Time History ,
Topof theTubosheet g Reference Tube Geometry ' 9-17
y
~-
a,c
%f i
i 1 l A r l 9 -
~
Figure 9-7 - Tube Stress Time History i Topof theTubesheet 240" Circumferential Crack Present 9-18
, _ , .- . . ._- _ ._m--- . . - . .. .. _ - , .
a.c I s S- .c h l l I l l 1 i - Figure 9-8 ,
- Wedge Group Locations l
i l l 9-19
wmpaw 4- emmege-Nm emsm- -'-w,eims.- 49,maar.am. m e l l
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