ML20235K937

Engineering Evaluation of San Onofre Nuclear Generating Station I Thermal Shield Supports

ML20235K937
Person / Time
Site:	San Onofre
Issue date:	02/28/1989
From:	Goossen J, Schwirian R, Yu C WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML13331B132	List: ML20235K937
References
WCAP-12149, NUDOCS 8902270309
Download: ML20235K937 (135)

Text

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. WESTINGHOUSE CLASS 3:

4 WCAP-12149 b.

i e

ENGINEERING EVALUATION OF THE SONGS I THERMAL SHIELD SUPPORTS February, 1989

. Authors J. E. Goossen D. R. Bhandcri R. E..Schwirian D. R. Forsyth C. Yu M. F. Hankinson N. R. Singleton A.'J. Kuenzel R. R. Laubham Approved by: Y.

C. H. Boyd, Manager

-RPV System Analysis Work Performed Under Shop Order SNNP-25 1

i' WESTINGHOUSE ELECTRIC CORPORATION

,. Nuclear and Advanced Technology Division j P.O. Box 2728 -

Pittsburgh, Pennsylvania 15230-2728 c.

r 8902270309 890217 DR ADOCK 0500g.'Cy6 >

3588s-020289 10

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WESTINGHOUSE CLASS 3 1

TABLE OF CONTENTS Section o ' Tit 1e Page

1.0 INTRODUCTION

/ BACKGROUND 1-1 )

2.0 DESCRIPTION

OF SUPPORT DESIGN 2-1

)

J 3.0 HISTORY OF SONGS I THERMAL SHIELD 3-1 I 4.0 SCE ACTION PLAN BASED ON HADDAM NECK SITUATION 4-1 5.0 OBSERVED DEGRADATION 5-1 6.0 ANALYSIS TO DETERMINE CAUSE OF DEGRADATION 6-1 6.1 Vortex Shedding Analysis 6-2 6.2 Vibration Analysis 6-8

.6.3 Thermal Evaluation 6-26 6.4 Structural Analysis and Explanation of History 6-36 6.5 Conclusions Regarding Expected Progression of 6-45 Degradation 6.5.1 Lower Support Blocks 6-45 6.5.2 Remaining Flexure 6-47

,, 7.0 CONSEQUENCES OF FURTHER DEGRADATION 7-1 7.1 Definition / Evaluation of Worst Possible Degraded Cases 7-1 7.2 Stability Analysis 7-5

- 7.3 Vibration Analysis 7-17 7.4 Seismic Analysis 7-18 7.5 Impact Loads Evaluation 7-27 7.6 Effect of Seismic, Vibratory and Thermal Loads on 7-33 Core Barrel 7.7 Wear Evaluation for Thermal Shield Support Blocks 7-36 8.0 LOOSE PARTS ANALYSIS 8-1 9.0 EVALUATION OF WORST CONCEIVABLE DEGRADED CONDITION 9-1 9.1 Hydraulic Evaluation 9-1 9.2 Thermal Shield Drop Analysis 9-6 10.0 MONITORING PROGRAM 10-1

11.0 CONCLUSION

S 11-1 la i*

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WESTINGHOUSE CLASS 3 4 l-I SECTION

1.0 INTRODUCTION

/ BACKGROUND SONGS I is one of the three Westinghouse PWR's which utilize a thermal shield that is supported as an inverted pendulum. In this design, fixed supports

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clamp the bottom of the thermal shield to the core barrel and the upper end of the thermal shield has gapped upper supports (called displacement limiter keys) plus two of the three plants had additional supports (called flexures) connecting the top of the shield to the core barrel. In both SONGS I and the other domestic plant which utilizes the inverted pendulum configuration, some l or all of these flexures cracked or broke early in plant life. Starting with the post hot functional inspection which was performed in 1966, vibration of the thermal shield has resulted in various issues associated with the thermal shield. A detailed history of these issues is contained in Section 3.0.

In order to conduct the required 10 year in-service examination of the reactor vessel and the reactor internals, it is necessary to remove the lower internals package from the reactor vessel. When this was performed in August of 1987 at the other domestic plant which utilizes the inverted pendulum support design, it was found that some of the fasteners (bolts and dowel pins) in the lower thermal shield supports were not in place and damage was noted to some of the irradiation specimen holders which are mounted on the outside of the core barrel and the thermal shield. A summary of the 1987 experience at this other plant follows:

During initial attempts to remove the lower internals assembly, difficulties were experienced, and it was necessary to increase the allowed lift load in order to pull the internals free. This difficulty in removing the internals was the first indication that the condition of the lower internals might have degraded. Based on observations made after the internals were removed, it is now eviden' that some of the dowel pins in the thermal shield support block were gouging into some unused support blocks on the reactor vessel and resisting attempts to lift the internals.

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3588s-020289.10 1-1 g

WESTINGHOUSE CLASS 3

.The original support system used .on the thermal shield at that plant-consists of six fixed support bic ks on which the bottom of the thermal

• shield rests and four lateral displacement _ limiters at the top of-the shield. The lower supports are bolted and dowelled to'the core barrel

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with one inch bolts and dowel pins and the six supports fit into a circumferencial groove on the outside of the core barrel. The fasteners-were originally locked in place by the use of washers which were welded to the pins and bolts and the thermal shield. Once the lower internals were removed from the vessel, it was noted that a number of the dowel pins had broken their lockwashers and " backed out" of the shield. Three of the eighteen visible bolts had also broken their lockwashers and  !

" backed out" of position. These bolts were ultimately retrieved from the bottom of the reactor vessel. Although the remaining fifteen visible bolts did not show visual evidence of damage, it was decided that these bolts should be removed and inspected for damage since any motions that-were significant enough to cause some of the bolts and dowel pins to come out, might also have damaged the other bolts.

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During removal of.the remaining visible bolts, the majority of the other visible bolts were also found to be failed. Only at one support block location were all three visible bolts found to be intact and not visibly I cracked. The torque required to remove the bolts ranged from essentially zero (broken bolts) to 300 ft-lbf.

In addition to the visible bolts, the lower supports also contain j

" hidden" bolts which could not be visually examined or removed without a major field modification effort. Although these bolts were originally -

l included to facilitate installation of the thermal shield onto the core barrel and not necessarily as major structural fasteners, they do help maintain a tight joint between the shield and the barrel, and it was ,

desired to learn something about their condition as well. Based on the results of this inspection, we were advised that the hidden bolts all

• - exhibited indications of cracks. Since there were indications in the hidden bolts, even at the one support block where the visible bolts were l' not broken, there still remains some question as to whether these are truly cracks, but for conservatism, it was assumed that all the hidden bolts are indeed cracked.

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1 WESTINGHOUSE CLASS 3 I

The lateral displacement limiters near the top of the shield serve to limit the relative lateral (and to some extent, radial) displacement o

between the core barrel and the shield. Fiberscope examinations plus feeler gauge measurements have shown that the gaps in these displacement limiters have worn over time from roughly 20 mils at installation up to as much as 181 mils. This means that the amplitude of vibration of the thermal shield was substantially higher than it was when the plant initially went into operation. This increase in vibratory levels results in increased loads on the thermal shield support blocks and the fasteners in those blocks.

Metallurgical evaluations of the failed 316 cold worked bolts led to the conclusion that cracks initiated as a result of high cycle fatigue and that these cracks propagated as a result of environmentally assisted corrosion fatigue. Engineering analyses have led to the conclusion that mechanical wear at the lateral displacement limiters over a period of time resulted in increasing cyclic stresses in the bolts which ultimately 1ed to fatigue failures of at least some of these bolts. It is believed that as the bolts began to lose function, the resulting decrease in stiffness of the thermal shield support system resulted in increased vibratory levels of the shield and this further accelerated the degradation of the bolts and the wear at the displacement limiters.

While stability analyses indicate that the shield did not become unstable, they do indicate that vibratory levels would have increased significantly given the observed condition of the support system at that plant.

As a result of the situation at this plant, Westinghouse and SCE reviewed similarities and differences in an attempt to determine the proper course of action for SONGS 1. Although the designs were very similar, there were some differences which suggested that if a similar situation existed, it might not .

be as extensive. For example, all of the flexures at the other plant had

=

failed after only a short period of operation, but the flexures at SONGS have l failed gradually. In fact, it was learned during the present inspection that l'

one of the flexures at SONGS is still intact. Also, whereas the displacement limiters in the other plant limited tangential motion, those at SONGS limit 3588s-020289 10 gg

WESTINGHOUSE CLASS 3 the radial motion of the shield. Another difference is the fact that the SONGS thermal shield is not as thick (2.5" vs. 4.2"), is- not as heavy (48,000 lbf vs. 86,000 lbf) and experiences somewhat lower flow velocities. In order to proactively address the questions raised, SCE initiated a plan to perform an inspection of the support system and an analysis effort to predict the likely condition of.the shield. Section 4.0 of this report presents the overall action plan that was developed for the issue.

Now that the analytical and inspection results are available, and since those results confirm that the condition of the SONGS support system is not as severely degraded as observed at the other plant, it has been concluded that deferral of a repair for another fuel cycle can be accomplished without adversely affecting t.afety and, it is believed, with little additional economic risk.

The primary purpose of the present report is to explain the bases for this conclusion and to demonstrate that SONGS I can return to service and operate for an additional fuel cycle without experiencing substantial additional support system degradation. At that time, the reactor internals will be removed from the reactor vessel in order to perform the ten year in-service examination. This will facilitate performing detailed inspection and repair  !

activities.

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WESTINGHOUSE CLASS 3 SECTION

2.0 DESCRIPTION

OF THERMAL SHIELD SUPPORT SYSTEM o

As shown in figure 2.0-1, the thermal shield in the SONGS I plant is supported as an inverted pendulum; there are six support block assemblies which fix the bottom end of the shield relative to the core barrel. These lower supports are fitted into a circumferential groove in the core barrel (approximately 0.6 inches deep) and are attached by a series of bolts and dowel pins. These supports are azimuthally located 60 degrees apart (0, 60, 120, 180, 240 and 300*). Figure 2.0-2 shows the details of the support block assemblies. A center bolt which clamps the top of the support block to the core barrel during initial assembly is 3/4" in diameter and two long bolts, which are 7/8" diameter, clamp the thermal shield to the top of the block and the core barrel. At the bottom of the block, there are two bolts which clamp the bottom of the block to the core barrel and two dowel pins to limit the vertical motion between the block and the core barrel. In

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between the upper and lower fastener elevations, there are two dowel pins which

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limit the relative vertical motion between the thermal shield and the core )

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barrel. The dowels obviously limit the bending and shear loads experienced by the l

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bolts. Any relative radial displacements between the barrel and the shield are j limited by the bolts. The support blocks used in the other plant have a similar arrangement but there are differences in the number, size and locations of the bots and dowel pins. In addition, the bolts in the two plants use a different locking design. Whereas the other plant uses lockwashers which are welded to the heads of the bolts and to the thermal shield, SONGS I uses a lockbar which fits into a groove in the bolt head and is welded to the thermal shield. The primary advantage of the lockbar design is that it is not susceptible to preload losses that might occur when the lockwasher is welded to the head of the bolts.

There are also four upper displacement limiters which are located roughly 9.5 inches below the top of the shield. These displacement limiters are located 90 degrees apart (0, 90, 180, and 270') and initially (at the beginning of plant operation) limited the relative radial motion between the barrel and thermal l l* shield at the key locations to approximately 15 mils (.015 inch) at normal

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operating conditions. Initially, these keys limited the lateral or tangential motion between the two components but had a large radial gap. The design was modified following evidence of large relative radial displacements during the nu..cnne' 2-1 i

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Figure 2.0-2. SONGS I Thermal Shield Support Block Design 1588s-020289 10 g3

WESTINGHOUSE CLASS 3 initial hot functional test program in 1966. The initial design of these keys was essentially identical to those which were used in the other domestic plant o

which has an inverted pendulum support arrangement. Unlike SONGS I, the original tangential restraints were not modified for that plant.

Another modification that was made after the SONGS hot functional test was to install six (6) flexure supports between the top of the thermal shield and the core barrel. The flexures are located at 21*, 85*, 124*, 205*, 244* and 325*.

As discussed in Section 3.0, only one of the flexures has been fully func-tional (intact) since sometime between 9/76 and 9/78. This is the 124' flexure and the January, 1989 inspection confirmed that this flexure is still intact and functional. The remaining flexures were also inspected to the extent possible and they are still inplace even though they are cracked or broken.

This means that these flexures do offer some restraint against excessive thermal shield motion but it is not possible to quantify their effectiveness so the present evaluations assume that only the 124* flexure is functional. It

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is noted that this is one of the significant differences between SONGS I and

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the other plant. In that plant all six flexures failed during the initial cycle of operation (17 years prior to observing significant degradation at the lower support blocks ) and they were physically removed af ter that cycle of operation. Thus any restraint offered by the damaged flexures was lost. This plus the fact that SONGS still has one fully functional flexure is at least partially responsible for the apparently slower rate of degradation at SONGS as compared to the other plant.

The nominal core barrel thickness at SONGS I is 1.75 inches and the thermal shield is 2.5 inches. The relative thickness of these two components influences the thermal lag that occurs during thermal transients. Since the core barrel is thinner, cooldown type temperature transients may result in barrel temperatures dropping faster than the thermal shield temperature and this may influence the low cycle fatigue loads experienced by the bolts.

However, this effect is less pronounced for SONGS than for the other plant since that plant uses a much thicker shield (4.2 inch) and the same core

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barrel thickness. This plus the shorter operating time for SONGS I ( 13 EFPY vs roughly 17 ) means that any low cycle fatigue damage is probably less severe for SONGS I.

3588s-020289 10 g.4

WESTINGHOUSE CLASS 3 SECTION 3.0 h! STORY OF SCE THERMAL SHIELD SUPPORT SYSTEM PERFORMANCE a

In order to support development of an understanding of the cause of the observed degradation, a review of the history of issues associated with the SONGS I thermal shield supports was performed. A chronological summary follows:

DATE ISSUE 11/66 Post hot functional test examination revealed that welds between the limiter keys and the thermal shield were cracked and all 4 of the limiter keys were " loose". Also 3 of the 12 dowel pins had cracked welds. The test ran for 3 weeks with various combinations of 1 to 3 pumps in operation. The number of pumps in operatfec. significantly affects the loads-experienced by the bolts and the resulting number of fatigue I cycles but no documentation exists to clearly establish the

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length of operating time in each condition. The original limiter keys were of the same design used in the other plant which recently had support system problems; i.e., they had close tangential gaps and a large radial gap. However, after l the hot functional test at SONGS I, the keys exhibited j indications of impacting and wear in the tangential and radial

! direction. As a result, the limiter keys were modified to limit radial rather than tangential motion and 6 flexures were l installed at the top of the thermal shield in order to restrain the motion of the thermal shield. Axially flexible rather than stiff supports were used to accommodate any differential thermal expansion between the core barrel and the thermal shield. These supports (called " flexures") were stiff 'in the radial and tangential directions in order limit vibratory amplitudes. Finally, the lower support bolts were removed, inspected and replaced and the damaged dowels were repaired.

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WESTINGHOUSE-CLASS 3

. 12/66 After completion of the indicated repairs, a second, abbreviated, hot. functional test was performed. At'least 30' l~ hours'at hot, full flow conditions plus 4-days with 2 pumps:

(during heatup)-of run: time was accumulated. There was no-indication of contact at the new limiter keys. The only indication of._ damage was a cracked weld on the right hand dowel at the 60~' degree support block _and records indicatsd that this-weld was-cracked after-the initial hot functional test but had been inadvertently missed during the initial repair.

12/08 HOT FULL POWER REACHED

- 7/70 Westinghouse advised SCE that the flexures were likely to be broken. This conclusion was based on observations from an inspection that had been performed at the other plant and a comparison of the two plants. An inspection was recommended to confirm the condition of the flexures. It was.noted that the primary basis for satisfactory operation without the' flexures was the visual exam at CY which did not reveal any adverse

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impact on the remaining components.

- 1971 BORESCOPE INSPECTION OF FLEXURES Determined that flexures at 124', 205*, 244* .and 85* were not broken. The flexures at 21* and 325' could not be inspected.

1972 FOLLOWUP BORESCOPE INSPECTION OF FLEXURES Obtained same results as the 1971 flexure inspection. Roughly 50 months or *.1 EFPY years of operation had been accumulated.

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. nu.-ommo 3-2

WESTINGHOUSE CLASS 3 9/76 5th REFUELING / INSPECTION OF FLEXURES f Visual inspections revealed that 4 flexures were broken and 2 l were intact. The intact flexures'were located at 124' and 244*.

[' Flexures at 21, 85, 205 and 325 degrees were failed. It was also reported that the lower support blocks-and limiter key welds were intact with no evidence of excessive or' unusual movement of the' thermal' shield.

The best estimate of the total . operating time at this point is1 88 months or 7.3 EFPY. The basis for this is as follows:

Number Months of-Year Full Power Operation 1968 6 1969 11

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1970 10 1971 12

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1972 11 1973 8 1974 10.6 1975 10.5 1976 8.9 9/78' 6th REFUELING / FLEXURE INSPECTION It was reported that the 244' flexure was now broken but the 124' flexure was still intact.

The best estimate of the accumulated operating time at this point is 95.8 months or 8 EFPY.

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-2...-_-__ - - _ . - - . - . - - _ _ _ . - . _ _ - . - . - . - - - . _ _ . _ _ - - - _ . - - - - . . -

WESTINGHOUSE CLASS 3 4

12/88 REFUELING / FLEXURE INSPECTION / SUPPORT BLOCK INSPECTION-

• The 124* flexure is still not broken. However, limited degradation was. observed on the upper bolts in the O' and 240' support blocks. Further details are contained in Section 5.0 of this report.

.The best estimate of the accumulated operating time at this point is 156 months or 13 EFPY.

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I l- WESTINGHOUSE CLASS 3

)

L SECTION 4.0 SOUTHERN CALIFORNIA EDISON / WESTINGHOUSE THERMA'. SHIELD ACTION PLAN INTRODUCTION After the 1987 observation of damage to the thermal shield supports at one of the two other plants which utlize a support system similar to that used at SONGS I, SCE proactively initiated an effort to determine the condition of the SONGS I thermal shield and began developing contingency plans to address the potential implications for that plant. Because of the limited storage space available in the SONGS I spent fuel pool, it was not possible to offload the entire core and perform a full inspection of the core barrel assembly. A program was in place to allow trans-shipment of fuel to other units at the same site but uncertainties in the ability to obtain the necessary approvals and execute ths program led SCE and Westinghouse to jointly develop a program for inspection of the thermal shield support blocks without removing the core barrel from the reactor vessel. This meant that a partial visual inspection program could be performed without unloading the core. The key elements of this program were a video inspection of the inside of the core barrel in the region of the lower support blocks and the flexures plus a video inspection of the condition of the flexures and partial video inspection of the outside of the lower support blocks. Since all except one of the flexures had been previously reported to be broken, the primary objective of the flexure inspection was to confirm the condition of the flexure which was reported as intact during the previous inspection. The results of this inspection are reported in Section 5.0 of this report.

Prior to the refueling outage, Southern California Edison also decided to install accelerometers on the reactor vessel seal ring flange in an attempt to determine if significant impacting was occurring within the reactor vessel. In addition, some neutron noise data was acquired prior to the outage and inspection; utilizing the ex-core neutron detectors, in order to initiate

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development of a baseline for any future monitoring programs that might be implemented. Although the amount of data that could be taken was limited, it J

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was important to obtain data corresponding to the observed condition of the thermal shield.

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WESTINGHOUSE CLASS 3 As an additional contingency prior to plant shutdown, SCE authorized Westinghouse to perform a stability analysis to determine the minimum numbcr of support blocks that are needed to ensure static and dynamic stability of the thermal shield system. It was felt to be crucial to ensure that stability

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limits were not violated in order to ensure that impact loads and displacements would never increase uncontrollably.

In order to develop estimates for the magnitude of any degradation that might exist and to be prepared to quickly evaluate the ability to resume operation with some degradation of the thermal shield supports, SCE also authorized Westinghouse to perform an analysis (prior to the outage) to determine the most probable condition of the thermal shield supports. Although any analysis has some uncertahties, a similar approach had been taken by the remaining plant which has .he same support system design and the analysis had correctly predicted that no degradation existed at that plant. As a final contingency, Southern California Edison also initiated procurement of the long lead material and developed a repair program outline that would be needed if a

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repair had to be performed prior to resuming operation.

The analytical results indiccted that even if the remaining flexure were still intact, a potential existed for thermal shield support system degradation at SONGS 1. In particular, it was predicted that degradation was expected at the 240' and 300* support blocks. This expectation was confirmed by the results obtained during an inspection of the thermal shield support system during the current outage where degradation was observed at the 240* block and inferred at the O' block. Although this was not perfect correlation ,it was sufficient to inspire high confidence in the analysis and in fact, further analyses showed that if the upper dist tement limiter keys have worn, then the 0 and 240' blocks should be degradeu but the 300* block might not be.

Because of the constraints previously outlined relative to spent fuel storage, SCE and Westinghouse decided to determine whether it was necessary to repair the support system at this time or whether a repair could be deferred without adversely affecting safety.

i 3586s-020289 10 p,g

WESTINLiHOUSE CLASS 3 In order to. determine if a repair of the thermal shield support system at SONGS Unit 1 can be deferred until the end of the upcoming fuel cycle, an

' analysis and evaluation program was performed to demonstrate the ability of the plant to resume safe operation for an additional cycle without effecting a -

repair. This plan is summarized in Table 4.0-1.

BASIS FOR THE PROGRAM The program addressed the safety implications of continued plant operation with the observed thermal shield support system degradation and also addressed postulated worst case scenarios with respect to damage to the support block fasteners and wear at the upper limiters since the visual inspection does not necessarily provide absolutely conclusive evidence relative to the condition of the fasteners in the thermal shield support system.

The overall approach used for the evaluation program was:

1) Determine the loads acting on the fasteners and use this to determine the most probable condition of the support system.

2) Evaluate the structural integrity of the thermal shield, core barrel and the thermal shield support system during plant operation under.

normal and seismic conditions when the support system is degraded.

3) Determine the structural consequences of unexpected, postulated failure of the thermal shield support system.

4) Prepare a licensing and safety evaluation to justify continued operation for one fuel cycle in the degraded condition.

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WESTINGHOUSE CLASS 3

SUMMARY

OF TASKS PERFORMED IN SUPPORT OF THIS PROGRAM' k

A listing. of the analyses and/or evaluations that were performed to support _

the preparation of the justification for continued operation follows: l

.o Vibration analyses for degraded support conditions o' Structural analysis for postulated failure of support system (Thermal shield drop analysis) o Loose parts analysis o Seismic evaluation with degraded thermal shield supports o Scoping evaluation of thermal transients o Hydraulic effects of tilted or dropped thermal shield o Evaluation of remaining thermal. shield flexure o Summarize Stability analysis o Summarize Vibration and structural analysis o Evaluate impact of thermal shield against core barrel o Evaluate Core Barrel Structural integrity o Evaluate support block function with degraded fasteners SCE will also implement methods of monitoring the condition of the thermal shield in order to enhance control of the issue and obtain an early indication of dramatic changes in thermal shield behavior in order to minimize the economic risk that would accompany failure of the support system.

me.-e2cas io 4_4

WESTINGHOUSE CLASS 3 TABLE 4.0-1 ACTION PLAN FOR SCE THERMAL SHIELD ISSUE C

GOAL: DEVELOP A JUSTIFICATION FOR CONTINUED OPERATION FOR THE NEXT FUEL CYCLE (18 MONTHS)

ACTIONS:

1. Use analysis to demonstrate acceptability for one fuel cycle.

o What further degradation might occur and is it acceptable?

o Evaluate effect of normal vibratory, thermal and seismic loads on structural integrity o Demonstrate that shield will be adequately supported and will not tilt or fall.

2. Determine Consequences of Postulated Failure of support system o Evaluate structural consequences of Tilting or drop of shield Detectability Determine if the loose parts and vibration monitoring program can detect substantial degradation of the support system.

3. Implement a monitoring Program o Neutron noise monitoring o Loose Parts monitoring o Fuel Damage / Coolant Activity

5. Perform Loose Parts Analysis

6. Inspect Bottom of Vessel

, 7. Develop 10CFR50.59 Safety Evaluation and JC0 O

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WESTINGHOUSE CLASS 3 SECTION 5.0 OBSERVED DEGRADATION

\

Because of the inability to fully unload the reactor core, it was not possible to remove the lower internals and perform a full examination of all the thermal shield supports with the core barrel installed. Therefore, an ,

inspection program and the required tooling was developed to facilitate an inspection of the upper and lower thermal shield supports. Further, since a l gross indication of the condition of the thermal shield supports could be obtained by inspecting non-support components that would likely be damaged by excessive thermal shield motion, seven of the eight irradiation specimen holder tubes were also included in the program. At the Haddam Neck plant, severe degradation occurred to some of the specimen holders as a result of the relative motion between the core barrel and the thermal shield. Although the specimen holders in that plant are more susceptible to flow induced vibration damage since they are welded to both the barrel and the shield, whereas the SONGS holders employ a spigot fit between the two sections (the lower part is welded to the thermal shield and the upper part is welded to the core barrel);

very large amplitude thermal shield vibration would still be expected to lead to damage since the clearance between the two parts of the holder is only 0.1 inches. Visual examinations of the outside of ;i:e support blocks were conducted by inserting cameras down through holes in the core barrel flange.

These holes are located at 45*, 135*, 225* and 315* and are normally used to attach the lift rig to the internals. The through diameter of these holes is only 2.85 inches,'and this limits the access available for inspection tooling. Nevertheless, it was possible to visually inspect the outside of the entire 240* block and to perform a limited outside inspection of the 60*, 120*

and 300* support blocks. It was not possible to observe the outside of the 0 and 180* support blocks because access to these areas from the core barrel lift holes is not possible. In addition, it was possible to examine the 124*,

205*, 244*, 325' and 21* flexures. It was not possible to examine the 85*

flexure, but since it was known that this flexure was previously broken, this

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was not a shortcoming. Finally, it was possible to perform a visual examination of the outside of five of the irradiation specimen holder tubes plus an interior examination of three of the tubes.

nu,-enne io 5-1

WESTINGHOUSE CLASS 3 A second part of the inspection program was to insert a camera down through

-the 2.31 inch diameter holes in the lower core plate in order to examine the

- inside of the core barrel and the back of all the bolts plus the lower dowel pins at all six (6) of the support blocks. This was accomplished by -

selectively moving a few fuel assemblies which are located near the block-locations in order to gain access.

In addition, a visual examination was performed of the inside of the core barrel in the vicinity of each of the flexure supports in order to confirm.

that the bolts between the flexure and the core barrel were intact. Where access permitted, a similar inspection was performed on the outside of the core barrel at the flexure locations. The results of the inspection program are summarized as follows:

Exterior Inspection o Flexures at 21, 205*, 244' and 325' were confirmed to be cracked or broken.

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o .The flexure at 124' was confirmed to be. intact. It was also confirmed that the attachments between the flexure, core barrel and thermal shield were intact.

o At the 240* support block, the top left-hand bolt appears to have a broken lock weld and a portion of region around the lock weld appears to be missing. Also at the 240' block, markings were observed on one of the dowel pins which might indicate a cracked weld. However, it could not be determined if this was real or some optical effect. It should be noted that the analysis which was performed prior to the inspection indicated that even if the 124' flexure were still intact, damage could be expected t at the 240' block. I i- o The other support blocks could not be directly viewed from the outside but those which were viewed from an angle did not appear to have bolts or dowel pins protruding (as in the other domestic plant) and they showed no evidence of gaps between the block and the thermal shield or core barrel.

nu.-e2o2n te 5-2 L__-_--____-_-_______-_________________

i WESTINGHOUSE CLASS 3 o- A camera inspection of the bottom of the vessel was performed at the 60',

120', 240' locations and the center of the reactor vessel. Nothing was found which could have come from thermal shield support degradation.

This was quite a contrast from the other plant where bolts, bolt -

l~ fragments and pieces of the irradiation specimen holder were found in the bottom of the vessel.

Interior Inspection The interior inspection at the 60',120',180' and 300* did not reveal any evidence of degradation. The primary observation that is used to make this determination is the relative amount of protrusion of the ends of the bolts into the inside of the core barrel. A tolerance stackup was performed and used to determine the difference in length that could exist given the tolerances on bolt length, core barrel, thermal shield and support block thickness. If some bolts were protruding more than the other bolts, especially by an amount that exceeded tolerances, then that bolt could be broken and in the process of rotating in toward the center of the core barrel.

At the O' block, the middle (hidden) bolt is clearly broken since it is protruding in roughly 1 1/2 inches (15 threads). The other bolts do not exhibit any visual evidence of being broken, but the evaluation assumes that all the top bolts are broken since failure of the hidden bolts results in increased loads on the remaining top bolts.

Similarly, at the 240' block, the middle bolts appears to be protruding inward 1/2"-3/4" and the right top bolt is inboard by 10-12 threads (3/4"-1"). Thus the evaluation also assumes that all three top bolts in this block are also broken.

Based on the observations, analysis and the experience at the other plant, no failures of any of the lower bolts are presumed. It should also be noted that the ends of all of the lower dowel pins were observed and they appear to be recessed by the correct amount; none have moved outward. The ends of the

~

upper dowel pins are not inspectable, but based on the exterior observation at the most obviously degraded block (240*), it is reasonable to conclude that the remaining dowel pins are also still in place.

nu.- """ 5-3

+e WESTINGHOUSE CLASS 3-A summary of the interior observations follows:

0-SCE-THERMAL SHIELD INSPECTION

SUMMARY

ITEM COMPONENT CONDITION 0' Block Center Bolt Inboard approx. 15 Threads (1 1/2 ")

lop Long Bolts OK Lowet Bolts OK Lower Dowels OK 60' Block Center Bolt OK Top Long Bolts OK Lower Bolts OK Lower Dowels OK

~

120* Block Center Bolt OK Top Long Bolts OK Lower Bolts OK Lower Dowels OK 180* Block Center Bcit Inboard approx. 4-6 Threads (5/8-3/4")-0K Top Long Bolts Inboard 1/4-3/8"-0K Lower Bolts Inboard same amount as Center Bolt-0K Lower Dowels OK 240* Block Center Bolt Inboard approx. 1/2-3/4 "

Top Long Bolts Right inboard 10-12 Threads (3/4-1")

Lef t Bolt - OK Lower Bolts OK Lower Dowels OK 3568s-020280 10 5-4

__ _a

WESTINGHOUSE CLASS 3

300* Bloc'k Center Bolt OK' Top Long Bolts OK Lower Bolts OK Lower Dowels OK' h

TV Inspection of Bottom of Reactor Vessel o inserted Camera at 60, 120, 240 and Center; No support block hardware observed.

CONCLUSION:

0* and 240' Blocks are Degraded. Assume Top 3 bolts are failed.

Lower Bolts are intact.

Remaining four blocks also appear intact, undegraded.

Inspected inside bolts at all 6 flexures; undegraded.

O .

O nu.-emo2u,o 5-5 ,

i

1 WESTINGHOUSE CLASS 3 SECTION 6.0 ANALYSIS TO DETERMINE CAUSE OF DEGRADATION Prior to the inspection program, analyses were performed to determine the i*

loads acting on the thermal shield supports in an attempt to determine if

]

degradation was likely. As previously reported, those analyses indicated that ]

degradation might exist at two of the support block locations if the remaining  !

1 flexure was intact. Following the inspection, further analyses were performed on a best estimate basis to explain the current observed condition of the thermal shield lower support blocks. Analyses were also performed to predict the expected progression of the thermal shield degradation over the next fuel cycle (18 months).

The evaluation of the plant operating history on the degradation of the lower support blocks focused mainly on the vibratory loads experienced by the thermal shield. To evaluate these vibratory loads and their effect on the thermal shield supports detailed analytical models were generated using finite elements. The vibrational models were run for a number of different support boundary conditions to understand the history and to predict the expected extent of the current thermal shield support degradation. Once the model was baselined with the history of the plant operation and to the observed current degradation, a prediction of the expected future degradation to the thermal shield supports over the next 18 months of operation was made.

In support of the vibrational work, a vortex shedding analysis and thermal evaluation were performed to eliminate any concern that another mechanism may have been the cause of the observed lower support degradation.

The following sections will discuss in more detail the analyses and evaluations performed. Also, the expected extent of any future degradation of the thermal shield supports will be presented and commented on.

e nu.-onne io 6-1

WESTINGHOUSE CLASS 3 6.1 Vortex-Shedding Vortex-shedding was considered because of some concerns in the literature to the effect that trailing edge vortex-shedding was a viable mechanism for excitation of thermal shield modes (Reference 6.1-1) and also because of data (Reference 6.1-2) which showed the presence of some higher order thermal shield modes at frequencies approximately equal to the trailing edge vortex-shedding frequencies. To address this issue, an ACSTIC (Reference 6.3-3) model of the fluid on both sides of the thermal shield was created (Figure 6.1-1) and used to calculate acoustically-induced pressure differences across the thermal shield for typical trailing edge vortex-shedding frequencies. The calculational model used trailing edge vortex-shedding pressure differences and Strouhal numbers obtained from Kennison (Reference 6.1-4). Kennison inferred the lift coefficients and Strouhal numbers for a thermal shield based on data for cylinders in crossflow and scaling laws to allow extrapolation from cylinders to trailing edges. These pressr e differences were distributed azimuthally in such a way as to make shell mode

~

excitation most likely, i.e., distributed as a full sine wave over a full azimuthal wave of a thermal shield shell mode. The magnitude of the trailing

~

r -

edge delta p was determined (Reference 6.1-4) to be L . A map

_ e.

of delta p fluctuation amplitudes across the thermal shield for a beam m,eode type distribution of trailing edge pressure fluctuations is shown in Figure 6.1-2. Amplitudes with shell mode type distributions are even lower.

It can be seen that the pressure difference amplitudes decrease rapidly as distance from the thermal shield trailing edge increases. The reason for this l

can best be understood by considering the acoustic behavior of the fluid. For  !

the model shown in Figure 6.1-1, the acoustic mode frequency, based on the i full length of the thermal shield and an n=2 thermal shield mode distribution

- of pressure fluctuations at the trailing edge is estimated to be on the order of [ ]. For higher order modes, this frequency will be higher because the dominant characteristic length will be the circumferential wave length of the mode. Clearly, with frequencies of this order of magnitude, there is little possibilityofanacousticalmatchupwiththe[ ]wevortex-shedding

~

frequency. The possibility of a matchup with the dominant structural 1 frequencies of[ ]is clearly even less. j b,e l

1 nu,-anem 6-2 j i

WESTINGHOUSE CLASS 3 In addition to this calculation, there are other reasons why trailing edge vortex-shedding should not be a significant effect as far as thermal shield a excitation is concerned. These are discussed below:

1. Correlation Length In order for vortex-shedding to excite the thermal shield, vortices have to be correlated at the trailing edge in the "up-down" fashion discussed earlier. For this to happen on a fluid-mechanical basis alone is implausible; for cylinders in crossflow, correlation lengths of only a few diameters are typical for stationary cylinders. For the SCE thermal shield trailing edge, this amounts to a few multiples of 2.5 inches, the Larmal shield thickness.

Thus, given the thermal shield circumference of about 413 inches, or even a fraction of 413 inches to allow for the shorter wave lengths of higher modes, this is unlikely.

It is also unlikely that coherent motions of the thermal shield could

~

correlate the trailing edge. vortices. Results of tests with cylinders (Reference 6.1-3) indicate that such " lock-on" does not occur until vibration amplitudes of 10-20% of the cylinder diameter are exceeded.

For the trailing edge of a thermal shield, cylinder diameter is equivalent to shield thickness, 10% of which is 250 mils. Thermal shield vibration amplitudes of this magnitude at the support block level are considered to be rather unlikely.

2. Geometry Well-defined vortex-shedding from blunt bodies like cylinders and thermal shield occurs when the free stream is downstream and is relatively free of obstructions. If, however, one places a splitter plate from the trailing stagnation point of a cylinder, oriented along the mainstream direction, " shedding" can be stopped. What happens is that vortex-switching from one side of the cylinder to the other cannot occur and the result is two steady-state recirculation zones on either side of the nu.-on2"

• 6-3

\. . .. .. . . .

WESTINGHOUSE CLASS 3 splitter plate. 'While there is no splitter plate on the SCE. thermal shield, the confinement presented by the the core barrel and vessel act l- in much'the same way. On a strictly geometrical basis, the core-barrel-thermal shield-vessel annulus resembles a single sudden expansion more than it does a blunt body in an-infinite free stream.

l

{

To summarize, vortex-shedding is not considered to be a mechanism which will  ;

cause significant vibration of the thermal shield in SCE. Acoustical calculations using Strouhal numbers and lift coefficients from the literature indicate that the induced pressure differences across the thermal shield at trailingedgevortex-sheddingfrequencies[' ]gli be small. The conditions assumed in this analysis are conservative in that they postulate vortex-shedding correlation along the trailing edge which is extremely unlikely, given the fact that thermal shield amplitudes are not high enough to produce such correlation.

Finally, the confinement in the barrel-shield-vessel region is such that vortex-shedding, if it takes place, will be greatly diminished in magnitude as compared to a more " free" condition. In fact, it is expected that the flow leaving the thermal shield will behave more like a sudden expansion than the wake of a single cylinder in crossflow.

REFERENCES 6.1-1 Fabic, S., " Investigation of Methods for Coupled Structural Hydrodynamic Analysis of Reactor Internals," Proceedings of the Conference on Flow-Induced Vibrations in Reactor System Components, ANL-7685, pp. 290-303.

6.1-2 Singleton, N. R., et al, "Four Loop PWR Internals Assurance and Test Program," WCAP-7879, July 1972.

6.1-3 Schwirian, R. E., ACSTIC Ccmputer Program User's Manual, August, 1983.

1 I

m e.-o m enio 6-4

WESTINGHOUSE CLASS 3-

/

6.1-4 Kennison, R. G., " Theoretical Analysis of the Flow-Induced Vibration of Flat Plates Caused by Trailing Edge Vortex-Shedding," KAPL-M-7198, July

• 15, 1971.

6.'l-5 Griffin, 0., " Vibrations and Flow-Induced Forces Caused by Vortex-Shedding," from " Excitation and Vibration of Bluff Bodies in Cr assflow, Vol.1 ( ASME),1984, pp.1-13.

O

na e l e

'"**-"2"  !

6-5

WESTINGHOUSE CLASS 3 i

1 i

1 2

b,*

I Figure 6.1-1. Vortex-Shedding / Acoustic Model l

nes.-cana ' 6-6

3 i

WESTINGHOUSE CLASS 3 1 l

\

)

i i

i i

l i

I l

i  !'

l f b,s Figure s 1-?. Calculated Thermal Shield Pressure Difference Amplitudes for Beam Mode Type Trailing Edge Vortex-Shedding Distribution ms.-czo2se so g_y

WESTINGHOUSE CLASS 3 6.2 Vibration' Analysis For a structure under a turbulent, random, flow field, the vibrational amplitude for.each mode of the structure can be calculated using the following -

equation where the~mean squared displacement and the pressure-spectral density are related by':

(6.2-1) where l

L 1

be y

l, 1. To determine the natural frequencies (fn), and the associated eigenvector ($n), and modal stiffness (kgn) of the thermal shield-core barrel structure, a finite element model of the thermal nu.-aane io 6-8 1

. WESTINGHOUSE CLASS 3 shield and the core barrel was developed. This model, shown in the following figure, included the full length of both the thermal shield and the core barrel. The lower thermal shield supports, the limiter keys and the flexures at the top of the thermal shield were also included at their respective locations around'the circumference. The modal stiffness and the natural frequencies and the eigenvector for each mode were determined.

2. A turbulent vibration analysis was performed using pressure spectral density (PSD) 6 eting functions which are based on theoretical, random vibration forn.8c, and are normalized by comparisons of predicted and measured modal amplitudes for a reference plant. These forcing functions were scaled to the SONGS plant conditions and were used in conjunction with the modal results to determine the vibrational response of the core barrel and thermal shield.

3. Based on the results of Items 1 and 2 above, the forces in the support 4 blocks, bolts and the displacements of the flexures and the limiter keys were determined.

The following paragraphs describe the above approach in more detail:

1. Model Development 0%C l

l e

0 3

me.-o2o2 e io 6-9

~

~ WESTINGHOUSE' CLASS'3 i*

C Q.o-  !

j

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'2. Random (Turbulent)ExcitationAnalyses A turbulent excitation analysis of the SCE reactor plant is complicated at the outset by the. absence of any specific information on thermal shield displacement amplitudes and pressure spectral density (PSD) functions for this plant. However, such information is available from test' data obtained in plant and model tests of a reference plant (References 6.2-1 and 6.2-2). The approach, therefore, was to use these reference plant data to calibrate a forcing function (PSD) model, which could then be extrapolated to SCE conditions via appropriate scaling laws. This involves development of finite element models of both the reference plant and SCE, as well as the development of the appropriate forcing functions. The procedure is as follows:

1

1) Create finite element models of reference plant core barrel-thermal shield system.

2) Use test data on pressure spectral densities, correlation lengths, and attenuation factors, to define appropriate forcing functions, j and apply to the reference plant finite element model to calculate

)

displacement amplitudes.

]

1 me.-un io 6-10

l l

~

[ ' WESTINGHOUSE CLASS 3

]

~

23) Calibrate thelforcing function model'so that the calculations match-the : test ' data amplitudes.

4)- Scale theLforcing function model to SCE conditions. -

4

5) Apply forcing function model to SCE finite element model to calculate modal amplitudes.

l- Calculate. modal force and' displacement amplitudes at the ' support 6) blocks and radial keys. Combine using the square root of sum of the squares (SRSS) to get peak forces and amplitudes.

~

7) Evaluate loads on support block bolts and blocks using the results of step (6).

This overall development is outlined below.

~'

Pressure Spectral Densities (PSD)

~

The forcing function for the analysis are pressure spectral density (PSD) spectra obtained from reference plant data (Reference 6.2-1) at the top of the thermal shield. These were taken on each side of the thermal shield and thus represent' spectra for the two (outside and inside) flow channels. For each mode, the rms amplitudes were calculated using Equation 6.2-1 in a discretized form to match the finite element model.

a3C (6.2.2) 1 l

El nu. on2n in 6-11

(

WESTINGHOUSE CLASS 3

/c.

where 6c.

4 From the measured pressure spectral densities of Reference (6.2-1), an empirical formula for the pressure spectral density in any reactor downcomer annulus was derived. This formula matches the amplitudes and frequencies at the peaks of the "inside" and "outside" test spectra from the reference plant test data. At high frequencies it also obeys the power dependence of the spectrum on reduced frequency'which has been observed in Reference 6.2-1 and other tests. The variables in this formula are the flow dynamic pressure, the reduced frequency based on the gap, and the downstream to upstream gap ratio, that is, the ratio of the

~

~ flow annulus gap at the point on the thermal shield where the PSD is nu.-omie io 6-12 L

v ~

WESTINGHOUSE CLASS 3 desired to.the. flow annulus gap above the thermal shield. These are:

a considered to be the primary variables on which PSD should depend. 'The form'of-this empirical. formula is:

a,c

. The approach described here assumes the spectra acting on the thermal ,,.

shield are initially generated at the inlet nozzles and are modified and

. attenuated as'the flow proceeds downward. Self generated wall turbulence

-in the flow annuli is considered to be negligible.

Correlat' ion Lengths The weighting factors C$ in Equation (6.2-1) depends on correlation length. The correlation lengthe used in the present analysis are based on Au-Yang's data (Reference 6.2-3) and DeSanto's observations (Reference 6.2-4). With this formulation, correlation length is inversely proportional to frequency at low frequencies and approaches an asymptote _ _6;a ,c at high frequencies. The azimuthal correlation length is approximately_

times larger than the axial correlation length.

e me.-ensio 6-13

WESTINGHOUSE CLASS 3 l

i Attenuation of Turbulence

_ b;Cp

'O ~~

This observation was used, together with Taylor's analysis (Reference l

6.2-5) for the decay of turbulence downstream of a grid, to obtain a continuously varying attenuation factor over the entire thermal shield.

While Taylor's approach is based on grid generated turbulence, it is expected to be applicable to the present case because the decay of eddies should be dependent primarily on their initial size, and not the means used to create them.

Reference Plant Calculation / Test Data Comparison There are two sources of reference plant modal amplitude data, plant tests

~

(Reference 6.2-1) and scale model tests (Reference 6.2-2). The two sets of data generally have the same overall behavior, but there are some differences. In addition, the core barrel beam mode data were considered reliable only in the scale model test because of the unknown effect of the radial keys on the beam mode amplitudes. Table 6.2-1 summarizes the plant and scale model data for what are considered to be the four most significant types of modes, the core barrel beam modes, thermal shield n=2 modes, thermal shield n=3 modes, and core barrel n=2 mode. The small contribution of the reactor coolt.u pump (RCP) to the core barrel n=2 mode is also shown for completeness.

To calibrate the calculational model, what was considered the best test datum was used. This is the displacement amplitude of the core barrel beam mode measured in the scale model tests. Table 6.2-1 shows a

_ _ ta C calculated val,ue of[ _

mils for this mode, as opposed to a measured value of__ 1 mils. The calculation is, therefore, conservative. The othermodalamplitudeswerealsocalculatedandgreshownasCase1in Table 6.2-1. Except for the mobe,thesecalculations agree reasonably well with one or the other of the plant or scale model test data.

asa .-ox a io 6-14

, WESTINGHOUSE CLASS.3 i

3. Results of Vibration Analysis The results of 'the' vibration analysis at.various configurations are presented in Table 6.2-2. This table is arranged in the chronological order from left to right to show the sequence of degrading.

-In the table, the case number represents the numbar of. flexures, the number of limiter keys and the number of support blocks. For. example, Case 246 represents the condition that there are two flexures, four unworn limiter _ keys and six undegraded support blocks, while Case 104 represents the condition that there'are one remaining flexure, four worn keys and four undegraded support blocks.

The. frequencies of the more important modes, namely the core barrel and the thermal shield beam modes and the thermal shield n=2 shell modes are tabulated for each case. The rest of the table is self explanatory.

Note that the results, such as the loads for the support blocks and the displacements at tne limiter keys and the flexures, presented in this table are obtained based on an assumed forcing function. A factor to be discussed in the next section adjusts the assumed forcing function to the more realistic plant conditions. The adjusted loads and displacements are then used to calculate the fatigue usage factor for the various bolts.

4 e

m e.-oroneio 6-15

WESTINGHOUSE' CLASS 3

- REFERENCES O 6.2-1 Singleton, N.R., et al, "Four Loop PWR. Internals Assurance and Test Program." WCAP-7879, July, 1972. -

6.2-2 Lee, H., " Prediction of the Flow-Induced Vibration of Reactor Internals by Scale Model Tests," WCAP 8303, March, 1974.

6.2-3 Au-Yang, M. K., Nuclear Engineering and Design, Vol. 58(1980),pp.

113-125.

6.2-4 DeSanto, D. F., " Turbulence Pressure Forcing Function Acting on a PWR Core Barrel," 83-1E7-DYNRI-R1, January 9, 1984.

6.2-5 Bird, R. S, et al, " Transport Phenomena," John Wiley & Sons, Inc.,

1960, pp. 173-174.

e e

O m e.-o m aeio 6-16 L___-_____-___________________----_

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Figure 6.2-1. Reference Plant and SCE System Models is...-e:c w o 6-18

WESTINGHOUSE CLASS 3

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- Figure 6.2-2. Reference Plant and SCE Support Models l

n u.-on a.io l 6-19 l

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_. ._. ._______.___Q

WESTINGHOUSE CLASS 3 6.3 Thermal-Transient Assessment 6.3.1 Objective i

l The objective of this thermal transient assessment was two-fold. First, it was performed to identify a set of pNry coolant system thermal transient events that could be considered representative for this scoping study of the San Onofre Unit No. 1 (SCE) nuclear power station. Second, these transients were used to estimate time history differences between average component temperatures of the SCE core barrel and thermal shield.

6.3.2 Technical Approach Basically, the approach taken to perform the assessment involved identifying specific fluid system thermal transient events, determining the time history impact of these events on the SCE thermal shield and core barrel average component temperatures separately, then deriving the corresponding time history differences between these component average temperatures.

. 6.3.2.1 Identification of Transient Events Initially, attempts were made to identify design bases thermal transient events specific to SCE. This was done by reviewing the currently applicable Final Engineering Report and Safety Analysis (Reference 6.3-1) for SCE, and the SCE Reactor Vessel Equipment Specification (Reference 6.3-2).

Unfortunately, neither document identified applicable thermal transient events. At this point, it was decided to use the normal and upset design bases thermal transient events from Reference 6.3-3. The terms " normal" and

" upset" are defined as follows per the ASME Code (Reference 6.3-4).

Normal Conditions Any condition in the courses of system startup, operation in the design power range, hot standby and system shutdown, other than Upset, Emergency, Faulted or Testing conditions.

l 3568s-020289 10 g.gg

WESTINGHOUSE CLASS 3 Upset Conditions Any deviations from Normal Conditions anticipated to occur often enough that design should include a capability to withstand the conditions without operational impairment. The Upset Conditions include those transients which result from any single operator error or control malfunction, transients caused by a fault in a system component requiring its isolation from the system, and transients due to loss of load or power. Upset Conditions include any abnormal incidents not resulting in forced outage and also forced outages for which the corrective action does not include any repair of mechanical damage. The estimated duration of an Upset Condition shall be included in the design specification.

Additionally, the September 1988 structural analysis c# the SCE steam generator feedwater nozzle (Reference 6.3-5) identified specific transient events used as an analysis basis. The transient events identified by References 6.3-3 and 6.3-5 are apparently the same, except the latter

. identifies a Feedwater Cycling transient, and the former identifies an Inadvertent Auxiliary Spray, Per Reference 6.3-3, however, the Inadvertent

. Auxiliary Spray evt.nt is not accompanied by a temperature change in the primary coolant system with the exception of the pressurizer itself.

Regarding the Feedwater Cycling transient, the two references do not provide a primary coolant temperature transient description. Therefore, Reference 6.3-6 was used to obtain this description. Primary coolant temperature descriptions for all other transient events considered were identified in Reference 6.3-3.

Table 6.3-1 below identifies the thermal transient events that were used in this scoping thermal assessment. Also, this table indicates the estimated number of occurrences based on a 40 year plant design life, and the estimated number of occurrences on a yearly basis.

e nu.-oxne' 6-27

WESTINGHOUSE CLASS 3 i

TABLE 6.3-1 THERMAL TRANSIENT EVENTS AND ESTIMATED OCCURRENCES No. Design Life

• No. Yearly Transient Event Occurrences Occurrences Heatup (to no-load condition 200 5 Cooldown (from no-load. condition) 200 5' Unit Loading (at 5% Full Power / min) 18,300 458 Unit Unioading (at 5% Full Pwr/ min) 18,300 458 Step Load Increase (90-100% Full Pwr) 2,000 50 Step Load Decrease (100-90% Full Pwr) 2,000 50 Large Step Decrease (100-5% Full Pwr): 200 5 Feedwater Cycling (at no-load cond.) 20,000 500 Loss of Load 80 2 Loss of Power 40 1 Loss of Flow (in one loop only) 80 2 Reactor Trip 400 10

. Inadvertent Auxiliary Spray 10 1/4 (inpressurizer)

-

• Based on a 40 year design life.

e me,-exue in 6-28

WESTINGHOUSE CLASS 3 At this point it was necessary to estimate a more realistic number of l

occurrences of each transient event corresponding to the time period between l- plant startup.and the end of July 1990 (i.e., 18 months after February 1989

~

corresponding to the next scheduled fuel cycle). This was accomplished by using the following SCE operational history from Reference 6.3-5:

TABLE 6.3-2' SAN ONOFRE OPERATIONAL HISTORY: 1974 - 1988 Year (19--) Months Operating Months at Cold Shutdown) 74 10.6 1.4 75 10.5 1.5 76 8.9 3.1 77 7.8 4.2 78 9.7 2.3 79 10.9 1.1 80 3.3 8.7

. 81 4.5 7.5-82 2.0 10.0 83 0.0 12.0 84 1.1 10.9 85 9.3 2.7 86 4.9 7.1 87 10.1 1.9 7/88 1.5 5.5 Totals 95.1 79.9 e

n u.-o w ne to 6-29

WESTINGHOUSE CLASS'3 l Based on this data, SCE maintained operation on an average of 54.3% of the calendar time. During.1979 SCE maintained their highest yearly operational percentage of 91%. This information was used as follows to estimate a more

' realistic number of transient event occurrences for use with the results of 1

. \

- this assessment: {

.' 1 No.' Occurrences per = No. Occurrences + No. Occurrences Event (N0PE) 1968 - January 1989 February 1989 - July 1990 (259 Months) (241 months) (18 Months)

NOPE = No. Design Occurrences

• 241 * (.543) + 18 * (.91)

Per Year 12 12 N0PE = No. Design Occurrences

• 12.2703 Per Year The results of this estimation appear in Table 6.3-3.

6.3.2.2 Estimation of Transient Component Temperatures Transient thermal shield and core barrel average temperatures were estimated corresponding to those transient events identified in Table 6.3-3. A simple-finite difference algorithm (Reference 7) was used to create separate linear finite difference models of the thermal shield and core barrel. The transient primary coolant cold leg temperature profiles were used in conjunction with these models in order to generate time history average component temperature data.

One particular area of uncertainty involves thermal energy generated within the thermal shield and core barrel due to gamma ray absorption. To date, a comprehensive analysis has not been performed for SCE. For the purposes of this scoping assessment, a one-dimensional discrete oridinates transport calculation has been performed to estimate the gamma ray heat generation in the SCE core barrel and thermal shield.]

O

~ bsC me.-enas to 6-30 .

- -__M--___m__.____m___ _ _ _ _ _ _ _ _ ... ' _ _

. WESTINGHOUSE CLASS 3

' TABLE'6.3-3 TRANSIENT EVENTS AND ESTIMATED NUMBER OF OCCURRENCES FOR USE IN SCE THERMAL SHIELD SCOPING STUDIES Estimated No. of Cycles *

!- Transient Event (1968throughJuly1990)

Heatup (to no-load condition) 61' Cooldown(fromno-loadcondition) 61 Unit Loading (at 5% fu11 power / min'.) 5620 Unit. Unloading (at 5% full power / min.) '5620 Step Load. Increase (90%-100% full power) 614 Step Load' Decrease (100%-90% full power) 614

'Large Step Decrease (100%-5% full power) 61.

Feedwater. Cycling (at no-load condition) 6136 Loss of Load 25

~

Loss of Power 13 Loss of Flow-(one loop only) 25 Reactor Trip 123 Inadvertent Auxiliary Spray 4

• All fractional results rounded up.

4 '.

nu.-omm ia 6-31

WESTINGHOUSE CLASS 3

~6.3.3 Results The impact of core barrel / thermal shield temperature differences on the subject structural assessments is only significant when the average thermal shield temperature at any time is greater than the average core barrel temperature. From the set of thermal transient events identified in Table  !

6.3-3, only Feedwater Cycling, Large Step Decrease (100% to 5% full power), '

Loss of Power, Loss of Flow, and Reactor Trip would result in the thermal shield being hotter than the core barrel. The Loss of Flow and Reactor Trip transient events are the two most severe; therefore, only Figures 6.3-1 and 6.3-2 are presented for them as worst-case results.

. Figure 6.3-1 shows the core barrel minus thermal shield average temperature differences estimated for the loss of flow transient. After 140 seconds from event initiation, the plant would be returned to a no-load condition, thus reducing any average temperature difference between the components. However, before that would occur, the thermal shield could become as much as about

~~

-hotter than the core barrel.

_ 3,c Figure 6.3-2 shows the core barrel minus thermal shield average temperature differences estimatsd for the reactor trip event. After 100 seconds from event initiation, the plant would be returned to a no-load condition, thus reducing any average temp eature difference between the components. However, before that would occur, the thermal shield could also becomo as much as about

~

-hotter than the core barrel.

Conclusion The analysis predicts that during steady state (100% load) conditions, the average temperature difference between the core barrel and tharmal shield is approximately] ]with the cors barrel bei.ng hotter. This condition with thecorebarrelhottertSanthethermalshieldwillhelpinmaintaininga tight joint at the lower supports. Although the bolt preload could be reduced slightly, the condition is considered to be in a beneficial direction.

1 The largest average temperature difference reported between the core barrel and thermal shield that would place a load on the lower support top visible n u .-om anto 6-32

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l WESTINGHOUSE CLASS 3 l 1

7 l bolts is about-- l ',c.eThis e average temperature difference is expect to produce a radial load at the support blocks of approximately4 wellbelowthejointpreloadofapproximately[__s [1 k.Thisis Therefore, this  !

additional thermal force is not expected to contribute significantly to the bolt fatigue loadings.

REFERENCES 6.3-1 " Final Engineering Report and Safety Analysis: San Onofre Nuclear Generating Station Unit 1," Southern California Edison Company / San Diego Gas and Electric Company.

6.3-2 E-Spec 569259, Rev. 4, "SCE Project Reactor Vessel," 5/14/64, L. R. Katz/H. J. Von Hollen (Original Issue Data: 1/23/61).

6.3-3 SSDC 1.3, Rev.1, " Systems Standard Design Criteria - Nuclear Steam Supply System Design Transients," 4/2/71, R. K. Lindefelt (W

, Proprietary).

. 6.3-4 ASME Boiler and Pressure Vessel Code,Section III, NB3113 and NB3114, July 1971 Edition.

i 6.3-5 WCAP-?.1947, Rev. O, " Structural Analysis of San Onofre Unit 1 Steam Generator 'B' Feedwater Nozzle with Deformed Thermal Sleeve,"

September 1988, R. A. Taylor and A. L. Thurman (W Proprietary).

6.3-6 SSDC 1.3 (Rev. 2, " Systems Standard Design Criteria - Nuclear Steam Supply System Design Transients," 4/15/74, W. R. Snyder (W Proprietary).

6.3-7 MED-RPV-1947, " User Guide for TCART," 7/19/88, K. B. Neubert (W Proprietary).

~

6.3-8 PSE-REA-89/247,i' Estimate of San Onofre Unit No.1 Core Barrel and Thermal Shield Heat Generation Rates", January 18, 1989, S.L. Anderson, (Westinghouse Proprietary).

mwozone to g_39

WESTINGHOUSE CLASS 3 6.4 Structural Analysis and Explanation of History 6.4.1 Structural Analysis The focus of the structural analysis was on the high cycle fatigue evaluation of the thermal shield lower support top visible bolts. The top visible bolts were judged to be the most susceptible to the thermal shield vibratory loads.

These bolts are subjected to the direct loads resulting from the vibration of shield in both the vertical and radial direction. Also, failure of these bolts would lead to a loss of clamping force at the support block, since these bolts clamp the thermal shield, support block and core barrel together.

The vibration analysis discussed in the previous section gives both vertical and radial deflections and loads at the lower support blocks. To evaluate the vibratory vertical loads, the top visible bolts are idealized as a guide cantilever beam. The vertical displacements are applied to the beam to compute the bending stresses in the bolts. The axial stresses in the bolt are determined by applying the vibratory radial loads to the bolt taking into account the expected joint efficiency. The resulting stress in the bolt is a combination of both membrane and bending stress. This is an alternating stress since the loads and deflections obtained from the vibrational analysis are cyclic.

To evaluate the bolts for fatigue, a concentration factor was applied to the alternating stress. Based on the experiences at the other plant, the most likely spot for failure was at the bolt threads. Therefore, concentration factors were determined for the bolt thread region. A different concentration factor was determined for axial stress and bending stress.

When evaluating high cycle fatigue, the alternating stress is usually below the cyclic yield of the material. When the material is cycling in the elastic range, 3 notch sensitively factor can be used to lower the concentration l factor. In the case of the 316SS bolts, a material notch sensitivity factor of q = .6 is applicable, Reference 6.4-1. The resulting fatigue notched i, reduction factor Kf, is determined by. j l

Kf = 1 + q (Kt - 1) l nu. c2o2n to 6-36 I

1

WESTINGHOUSE CLASS 3 where: Kf = fatigue notched reduction factor .

Kt = elastic stress concentration factor

~

However, if the alternating stress is above the material cyclic yield, a strain concentration factor needs to be determined and applied. Reference 6.4-2 provided these values.

The vibration evaluation provided loads and displacements that are RMS values, and these values need to be factored to obtain peak values. To accomplish this, a normal distribution is assumed for the distribution of 1 to 4 Sigma values. Using the normal distribution, partial usage factors for several Sigma values are obtained and then these values are summed to obtain the total fatigue usage on the bolt over some defined period of time.

The fatigue usage factors obtained are based on the failure curves for stainless steel and a combination of 2 curves are used. One is based on ASME j data and this curve is used for alternating stress (Sa) values up to 1E8 I cycles, and the other is based on work by M. J. Manjoine and is used for Sa values beyond 1E8 cycles. Both of these failure curves are presented in Reference 6.4-3.

Using the approach discussed, a history of bolt fatigue usage can be determined starting from the hot functional test to the current observed condition and extended beyond this point for an additional 18 month period. j l

6.4.2 Explanation of History 1

This section will present an explanation for the currently observed degradation of the thermal shield lower supports. This explanation is based on a best estimate high cycle fatigue evaluation of the lower support block top visible bolts.

1 I

l l

l mwnme ,e s.3y 1

WESTINGHOUSE CLASS 3 i

A vibrational evaluation was performed for different time periods of operation j using the support boundary conditions observed prior to that period of  !

operation. A fatigue evaluation, using only vibratory loads, of the top I visible bolts was performed for each period of operation at all or selected support blocks. '

Hot Functional Test i

During the hot functional test, the vibrational analysis predicts that the l

highest loads were placed on the blocks at the 0 degree and 180 degree locations. For the analysis, the test was considered to be 21 days in duration giving an estimated total number of cycles of approximately 11E6.

The initial conditions before the hot functional test were 4 limiter keys with tight tangential gaps, 6 support blocks and no flexures.

After the hot functional test was completed in late 1966, the lower support top visible bolts were all removed and examined. None of the bolts were found to be cracked or broken. This fact was used to baseline the fatigue

~

evaluation of the bolts. An unadjusted fatigue evaluation indicated that the 6., e bolts should have been locafeions.

~

To calibrate the fatigue evaluation to the observed condition, an adjustment factcr was applied to the loads. A factor was selected such that the usage factor in the bolt threads was just below the failure usage factor of 1.0.

This factor was used to adjust the load results from each time period evaluated.

Conclusions After Hot Functional Test

1) None of the bolts were broken or crheked. I

2) The usage factor in the top visible bolt threads at the 0* and 180' blocks is assumed to be just below 1.0. A load adjustment factor was selected to accomplish this in the fatigue evaluation.

3) The usage factor calculated in the top = visible bolt threads at the

' -.e,:,

_ _ applying the adjustment factor i determined in Conclusion 2 is approximately 0.0.

nu.ane io 6-38

i WESTINGHOUSE CLASS 3

l. Operation with 6 Flexures, 4 Limiter Keys and 6 Support Blocks l* After the hot function test, the plant operated for a total of 7.3 effective power years before the internals were inspected again. The initial conditions before this period of operation were 6 flexures at the top of the thermal shield, 4. limiter keys with close radial gaps.and 6 support blocks. A vibrational analysis and a fatigue evaluation of the top visible bolts were performed for these conditions over this period of time.

Conclusions After This Period of Operation i

1) A fatigue evaluation of the top visible bolts at the threads assuming the initial conditions of this operating period and applying the adjustment factor developed in the hot functional test case analysis produced a usage factor of approximately 0.0 at all block locations. Therefore, with 6 flexures, 4 limiter key and 6 support blocks, no additional fatigue usage of any significance is predicted for the bolts.

, 2) Inspection of the internals after this period showed that 4 of the 6 flexures were now broken.

Operation With 2 Flexures, 4 Limiter Keys and 6 Blocks This period of operation was for two calendar years but only one effective power year before the internals were inspected again. The initial conditions at the start of this operating period are 2 flexures intact, 4 limiter keys cnd 6 blocks.

Conclusions After This Period of Operation

1) A fatigue evaluation of the top visible bolts at the threads assuming the initial conditions of this operating period and applying the adjustment factor produced a usage factor of approximately 0.0 at all block locations. Therefore, a condition with 2 flexures, 4 limiter keys and 6 support blocks is not expected to produce any significant fatigue usage for the bolts.

2su.mone to 6-39

i

!! WESTINGHOUSE CLASS 3 l 2) Inspection of the internals after this operating period showed that

! only one flexure at the 124o location remained intact.

Operation With 1 Flexure Intact Thi.s period of operation was for 5 effective full power years before the i internals were inspected again with the core barrel installed. This period 1

ends at the current observed support degradation. During this operating period, various support block and limiter key configurations were assumed to predict the history of block degradation.

1 Flexure, 4 Limiter Keys and 6 Support Blocks The initial conditions for this assumed case are 1 intact flexure, 4 limiter keys and 6 support blocks. The period of operation is five effective power years.

Conclusions Based on These Assumptions

. 1) d fatigue evaluation of the top visible bolts at the threads assuming these conditions (and the adjustment factor) shows that fatigue usage would only be accumulated at the~ _ ((b oN location. For a 5 year period of time, the usage factor is expected to be much less than 1.0. Therefore, high cycle fatigue failure of the bolts is not expected for these assumed conditions.

2) Since bolt fatigue failure is not expected with this set of assumptions, a new set of assumptions are needed in order to match the current observed support block degradation.

O ms ic2o2se to 6-40

_---_--------J

WESTINGHOUSE CLASS 3 1 Flexure,fornLimiterKeys,and6SupportBlocks The initial conditions for this assumed case are 1 intact flexure, 4 worn limiter keys and 6 support blocks.

Conclusions Based on These Assumptions

1) A fatigue evaluation of the top visible bolts at the threads (applying the adjustment factor) shows that the usage factors at the[ Qfohations would exceed 1.0 before the end of the5yearoperatingperiod. However, the usage factor at the top bolts at the] h ock location exceeds 1.0 before the usage factor ofthetopboltsatthe[ 11oidion. Therefore, the bolts at the 0* block location are expected to fail first. Also, the bolt D, Cat fatigue usage at thel .1 block location would increase but is expected to remain below 1.0 for these assumed conditions.

, 2) The results of this assumed condition match part of the currently observed support block degradation, i.e. bolts have failed at the O'

. block location. This suggests that the assumptions of this case are consistent with the plant history.

1 Flexure, Worn Limiter Keys, O' Block Degraded and 5 Blocks Intact The initial conditions for this assumed case are 1 intact flexure, 4 worn limiter keys, 0 block degraded and 5 blocks intact. ,

Conclusions Sar.ed on These Assumptions

1) A fatigue evaluation of the top visible bolts at the threads shows thattheusagefactorsatthe[ {bk$$klocationswould l exceed 1.0 before the end of the 5 year operating period. The bolt  ;

usage factor at thel lbYo'ck would exceeg l.0 tzuch sooner in tine than the bolt fatigue u: age at thel 3 block locatien. Therefore, l itisjudgedthattheboltsatthel Eb$oc$werethenexttofail.

mwonus to 6-41

WEST!NGHOUSE CLASS 3

'2) The results of this assumed condition match the remaining observed support block degradation, i.e.. bolts at support block 240' had failed. -This suggests the sequence of historical events is consistent with the analysis assumptions.

c ., __ _

_ b, c . e.

3) The bolts at the remaining blocks ,, _have not accumulated any-significant amount of fatigue usage for any assumed' conditions so far in this 5 year operating period.

'l Flexure, 4 Worn Limiter Keys, O' and 240' Block Degraded and Remaining Blocks Intact The initial conditions for this assumed case are 1 intact flexure, 4 worn limiter keys, O' and 240' blocks degraded and 4 blocks are intact.

Conclusions Based on These Assumptions

1) A fatigue evaluation of the top visible bolts at the threads

. (applying the adjustment factor) shows that some factor usage is

~ - . b ,C ,%

expected to accumulate at the _ block location. However, this fatigue usage is expected to be below 1.0 for a 5 year operating condition with the assumed boundary conditions of this case and the usage is expected to remain below 1.0 for an additional 1.5 years of operation.

2) Thevisibleboltfatigueusageattheotherintactblocks[

s approximately 0.0 for this set of assumptions, b,c,e

3) With this set of assumptions the visible bolts at thel 3 block

^

{

would not accumulate any additional usage. However'c.if b e the previous caseisreferredtothevisibleboltsatthe[ 1blockwouldhave been predicted to have fatigued if the 240* block remained intact.

l' The assumptions of this case assume the 240* block to be completely

. degraded. It is possible that the 240' block remains functional enough that the visible bolts in the[ lblock could fail.

b,C ,e me.monoao 6-42

Q_,e ~

WESTINGHOUSE CLASS 3-1 Flexure, 4 Worn ~ Limiter Keys, 0*, 240' and 300* Blocks Degraded,

~

4 and 60, 120 and 180 Degree Blocks Intact

~

The initial conditions of this assumed case are 1 intact flexure, 4 worn limiter keys, 0, 240 and 300 degree blocks are degraded, and the.60,~120, and 180 degree blocks are intact.

p

' Conclusions' Based on These Assumptions

1) - A fatigue evaluation.of the top visible bolts at the. threads .

-(applying the adjustment factor) shows that a factor. usage of approximately 0.0 would be at-the remaining blocks, 60, 120.and 180:

degrees for this set of assumptions.

Conclusions of The Analysis The= analyses performed were on a best estimate basis and were baselined to the conditions of-support blocks observed after hot functional testing. Table 6.2 summarizes the previously discussed usage factor results.for:the top. visible

.- bolts. The observed condition of the lower support blocks shows that some of

.the bolts at the O' and'240' block location have definitely failed and based on this, probably all of the top bolts at these locations are cracked or broken. The conclusions of the analysis on the predicted extent of the damage is as follows:

b ,c. ,e.

1) Thetopboltsatthe[3blockfailedfirst,butnotuntil1 flexure remained and wear at the limiter keys occurred.

e,c. ,e 2)' The top bolts at the[ 3blockfailednext.

~

3) The top bolts at the 300* block are probably failed, but this is not completely clear by the analysis performed.

b, c. , e

4) The top visible b dts at the[ ] block location have probably accumulated some fatigue usage, but even considering the loading case based on the 0*

and 240' blocks degraded for a 6.5 year period, the bolts are not expected to fatigue due to vibratory loads.

mwenne io 6-43

WESTINGHOUSE CLASS 3

5) The top visible bolts at the 180' block location are assumed to accumulate a large amount of fatigue usage during the hot functional test, but the best estimate analysis does not predict any additional usage.up to the last assumed operating conditions.

The best estimate of the worst current condition is that the top bolts at the 0*,.240' and 300' blocks are probably cracked or broken. This compares closely with the physical damage observed at the lower thermal shield supports with the exception- that the 300* block does not visually exhibit evidence of degradation.

REFERENCES 6.4-1 " Engineering Evaluations for the Haddam Neck Thermal Shield Issue,"

WCAP-11729,AppendixC,4/88(Proprietary).

6.4-2 " Elastic-Plastic Analysis of Blunt Notched CT Specimens and Applications," W. K. Wilson (Journal of Pressure Vessel Technology).

6.4-3 " Techniques for Fatigue Testing and Extrapolation of Fatigue Life,"

M. J. Manjoine and E. I. Landerman.

l auwonae io 6-44 L_____-_-----------_-_----------------------

i 1

WESTINGHOUSE CLASS 3 6.5 Conclusions Regarding Expected Progression of Degradation This section will give the conclusions of how the observed and predicted current degradation of the thermal shield supports i's expected to progress over the next fuel cycle (18 months) considering only vibratory loads. The areas discussed are lower support blocks and the remaining flexure at the 124*

location.

6.5.1 Lower Support Blocks There are several scenarios that can be theorized on how the degradation of the lower support blocks will progross. However, each depends on the current conditions of the support blocks.

One Flexure Intact and The 0* and 240* Blocks Degraded The first scenario assumed is based on the currently observed condition of the thermal shield. The condition observed shows that only two blocks are degraded, the 0 and 240* blocks, and a flexure is intact at the 124*

location. This condition was discussed in the previous section.b,Thec ,e.

conclusion of this discussion was that the top bolts at thel ^} block were the only ones expected to accumulate any significant fatigue usage for this set of assumed boundary conditions. Also, the fatigue usage was not expected to exceed 1.0 for a 5 year period or for an additional 18 month period, assuming the initial boundary conditions.

Therefore, if enly the 0* and 240* blocks are degraded and one flexure remains intact, the degradation for this assumed condition is not expected to progress I

to the other block locations over the next 18 month period. This also assumes the 0* and 240* block do not beceme dislodged.

e me,mone ,o g.4s

WESTINGHOUSE CLASS 3 One Flexure Intact and 0', 240' and 300* Blocks Degraded

. The previous section of this document concluded that the worst expected .

current condition of the lower supports is that the O', 240' and 300* blocks are degraded. In this condition, with the one flexure intact, the top bolts at the 60',120' and 180' are not expected to accumulate any significant amount of additional fatigue usage for a'6.5 year period (5 years + 18 l l

additional months).

Therefore, if the lower supports are degraded in the worst expected condition, I

the degradation is not expected to progress to the remaining blocks (60', 120' and 180') over the next 18 months, provided the flexure remains intact and the degraded blocks do not dislodge.

One Flexure Intact and 0*, 180', 240* and 300* Blocks Degraded l

This condition assumes that the 180' block degrades leaving only the 60' and 120' blocks intact along with the flexure. For this condition, a fatigue  !

evaluation of the top visible bolts at the two remaining block locations shows

. that some fatigue usage is expected to accumulate. However, this usage is .

expected to be less than 1.0, even considering a 6.5 year operating period (based on'5 years + 18 months).

Therefore, the 60* and 120* support blocks are not expected to degrade over the next 18 months providing the assumed degraded blocks do not dislodge and the flexure remains intact, i Flexure Broken and 0', 240* and 300* Blocks Degraded This case assumes the flexure breaks over the next 18 mogtg. If this occurs, the predicted usage factor in the top bolts at thel . A bl ock location would exceed 1,0 before the end of the 18 months. The degradation is then expected toprogresstotheremainingtwoundegradedblocksatthe{ ] "' '

locaticns. The expected result is that all of the blocks would be degraded before the end of the 18 month cycle, mwman to 6-46

WESTINGHOUSE CLASS 3 Conclusions l

The current predicted worst condition of the 0*, 240' and 300* blocks degraded

~

is not expected to progress to the remaining blocks at the 60",120 and 180' locations over the next 18 months of operation. This is provided that:

1) the flexure does not break; and 2) a degraded block does not become l dislodged. The likelihood of either of these two events occurring is remote and i> aiscussed in more detail in the following sections.

6.5.2 Remaining Flexure History A history of examinations, starting at the conclusion of the original hot functional test of SONGS Unit 1, has provided an insight into the behavior of the reactor internals thermal shield flexura. The flexures were installed and were originally justified as an additional support for the top of the thermal shield. This additional support was necessitated, at both SONGS Unit 1 and a similar plant Haddam Neck, due to large motions observed during the first hot functional test. This is attributed to the unique bottom thermal shield support system of these early plants. Later plants incorporated a different system with the thermal shield supported at the top and flexure supports at the bottom.

After the flexures were installed, another abbreviated hot functional test was conducted to ensure the successful operation of the flexure. This operation covered a period of thirty hours, after which the flexure were examined and no problems were identified.

This flexure supported thermal shield functioned well during the initial fuel l

loads of the plant. Analytical evaluations for flow induced vibration of this l

condition indicate that the high cycle stresses are of a low magnitude when all six flexures are in operation.

O 3588s/020289 10 g_47

1 WESTINGHOUSE CLASS 3 l

\

After a period of 7.3 years of operation, an inspection of the reactor internals indicated that four of the flexures had failed. In reviewing the ,

two flexures that remained operational (at 124* and 244*) it is' noticed that j they are located in areas. of sof ter barrel flexibility, away from the hot leg I outlet nozzle and any support which exist inwardly from the reactor vessel.

Furthermore the.124' flexure had the largest shims between it and the core barrel which could indicate a different flexure support compared to the others. Analytical calculations indicate that as the flexures are removed from the analytical model, representing failure of the flexures, the flexure loads increase in the remaining flexures. '

Operations continued and in 1975, at the conclusion of another fuel cycle, another inspection was performed. This inspection indicated that the 244' flexure failed and that the only remaining flexure was at the 124' location. i Following this, plant operation was continued until the present shutdown time during which additional inspections were performed.  !

The present inspection has not indicated flexure failure. This is based on a visual examination of the flexure outside the core barrel and a visual

. inspection of the flexure attachment bolts from inside the core barrel. All of the flexure bolts seem intact and none of the threaded shanks are protruding inward towards the center of the core barrel, beyond what would be )

expected.

Another part of the present inspection has been to review the thermal shield attachment bolts at the lower end of the thermal shield. Recent video tapes of the inside of the core barrel, at the elevation of the thermal shield lower support blocks, have indicated that some of the bolt shanks have appeared to have failed, and that the threaded shank is progressing through the thresded hole towards the center of the core barrel. Analytical calculations have indicated that the flexure loads peaked.when the worn block analysis was completed at a condition identified as case 1,0,4. This is also identified as the case with one remaining flexure and four remaining functional blocks with worn keys. As further block degradation occurs, the flexure loads decrease on

i. a relative basis compared to the previous case identified as case 1,0,4.

mwona ie 6-48

WESTINGHOUSE CLASS 3 I Scenario The most likely scenario, based on recent analytical studies and inspection

~

observations is that: as a consequence of flexure failure the previous redundancy for radial thermal loads of six flexure supports, has been removed and the resulting radial thermal loadings should therefore decrease.

However, as a consequence of flexure failure the fatigue loadings on the remaining flexure have analytically increased. This increase is a relative increase indicating a trend and not an absolute value. In actuality, as of-the most recent observations, the 124' flexure has not failed. Analytical studies augmented by the visual inspection of the thermal shield support block bolts lead to the conclusion that the condition identified as case 1,0,4 must have existed for an appreciable amount of time. This leads to the conclusion that if high cycle fatigue failure were to occur, it would have occurred during this time frame. This conclusion is based on the observation that the broken support bolts appear to have been broken and " walked" inward over an appreciable amount of time. Appreciable, in this sense means that sufficient cycles have occurred to be considered a high cycle condition. Analytical e

studies of the hot functional condition indicate that high cycle fatigue, if it occurred, would occur in a relatively short amount of time.

This leads to the conclusion that operation at the Case 1,0,4 condition, with the high cycle fatigue loading, of the flexure have not exceeded the endurance limit. In other words, with this condition existing for an appreciable amount of time, it is apparent that the stresses in this component must be below the fatigue curve and that continued operation for the next fuel cycle can be continued, since the slope of the fatigue curve is very shallow in the high cycle fatigue range.

Furthermore, to add confidence to the above evaluation, continued operation for the next fuel load will be at lower flow velocities and lower flow loads due to the added tube plugging that was performed during this outage. These

^

lcwer flow rates lead to lower excitation levels further reducing the stresses I

below the fatigue curve. This adds further margin for the flexure and further justify continued operation of the next fuel cycle.

nu.mone io 6-49

WESTINGHOUSE CLASS 3 Seismic l

For the seismic event, the flexure has been evaluated for SSE seismic loads.

Theseismicevaluationindicatesthattheradialflexureloadis7 ]pYu'nds '

~

and'thetangentialloadis{ ]p'[u'nds. No significant verticai load is seen by the flexure due to the seismic event.

b,c,e 3Ksi,apeakshear These stress off seismic loadsstress h d i and a bending reduce o [ ] Ks to a direct stress of1 i w

- g,c,e clamped guided beam. Thesest,resgesresolveintostressintensitiesof17Ksi attheneutralaxisandl._,Ksi at the extreme fiber. At either location the stress is less than the 3.6 Sm limit of the faulted event, and the fle:ure is adequate for the seismic event.

O e

9 m e,ioso2n so 6-50

WESTINGHOUSE CLASS 3 SECTION 7.0 CONSEQUENCES OF PJRTHER DEGRADATION Expected Support Block Degradation Based on the results of the visual inspection, failures exist in a few of the upper visible and hidden bolts on two of the support blocks. Based on the analysis of the failure mechanism and the observations at'the Haddam Neck plant, it would not be surprising if all the upper bolts in the degraded support blocks are failed. However, failures in the lower support block bolts are not expected based on the observations at the other plant. In addition, unlike Haddam Neck, none of the dowel pins were observed to be protruding,  ;

even in the 240' block which had the most obvious damage to the bolts.

Further damage to the already degraded blocks is expected to be limited to migration of the ends of the broken bolts until they fall into the bottom of the vessel and the eventual failure of any top bolts which have not yet broken. Since the time required for these effects to occur car.not be readily predicted, the present evaluation conservatively assumes that this corresponds to the current configuration of the blocks. The heads of these bolts are expected to remain in place since they are held in place by lockbars and no significant loads act on these lockbars once the bolts have broken.

As previously discussed, substantial further degradation of the thermal shield support system is not expected to occur for the next 18 month fuel cycle since the flexure is expected to remain intact and since loads on the support block fasteners are equal to or less than they have previously experienced. However, since some degradation is present, conservatism dictates that an evaluation of the consequences of further degradation be performed.

'- 7.1 Worst Credible Degraded Case l

In the limit, the worst credible case (or " fully degraded case") would be failure of the remaining flexure, failure of all the support block bolts and

" loosening" of the dowel pins due to increased thermal shield motion. Because of the significant amount of weld on the dowel pins, the close fit at the end of the dowel pins and the fact that all of the inspected pins were still in an.-una m 7_i

WESTINGHOUSE CLASS 3 place; assuming that all the dowel pins somehow become loose is a sufficiently conservative assumption for the evaluation of the worst possible degraded case.

Even if all the upper bolts in all six of the support blocks were to fail, it is expected that the dowel pins plus the lower bolts would remain intact based on the Haddam Neck experience. This means that the support blocks would still remain in place, fastened to the core barrel. Axial movement of the shield would be restrained by the dowel pins and the ledge on the support blocks.

Kadial motion of the shield would be restrained by friction forces between the block and the shield as well as by the geometrical constraints inherent in the support design. Even in this extreme situation, the support blocks would ensure support for the thermal shield.  ;

Even in the more extreme case that all the fasteners (upper and lower bolts and dowel pins) in all the blocks fail, the blocks provide support against downward motion of the thermal shield. Because of the reduction in support stiffness, the vibratory levels would increase but the stability analysis, Section 7.3, confirms that the shield would not become unstable. A vibration analysis of this " fully degraded" situation has been performed in order to l estimate vibratory amplitudes and frequencies and to facilitate the estimation of impact loads between the core barrel, thermal shield and the support blocks. The resulting loads have been used to assess the structural integrity of the core barrel and the support blocks and those analyses confirm that even in this fully degraded condition, the support blocks would not shear off and the stresses in the coro barrel would be acceptable. This extreme situation could also lead to impact loads at the limiter key locations and these loads were also estimated. The resulting stresses induced within the core barrel would also be within acceptable limits.

The effect of operating in this condition has been explicitly evaluated in order to determine the resulting flow induced vibration and seismic loads plus the effect of these increased loads (including impact loads) on the structural integrity of the thermal shield support system and the core barrel. In addition a stability analysis has been performed to demonstrate that even in

^

this postulated worst credible case condition, the shield would still remain stable and that the supports would not fail and allow the thermal shield to fall downward, me.-mae to 7-2

[ WESTINGHOUSE CLASS 3 1

The justification for these conclusions regarding the worst credible degraded conditions follows:

,3 As'shown in Figure 2.0-1, the thermal shield rests on six (6) support blocks and these support blocks fit into a groove which is machined in the core barrel. Near the top of the blocks, there are two long bolts l

which clamp the thermal shield, block and core barrel together. These bolts restrain the relative radial displacement between the shield, core barrel and support blocks. Dowel pins are inserted through the shield and partially through the support blocks and these pins restrain the axial movement of the shield relative to the core barrel. There is also a Lolt (hidden-bolt) which clamps the top of the block to the core barrel and restrains the relative motion between the core barrel and the support block. There are also two dowel pins and two bolts near the bottom of the support block and these fasteners connect only the barrel and the support block. As with the upper fasteners, the dowels restrain the relative axial motion between the barrel and the shield and the bolts restrain the relative radial motion. Any moments generated by vertical loads on the shield are reacted by the contact between the blocks and the core barrel as well as by the upper visible and hidden bolts. It is also noted that even if the dowel pins somehow lost function, the axial loads exerted by the shield are reacted into the core barrel through the groove which supports the blocks. Thus, the design of the support block assembly is conservative from the standpoint that even if all the bolts and dowels were to fail, the support would still support the thermal shield.

Another means of ensuring axial support for the thermal shield is a "self-locking" feature that is shown in Figure 7.0-1. Even in the extremely improbable event that the ledge on which the shield rests somehow fractures and falls off the block, the downward motion of the shield is restrained by a stop on the support block.

Worst Conceivable Degraded Case .

Ultimately, the worst possible scenario that could be postulated is a complete failure of the support system such the shield either tilts because of a loss of some of the supports or the shield drops because of a loss of all the lower M80s-02028910 73

WESTINGHOUSE CLASS 3 O

I i

Figure 7.0-1. Thermal Shield Support Block isa.-cro2n in 74  !

WESTINGHOUSE CLASS 3 supports. The analysis of the fully degraded case (above) has shown that this is not a credible scenario. Nevertheless, an evaluation of the structural o effects of this worst conceivable case has been performed and an estimate has been made of the resulting effects on flow and pressure drop as well as core r inlet flow distribution. The primary motivation for performing this evaluation was to ensure that even in this incredible case, the dropped shield would still be restrained by the lower radial supports and that core flow would not be significantly reduced. The evaluation of this hypothetical case is presented in section 9.0.

7.2 Thermal Shield Hydroelastic Stability Analysis The thermal shield / support block arrangement is a type of structure which has been called an " inverted pendulum," and is a configuration which can be unstable under certain flow conditions. Figure 7.2-1 illustrates the

" inverted pendulum" idealization of the thermal-shield / core barrel geometry.

Mark (Reference 1) performed a stability analysis of the inverted pendulum configuration and found it to be statically unstable beyond a given flow velocity, but had no dynamic or " flutter" instabilities. The velocity at which static instability occurred was proportional to the square root of the support stiffness.

The Mark (Reference 7.2-1) analysis is not direccly applicable to the SCE thermal shield structure for several reasons: a) equal channel widths were assumed (a=b); b) the stability limits presented omitted the effect of fluid friction, although fluid friction was included in the broader analysis; and c) the flows into each channel were assumed equal and invariant. This last condition is equivalent to assuming that the flows upstream of the thermal shield are separated by a barrier.

The present analysis was performed for conditions which are more appropriate for calculating the stability of the SCE thermal shield than was the Mark

- (Reference 7.2-1) analysis. In particular, the effects of unequal channel widths and fluid friction are included. In addition, the " flow barrier" boundary condition at the upstream end of the thermal shield was eliminated in favor of one which is more appropriate to the SCE thermal shield. This i

3569s-020289 10 7 ...]

WESTINGHOUSE CLASS 3 l

)

condition requires that the pressure and velocity of the flow upstream of the thermal shield be invariant. Other than this and the usual requirement that

. the steady-state flow in each channel satisfies a fluid momentum equation, no l

restrictions are placed on the flow (steady-state plus transient) I e distribution. -These boundary conditions and the Lienard-Chipart stability criteria (Reference 7.2-2) lead to the possibility of a dynamic or " flutter" type of instability as well as a static instability. However, it is shown that the support stiffnesses of the SCE thermal shield are sufficient to l

preclude the occurrence of either type of instability. j The analytical model shown in Figure 7.2-1 corresponds to only half the thermal shield. In using this type of model, it is inherently assumed that the motion of the thermal shield is rigid body rotation about an axis at the support block level. Only a half-model is required because, with rotation of this sort, the pressure loads on the two halves of the thermal shield are equal and opposite. The equation of motion of the thermal shield is therefore made to include the pressure loads on both halves by doubling the loads calculated with the half-model of Figure 7.2-1.

The half model representation simulates the thermal shield support with a rotary spring K0 at the bottom and a linear spring Ks at the top. The forces acting on the thermal shield in addition to these spring forces are its weight Mg and the fluid pressure forces acting in Channels 1 and 2 on either side of it. Thase pressures will vary axially in each channel and will vary as the channel widths A and B vary with the thermal shield motion. The .

t problem is therefore very much a dynamic one, and stability will depend on whether the fluid-structure interaction tends to amplify or attenuate the thermal shield motion.

The equations governing this system are: a) the mass and momentum conservation equations for the fluid flow in each channel; and b) the equation of motion of the thermal shield. In the present analysis, only axial fluid

. momentum equations are considered. In the actual situation, there will be some circumferential flow passing between the two sides shown in Figure la.

l- The present analysis neglects this flow because it is conservative to do so and because it would add complexity to the analysis which is well beyond the 1

me.-mas ' 7-6

WESTINGHOUSE l CLASS 3

(

scope intended for this effort.. The present work'is comparable to that of

-Mark (Reference 7.2-1)-in its complexity and is discussed in detail in

.. Reference ,7.2-3.

D Stability' Limits-The' fluid and structure governing equations can be combined to yield a linear -

differential equation in the gap b alone. By substituting b = bo.exp (At),

the characteristic equation for. stability is obtained:

R13+R1 3 2 +RA+R4=0 3

(7.2-1)

(characteristic equation) where R 1

, R2 , R3 and R4 are functions of the fluid and mechanical-properties. of the system-and are defined in Reference 7.2-3.

~

For stability, the Lienard-Chipart (Reference 7.2-2) criteria require that, if R1 > 0.(which an inspection reveals is always true):

R2>0 (7.2-2).

R4>0 (7.2-3)

R2 R3>R1 R4 (7.2-4)

It can be shown that condition (7.2-2) is always satisfied. Condition (7.2-3) is the static stability condition and can be written:

-- b0E 3 3 m*

'N 4

. n...-e2ene io 7-7

WESTINGHOUSE CLASS 3 Equation (7.2-5) is a lower limit on the effective stiffness required as a function of the flow dynamic pressure, frictional, and irreversible hydraulic o losses. The terms on the right-hand side of equation (7.2-5) have the dimension (dynamic pressure / gap).

e

• Condition (7.2-4) represents a true dynamic, or flutter, instability because it specifies an effective stiffness limit which depends not only on fluid dynamic pressure and geometry, but also on the thermal shield mass and the damping'of the support system. Because of the complexity of the result, it serves no purpose here to expand inequality (7.2-4) beyond its present form.

It does, however, represent a lower limit on stiffness. The limiting stiffness is, therefore, the larger of the two limits determined from (7.2-5) and (7.2-4). Stiffness K,ff is defined as:

__ -.. b>C &

(7.2-3) where L is thermal shield length.

Applying inequalities (7.2-2), (7.2-3) and (7.2-4), the SCE thermal shield has the following hydroelastic stability limits:

b,c,e.

(7.2-7)

(7.2-8)

For the above calculations, fluid velocities characteristic of mechanical design flow were used.

3569s-02u28910 7.g

WESTINGHOUSE CLASS 3 Other Stability Considerations The previous sections have addressed the hydroelastic stability of the SCE thermal shield system. From these sections and Table 7.2.3, it can be -

concluded that the thermal shield will be stable, unless the support blocks are severely degraded. However, even if stable, the thermal shield can experience high amplitude vibrations if the value of K eff is sufficiently close to the statically limiting value. In such a case, the system stiffness is the difference between the rotational stiffness provided by the support blocks and the static stability stiffness limit. The characteristic equation for stability (Equation 7.2-1) assumes that the total pressures upstream of both channels are the same. If, instead, a sinusoidally fluctuating total pressure difference between the two streams is allowed, equation (7.2-1) is no longer homogeneous and contains terms on the right-hand side which are proportional to the tmplitude of this pressure difference. The resulting equation can be solved by standard techniques in which the channel widths a and b are set equal to t:0 exp (iwt) and gb exp (iwt).

The results are shown in Figure 7.2-2 for effective stiffnesses K eff which are 2, 5 and 10 times the static stability limit.

Amplitudes are expressed in the figure as inches per psi of total pressure difference between the two inlet streams. In the turbulent region around the inlet nozzles, it is estimated that pressure fluctuations can be as much as 0.5 psi - 1.0 psi.

However, the difference between the pressure fluctuations at the two channel inlets can be expected to be much smaller than these values. This is particularly true in light of the fact that Figure 7.2-2 assun.es this pressure fluctuation difference is maintained over the entire inlet, which is highly {

1 unlikely. It is, therefore, considered that more appropriate range for j pressure fluctuation difference is 0.05 psi - 0.1 psi. In this case, the j thermal shield vibration amplitude due to hydroelastic stiffness reduction is I less than-- . ~IinN, even with an effective block stiffness of twice the statically limiting value. While this is a significant amplitude, it is

- relatively small compared to those induced by turbulence.

nao.-cro2so io 79

WESTINGHOUSE CLASS 3 i i

Thermal Shield Rotational Stiffness For the purpose of calculating the actual effective rotational stiffness of the thermal shield, the limiter keys are assumed not to be engaged and support flexures are absent. This leaves no support at the top, and Ks in equation 7.2-6 is 0.

The rotational stiffness Ke provided by the blocks comes from several sources: a) compression of the support blocks; b) bending of the block; c) dowel pins; and d) bending of the bolts. In the case of a degraded block, the

, flexibility of the thermal shield and core barrel are assumed. Cont,idering these parameters, vertical stiffnesses are computed in the up and down direction at each block taking into account whether the block is degraded or undegraded. After sclecting the expected axis of rotation for the thermal shield an equivalent rotary stiffness is computed.

Table 7.2-3 presents rotary stiffness values fer various lower support conditions. The first condition, Case 1, represents the predicted current condition of the thermal shield supports. This condition assumes the 0, 240 and 300 degree blocks are degraded. The axes of rotation considered, 1-1 and 2-2, are based on the thermal shield beam mode results from the vibrational analysis using similar support conditions. The lowest expected rotary stiffness for this lower support configuration is in-lb/ rad. This value is times greater than the static stability limit, b,c , e.

- -" L a The next support condition considered, Case 2, in Table 7.2-3 is the completely degraded condition. TherotarystiffnesscomEutedforthissetof

_ __tsc w_ e,c, supports is Ke =L ._ in-lb/ rad. Thisvalueislltimeshigherthanthe static stability limit. Case 3, considers the completely degraded condition plus assumes the O' or 240' block dislodged or broken. This condition is not 4 credible and is extremely remote but was evaluated just to determine .

sensitivity. The results indicated that even for this extreme situation l although the static stability limit is violated and the shield could tilt, it would still not become dynamically unstable.

m e.- m a sio 7-10

~ WESTINGHOUSE CLASS 3 IThe next set of lower support conditions considered assumes the blocks at both the 0* aad 240* locations become dislodged. For the all block degraded case, exceptthe120' block,therotarystiffnessiscomputedtobfl[ ~ D ' '

[-

in-lb/ rad. This stiffness value'is just above the static. stability limit of j1 [I$-ib/ rad. Therefore, for this extreme case, the shiold is not expected to tilt or become dynamically unstable.

The expected condition of the thermal shield (upports at the end of g'he next 18monthperiodiseitherCase1}-_ '((orCase2}[ ((ilthe last remaining flexure breaks. For both of these cases, the computed rotary stiffness is high enough not to cause a stability concern.

e i

l l

l e

0 3589s-020289 10 7,, g }

J-WESTINGHOUSE CLASS'3 l-

. TABLE 7.2.3.

. ROTARY STIFFNESS FOR VARIOUS LOWER SUPPORT CONDITIONS-1 ..( Top End Assumed Free and-No Flexures) l f

l * .

Y a9 e

.1

-4

'4

__ b, C, e.

i nee.-onne io 7-12 i

WESTINGHOUSE CLASS 3 TABLE 7.2.3 (Continued) o d

i a

i e

9 i

e e

me..exas to 7-13

WESTINGHOUSE CLASS 3 REFERENCES.

' 7.2-1. Mark. W. P., "Aero/ Hydrodynamic Stability of Elastically Supported Plates in Narrow Channels with Upstream Flow Barriers Preventing

~*

Flow Redistribution," Journal of Vibratio'n, Acoustics, Stress, and Reliability in Design, Vol.107, July 1985, pp. 319-328.

7.2-2. Zadek., L. A. and Deseor, C. A., " Linear System Theory: The State Space Approach," McGraw-Hill, New York, 1963, pp. 419-420.

7.2-3. Schwirian, R. E., Bhandari, D. R. and Chandra, S., Stability of the NUSCO Nuclear Reactor Thermal Shf.:d With Assumed Degraded Support Conditions, Fluid Transients-in r'. aid-Structure Interaction

( ASME-FED-Vol . ' 56), pp. 79-86, ASME WAM-1987, Boston, Mass.

O i

i 3589s-020269 10 p t _---_-- _ _

e-c, b

s Y ;5

WESTINGHOUSE CLASS 3 7.3 Vibration Analysis for Worst Credible Degraded Case-The worst possible (credible) degraded case was defined'as follows:

o The. remaining flexure at 124' fails so all flexures are non-functional.

o All 'four limiter ' keys are worn to the extent that they are non-functional.

o All top bolts, hidden bolts and bottom bolts in each of the six (6) support block assemblies are broken.

o All dowel pins are worn and therefore loose rather than tight.

o The support blocks stay in place.

The results of the vibration analysis for this worst possible degraded case ,

are presented in Table 6.2-2 and are designated as Case 0 0 0. Note that since all the radial bolts are assumed to be broken, the radial loads are zero. The displacements resulting from this case were used in the impact loads evaluation and wear evaluation for the core barrel, support blocks and limiter keys.

1. 5 s

a we,-owse so 7-17

WESTINGHOUSE CLASS 3 7.4 Seismic Evaluations Non-linear time history analysis of the San Onofre reactor vessel / core barrel /

thermal shield system was performed for the most degraded conditions. The objective of this evaluation was to determine the dynamic response of the system consisting of displacements and impact loods at the limiter keys, core barrel flange and the lower radial keys for the most degraded conditions.

A simplified mathematical model of the San Onofre reactor vessel and the lower internals consisting of core barrel, thermal shield and the lower support plate is shown in Figure 7.4-1. For the applicability of this simplified model, the fundamental vibrational frequencies of the thermal shield and the core barrel ,

were matched with the ones obtained from a detailed three-dimensional system finite element model with degraded conditions.

A comparison of fundamental beam mode frequencies for the thermal shield and the core barrel obtained from the simplified model of Figure 7.4-1 and those obtained from the 3-D detail systeni model are shown in Table 7.4-1.

. The mathematical model shown in Figure 7.4-1 was developed using WECAN (Westinghouse Computer Analysis), a general purpose finite element computer code. The elements used in the model consisted of:

Three-Dimensional Straight Pipe Three-Dimensional Spring Rotary Spring Three-Dimensional Non-Linear Impact Spring General Matrix Input for Hydrodynamic Mass For time history response of the reactor vessel and its internals, synthesized I time history accelerations are required. The synthesized time history accelerations used in this analysis were based on a conservative design response spectra shown in Figure 7.4-2. Also shown in Figure 7.4-2 is the spectra generated from the synthesized time history accelerations used in the analysis. For the sake of comparison, the San Onofre design response spectra at the reactor vessel support elevation and the spectra used in the analysis mwo**

7-18 I

i

. , ~ , :

'WESTINGH00SECLASS3 lM L

TABLE 7.4-1 FREQUENCY COMPARIS0N Component Beam Mode Frequency l

Simplified Model 3-D Detail Model

__ b3C e' 2

Thermal Shield' Core Barrel e

e '

4 S

4 3580s-020289 10 7 yg i

WESTINGHOUSE CLASS 3 b,:, e.

t 5

i 1

I l- Figure 7.4-1. Simplified System Model me.-omse '

7-20 E_------_---__----- -- - - - . - . - - - - - - - - - - - - -

i 11i -

rmwH.zoxoCwm or>ww w e-c, b .

a r

t c .

e p

S s

i s

y .

l _._

a _

n -

A i

c

. m s

i e

S

. e -

r f

o n

O n

a S

2 4

7 e

r u

g i

F 10

. 9 8

2 0

7 0 .

s 9

8 5

3 uiO

WESTINGHOUSE CLASS 3 is shown in Figure 7.4-3. The San Onofre Design Basis Earthquake Spectra at the reactor vessel supports is shown in figure 7.4-4 and 7.4-5; and this O

Design Basis Earthquake is also referred to as 0.67 g modified Hausner Earthquake. It can be seen from Figures 7.4-2 and 7.4-3 that the response

, spectra generated from the synthesized time histories envelops the design response spectra of San Onofre Nuclear Station. It-is to be noted that the structural damping ratio for application in the Design Basis Earthquake (DBE) is four (4) percent of critical damping. From figure 7.4-3 it is seen that a j very conservative spectra is used in the analysis to determine impact loads.

The thermal shield flexure and support block loads were determined for the conditions: a) in which all six support blocks were assumed to be degraded, and b) in which only three support blocks were degraded. For the evaluation of flexure and support block loads, response spectrum analysis was-used.

Horizontal seismic excitations in both X and Z were applied simultaneously to determine the thermal shield loads. Frequencies in the vertical modes of vibrations for the thermal shield are significantly high (> 30 Hz) even in

, the most degraded conditions. Therefore, simplifying approach used for rigid structures (1g) is adapted for vertical excitations.

l The system loads generated from seismic analyse were used in detail component i evaluations and a summary of seismic loads is given in the following table.

l 1

s 1

1 1

! l 3569s-020289 10 7-22

I ,Iji rmmH EaxoCwm or>mw w e_

3 C,

b s _

i _

s y

l a

n _

. 3 A 4

i c

7 m s -

. e i r e u S g

i e F r f

o n

O n

a S

0 1

. 9 8

2 0

2 0-s 9

8 5

3 uI~w l;l'iL

WESTINGHOUSE CLASS 3'

'l i

- b, c , e.

l l

o a i

l

- Figure 7.4-4. Spectra Primary Horiz-North South (El. 10.93')

m e.-o202ee ,o 7-24  :

=

WESTINGHOUSE CLASS 3 bA,o 4

e t

i

- Figure 7.4-5. Spectra Primary Horiz-East East (El. 10.93')

35fl9s-02028910 7,

WESTINGHOUSE CLASS 3

SUMMARY

OF' MAXIMUM SEISMIC LOADS

. Component' Value- Description 1 -

b,c.e _

Limiter Keys ~ 4 ImpactLoadperkeyj

~

Thermal Shield Flexures Radial Direction-Tangential Direction Thermal Shield Support Block Loads

. Case a) All Six Blocks Degraded,-No Flexure Block Location Fr(1bs)' Fvert. (1bf)'

'(Degrees) b,c.e 0

300 >

240 180 120 60-Case b). Three Blocks Assumed Degraded-One Flexure Intact No Flexure Block-Location' Fr (1bs)- Fvert. (1bs) Fr (1bs) Fvert (1bs) b,c.e 0 i 300 240 180-120

'60

, , b,c.c

. ._ _ b,c e For vertical seismic excitations,.___ loads should be considered.

we.-oso,so ,o 7-26 i

WESTINGHOUSE CLASS 3 7.5 Impact Loads Evaluation The vibration analysis of the thermal shield-core barrel system treats the limiter keys and support blocks as linear springs, with spring constants

. chosen to represent a combination of free motion followed by impacting. This is adequate for estimating the overall flow-induced response of the system, but a more refined model is needed to obtain a reasonable estimate of impact loads. The model used here is one that has been used successfully in other reactor applications involving flow-induced impacting, and which is supported by test data. In this model, the vibration is considered to have a dominant frequency, and the vibration amplitude is assumed to start at zero and build up gradually, as in a lightly-damped system. When the amplitude reaches a value at which impacting will occur, the impact force is conservatively calculated as if the impact were to occur approximately one cycle after a near miss. The vibration is assumed to stop momentarily after the impact, and to then be followed by another buildup period, impact, etc. Basically, the model reduces the dynamics to that of a spring-mass-dashpot excited at its natural ,

. frequency fn by a force Fo sin (2n fn t). For a lightly-damped system with zero amplitude at time zero, the time history obeys the following

. equation:

b,c,6 (7.5-1)

~ ~ ~

__ - - b,c #-

Since the system is lightly-damped, the damping fraction is much smaller than i

unity. Consequently, the time constant of the exponential term in (7.5-1) is much longer than the period of oscillation of the sinusoidal term. Under these conditions the " envelope" of the oscillations is given by:

___ _ _- b , C 6 3

l.

(7.5-2) 3589s-020289 10 7y b_---------------------------------------- J

WESTINGHOUSE CLASS 3 where g, =

Impacting occurs when the amplitude Xe reaches a value equal to the limiter  !

key gap g. At this point it is assumed that a "near miss" takes place, and j that impacting occurs during the next vibration cycle when displacement again equals the gap g (see Figure 7.5-1). This assumption is made to maximize the calculated impact force. The impact rate is then the inverse of the total period T in Figure 7.5-1. The impact force is calculated by equating the kinetic energy of the impact to the peak stored energy in system during the impact.

Impact forces were calculated for the limiter keys and for the support blocks in the degraded condition. This was conservatively done by using a combined mode zero-to peak amplitude (g-), and the modal stiffness (K) of the dominant mode. Since the impact force is a function of the gap (g), a range of gaps from a minimum value to the maximum value of g- was considered. For the limiter keys, the limiting mode was found to be the n=2 mode, for which the following characteristics were determined from the vibration analysis:

Limiter Keys b, c, e_.

gc,e 6. C, c-Withtheseinputdata,apeakimpactforceok Zlbs at a gap of1 []nch was calculated. For the minimum gap of~ _

~]1nch,theimpactforcewas

. calculated to be Ibs. hCA

~

~ f.,,c c.

3589s-020289 1 7-28

l f' . WESTINGHOUSE CLASS 3

(:-

.. b,64 Figure 7.5-1. Vibration Impact Model 3580s-020289 10 7,

i

_ _ . _ . . _ _ _ . I

WEST 1NGHOUSE CLASS 3' The limiting mode for the support blocks was found to be a thermal shield beam or " rocking" mode. In this mode, the thermal shie rd rotates about an axis through two diametrically opposed support blocks, and the shield impacts vertically on two other support blocks, on one side of the " rocking" axis.

The characteristics of this motion are listed below:

Two Support Blocks

__ b,CA e,c,c _ b,C ,e Apeakblockimpactforceoff '31bs,wasfoundatagapof[ _ inch.

Each of the two blocks receives half this value. At the minimum gap of]

~ }

• inch,theimpactforcewascalculatedtobe[ 1bs.

~ b,C,C-

, It should be recognized that the impact forces calculated in this way are very conservative for several reasons. First, the calculations are based on the assumption of a severely degraded condition, in which vibration amplitudes are very large. Secondly, the impact is assumed to be absorbed by a single limiter key, with no energy dissipation. Thirdly, the impact transient is calculated in such a way as to maximize the impact forces. Finally, the full, combined mode, zero-to peak amplitudes are used, even though a limited number of modes dominate the vibration which leads to impact. For these reasons, the above loads are considered to be substantial over estimates of the impact loads that might, in fact, occur.

Structural Evaluation of Support Blocks and Bearing Surface In the very unlikely event that all of the bolts and dowel pins attaching the shield to the support blocks and core barrel were to fail, the blocks are still captured by the shield and core barrel. This section will demonstrate that if all the fasteners did fail, the blocks can still adequately support the shield.

me.-onae io 7-30

WESTINGHOUSE CLASS 3.

- In the' event all the fasteners did fail,- the downward force of the shield will-push down the outboard arm of the block, and cause an outward rotation of the

, upper part. The eccentricity is balanced by a radial: resistance by the shield

-against the block. It has been determined.that the following downward loads-exist on the outboard arm.

_ b.t k Fv1=Downwardforceduetoflowinducedvibration(FIV)=_ ~~ lb NOTE: This is a 4* Sigma value taken from a normal' distribution curve

' ~~

Fv2 = Down force / block due;to vertical seismic Fv3 = Down force / block due to horiz seismic' 4 _

The most. critically stressed section'is the horizontal bearing surface between the block and the core barrel. When tolerances are adversely added and the block is assumed to be radially touching.the shield (creates minimum bearing

. . . . area), it is found that[ ] sq.in, of bearing exist between each block and the barrel. The corresponding stresses are:

31 W

S2 S3, . i-S1 -

This is less than the yield strength and is acceptable for normal or faulted conditions. Note that the FIV and seismic loads are cyclic and therefore may be added by SRSS.

b,c,c The shear area at each block is:

1 .T--in2. The _ b,e shear stress is therefore: . .. ..

_jp si ., e.This is much

- less than 0.6*Sm and is acceptable for normal or faulted conditions.

e o

nee.-enne io 7-31 i-l _ _ . _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _

lO

[ WESTINGHOUSE CLASS 3 1'

H From the width-to-height ratio between the loads, the horizontal reaction is 0.483*Fv. The horizontal reaction between the shield and the block is applied a

2.38 in. above bottom surface. The radial force due to the loads presented above pxcluding seismic) is:

. b,cA '

Fhl = lb.

An additional radial reaction against the blocks must be considered due to the horizontal component of the seismic event. 'This load is:

__b,c.,e, F

h2 = radial reaction to to horizontal seismic [ _.

Therefore:

__ L,c,L The bending stress during the seismic event is:

_  % A c'-

S= ,

which is much less than 1.5*Sm and is acceptable for normal of faulted conditions.

An additional evaluation of the block is needed for a postulated case where '

not all of the block fasteners are degraded. Here,theboltedblocks(3 total) will see higher loads than those described above. For this case, the critical load is the shcaring stress at the base of the horizontal support.

For this case, the load at each block are:

- h, C, C,

• Note: FIV load determined from analyses described elsewhere in this report.

me.-emse io 7-32

WESTINGHOUSE CLASS 3 b,c.e 2

.The s' hear ~ area. at each block is: A=. in , The shear stress'is therefore~ 1 . ~ psi.

, This' is less than the 0.6*Sm and is acceptable for normal or faulted b 3c,e.

conditions.

w If the load is seen by the bearing surface (e.g. worn dowel pins), the bearing stress is:

E,c ,e. <

[- 1 psi,whichislessthan2*yieldandisacceptablefor faulted conditions. For norm:1 operation,~ the stress.is:

6,C g *~

psi, which is approximately equal to the actual yield strength at temperature and is therefore considered acceptable.

It should be noted that, based on the loads, no liftoff of the shield from the blocks is anticipated, and on that basis, no dynamic load factor would be used. As a matter of interest, it has been determined that a 10 inch drop on '

the radial keys would result in a factor of around 1.25. Since the shield can

. only lift off the blocks about 1 - 2 inch before becoming disengaged, and because the keys are considered stiffer, a dynamic load factor of close to 1.0

, would be realized for any postulated impacting between the shield and blocks.

A factor of 1.0 means the above stresses are unaffected by liftoff.

7.6 Effect of Seismic, Vibratory and Thermal Loads on the Core Barrel This section of the report addresses the adequacy of the core barrel at the attachment of the thermal shield and demonstrates that the core barrel can withstand the associated loads during the next 18 month fuel cycle. The two locations on the core barrel are 1.) the upper thermal shield load transmittal point which is the limiter key and 2.) the lower thermal shield shield support block attachment to the core barrel. The loads considered at the top support are the high vibratory impact load and the seismic loading coming from the l thermal shield. The loads considered at the lower support are the high cycle impact load, the seismic loads and the dead weight loads coming from the thermal shield. No specific thermal loads are identified at these locations l, as they relate to the thermal shield.

me,-none io 7-33

l I

WESTINGHOUSE CLASS 3 7.6.1 Impacts Between the Limiter Key and the Core Barrel To better understand the load paths and functioning of the upper limiter key, the next paragraphs provide a short history describing how the function of the limiter key as evolved.

The original design of the upper support of the thermal shield was a key - l clevis arrangement between the thermal shield and the core barrel. This was the original design concept before the introduction of the flexure attachment at the top of the thermal shield. The original key clevis design incorporated a key welded into the top of the thermal shield and a clevis machined into the outside of the upper core plate pin attached to the core barrel. The original concept was to transmit this load in the tangential direction, by designing a J

.0005 to .003 inch tangential gap between the key and the clevis. The original radial gap between the key and the clevis was .145/.150 inch.

With this original arrangement, the load transmitted from the the mal shield

. to the core barrel was a tangential load. The key, attached to tte thermal shield was in shear, and the corresponding load on the clevis, machined into i

, the back of the upper core plate pin, was a circumferential load applied to the barrel.

With the design modification to add the flexure at the top of the thermal shield, the load path through the key and clevis was modified. The tangential gap was opened and the radial gap was tighten. This changed the load path from that of a tangential direction to a radial direction. Note, this was intended as a backup load path, to complement the primary load path offered by the addition of the flexures. This new radial load path changes the stress distribution in the key and clevis components. The following explanation details the important stress areas in this new radial load path direction.

The three highest stressed areas for this radial load path are a.) the bearing stress between the key and the clevis, b.) the shear stress in the weld of the key to thermal shield, and c.) the shear stress in the lip of the upper core i, plate pin, which restrains the inward radial load. The most critical of these l

three items is the shear stress in the lip of the upper core plate pin.

l m...... m I 7 34

Other parts of the report have identified a vibratory radial load between the

-.4 c.c.c top of the thermal shield and the core barrel as+g a pounds at a frequency of[ ) Hz. ThisJscons_jdgedahighcycleloadcondition,sinceatthis frequency, almost g c a ycles would occur during the next 18 ty nth jue cycle. The shear area of the lip of the upper core plate pin isu 4n, i

-t 5 c>c 6ce and the radial load generates a shear stress of t-u 4 Ksi. Note: this value exceeds the 0.6 S, limit of the ASME code, but is less than .6 of yS if the actual yield is considered to be 21 Ksi.

,b,c,C For fatigue, a fatigue failure value of g jKsi,isconsideredat 0

10 cycles. Considering the total range stress as two times the maximum shear stress, valid since the load is in one direction only, the alternating t- -i stress is half that value and is p 4 Ksi. Considering a concentration factor of 1.5, a typical value for peak shear stress at the midpoint of a section, the resulting stress is, 4 Ksi. This provides a margin of 4 b.c .e 4 _+t 4 o the allowable value of 34 Ksi mentioned above. Note: margin here is I definedas(allowable / actual)-1.

.p_ q b,c.c For seismic considerations, an additional load of g a pounds must be applied to the limiter key. This increases the total stress to 44.3 Ksi which is greater than the 34 Ksi limit mentioned above. In the worst case, if the seismic event were to occur just at the end of the 18 month fuel cycle and if it is assumed that worst case impact loads for the worst credible degraded '

case acted throughout the entire cycle; the seismic load at the top of the

+ thermal

-i e,c.eshield, applied inward in the radial direction, coupled with the g 4 pound vibratory impact load, would fail the upper core plate pin lip in shear. The pin could then move ir, ward but would be restrained in its position, being entrapped by the upper core plate. Thermal shield motion would also be limited by the one-half inch gap between the thermal shield backing strips and the core barrel.

7.6.2 Loadings at the lower thermal shield support This section of the report addresses the loading on the barrel at the interface with the thermal shield support blocks at the bottom of the thermal shield. Other parts of this report have identified the downward force due to me,-c20'" "

7-35

WESTINGHOUSE CLASS 3 b , e. e.

flow induced vibration of[ [lbs. This section of the report will add the dead wo f t and the seismic load to determine the worst case loading onto the core barreb b, C , f-,

Theseismiceventaddsaverticalreactiono( 11bs.onthecorebarrel ledge from the vertical seismic excitation. The horizontal seismic excitation

- _ _ u , e.

also produces a vertical load on the suppgrtp and this value is 1bs forthefullydegradedcaseand[ ]lbsforapartiallydegradedcase which is the worst condition. The flow induced vibratio, load is multiplied by four to obtain a 4-sigma value taken from a normal distribution curve. The dead weight per thermal shield support is{ 11Is. This is added directly

~ eg.e . e . c. e-andthesumtotalsi- _

jbs which produces a bearing stress o ( _Ksi on the barrel, for the fully degraded case.

For the partially degraded case, only three of the blocks are considered active in supporting the load andc the

--- ,c. c.

flexure is assumed to be broken; the t i, e.

F.I.V. load increases to _ . libsresultinginatotalloadof__ Tibs

. and a stress ofL lbsY These values are less than two times yield, and ensures successful operation even for a seismic event.

9 7.7 Support Block - Thermal Shield-Core Barrel Interface Wear Analysis At the horizontal interface between the thermal shield, support blocks, and core barrel wear can occur when the two surfaces are in contact. In the degraded condition, these surfaces are likely to be in contact as the system vibrates, particularly in the " rocking" mode described in the impact load evaluation. The wear can come from two sources: a) Impact wear of the thermal shields on the support blocks; and b) Radial (sliding) motion of the thermal shield on the support blocks when the two are in contact, and c)

Sliding wear between the support blocks and the core barrel. These are considered separately below:

i me.-cune ,o l 7 3s

WESTINGHOUSE CLASS 3 Impact Wear at the Support Block-Thermal Shield Interface The basis for this calculation is given in the impact loads evaluation l section. The volume wear rate is calculated as the product of impact energy,

) ,

impact rate, and wear coefficient. The first two of these are obtained from i the results of the impact load evaluation. The wear coefficient for stainless steel-to-stginlesssteelimpactingisconservativelyestimatedtobe[_

_in2/lb. Using this approach, and converting volume wear rate t

~ ~ h,c,e linear wear rate by appropriate insertion of the two-block area of square inches, a first order differential equation for the rate of change of ,

gap with time was derived. This equation was numerically integrated to yield the gap wear-time history shown in Figure 7.7.1. The initial gap was taken to b,C C O .C ,0.

bel ] inch. With this model, a gap of1t?1inchwascalculatedtooccur after approximately 1.5 years, i.e., a wear depth of~' ~ inch.

Sliding Wear at the Support Block-Thermal Shield Interface

., In addition to impact wear, the possibility exists that the radial vibration of the thermal shield could cause some sliding wear during that part of the

, vibration cycle in which it is in contact with the support blocks. At some point in the cycle, however, the frictional force between the shield and the block will be large enough that the radial vibration will stop. When the frictional force again falls to values which allow radial motion, sliding resumes. A calculation of the wear induced by this mechanism was made using a

~

slidingwearcoefficientof[ 2 in /lb., a friction coefficient of 0.4, and a " free" radial shield amplitud/If[.diIc$ The result was{

inch of wear in a 1.5 year period. It should be noted that the wear

[

coefficient for sliding wear has a great deal of uncertainty associated with 2

it. The value --

in /lb.

--e. c 4 is

--considerde 4to be a best estimate lin/lb.to--- _ -2 2

value, with a range of variation of_ in /lb.

Most test data, however, fall in the low end of the range.

O m,-cac2a to 7-37

WESTINGHOUSE CLASS 3

'40,0

.d

.- Figure 7.7-1. Support Block Axial Gap Impact Wear Versus Time l

3589s-C20280.10 7,

4 WESTINGHOUSE CLASS 3 m

Sliding wear at the Support Block-Core Barrel Interface

. , If the block is_ assumed to-be free to slide laterally, a very conservative postulation will be made here that somehow it will move back and forth through

.. the available' clearance at each load cycle. If this happened, the horizontal friction reaction between the block and the core barrel could wear that surface.- A wear evaluation will be performed to show that the support ledge will not wear down to the chamfer of the core barrel, thereby ensuring support for the shield.

K1 = Approximate wear coefficient for stainless on stainless

=1 taken from tests elsewhere)

}(6,c.c.

h, C, C X1=Maximumavailableradialclearance=1_ 3in.

The table below gives the bearing load, number of cycles, and resultant volume removal for each load from sigma = 1 through sigma = 4, where these are

. deviation values from a normal distribution curve.

. SIGMA FIV LOAD # CYCLES TOTAL LOAD VOLUME BLOCK INC. WEIGHT REMOVED u 4,c.e_

• For this case, the horizontal reaction is only about1 ltimes the b, c, e deadweight. Therefore, sliding is not be expected.

b,C,C. l

, The maximum expected wear depth is therefore, .

I 3580s-020289 10 7 39

1 WESTINGHOUSE CLASS 3

{

There is at least 1.5 inches of material between the bottom of the support block and the elevation where the core barrel begins to narrow. Therefore, there is no concern that the support surface will wear away.

, It should be noted that there is no restoring mechanism to cause the unbolted  ;

block to slide back and forth at each cycle; therefore, the above sliding calculation is very conservative. In' reality, the loose block will likely tilt through the very small available clearance and wedge in place, and minimal or no sliding would be expected. l l

Wear Summary A calculation of the wear at the support blocks - thermal shield interface, inducedby"3nch rocking"ifweariofthethermalshieldandradialvibration, n a 1.5 year period. This includes both impact yields and sliding wear. The calculation is considered to be conservative because the degraded block condition of the thermal shield is assumed. The impact

., wearcalculatioriL ]inih$hasadditionalconservatismswhicharediscussed in the impact loads evaluation section. An additional wear evaluation of the

, interface between the block and the core barrel indicates that the support surface will not wear away.

1.

l \

l afb me,-onne so 7-40

(. .

WESTINGHOUSE CLASS 3 SECTION 8.0 LOOSE PARTS ANALYSIS As noted in Section 5.0, there are at least three broken support block bolts

. which are in the process of rotatirg out of the core barrui. Loose pieces of bolt originating from the thermal shield support will tend to fall and be carried by the downward inlet flow into the lower plenum. It is expected that these bolt fragments would tend to collect in the bottom of the reactor vessel I since that is the lowest point in the system. It is also possible that these fragments could fall out ar<d remain trapped in the region between the lower support plate and the lower core plate. There would be no adverse effect if this were to happen. Even if the pieces were lifted by the flow and passed through the lower core plate, they could not pass through the bottom nozzle of the fuel assemblies unless they broke up into smaller pieces. While this is unlikely, the ultimate consequence could be debris induced cladding perforations which would be manifested by increases in coolant activity. For the expected case where the fragments collect in the bottom of the reactor

. vessel, one issue which is normally assessed is impacting between the loose parts and the bottom mounted instrumentation. Since there is no bottom mounted

. instrumentation at the San Onofre Unit 1, there is no concern regarding impact from bolt fragments.

The other issue that needs to be evaluated is the case where the fragments lodge under the secondary core support in a cold condition. During subsequent heatup, the gap under the support will close due to differential thermal expansion between the reactor vessel and the internals. Any fragments which are lodged under the support would load the core support columns in compression. Note that this is only a concern if during the next fuel cycle the plant is cooled and reheated; if this does not occur, this issue doesn't exist.

Nevertheless, an evaluation was performed assuming that a cooldown and heatup does occur. During the cooldown, a maximum cold gap of approximately 0.981 inches will be realized between the reactor vessel cladding and the base of

, the secondary core support. Thus, the body of a support bolt (0.875 inch diameter) could place itself under the support. In this event, the resultant me,-c2 cue ,o g.1

WESTINGHOUSE CLASS 3 heatup would compress and yield the bolt. The following section is a quantitative description of what happens if.this occurs.

l The critical part of the secondary core support are 4 energy absorber bsA sections. Theseareneckeddowncolumnswithatensileareaof[~[sq. in.

each. The crushing of.the bolt fragments on subsequent heatup would impart the same type of yielding that the column is already designed for. The forces needed to crush the fragment and the subsequent yielding of. the column is summarized as follows:

Syl = . Yield strength of energy absorber = 18.5 ksi at 550*F -

b, o , C-Fy1 = Force to yield.each energy absorber =[_

Jb From Roark, " Formulas. for Stress and Strain", 5th edition, page 517, the applicable equation for deforming a cylinder between 2 surfaces may be taken as:

Max Sc = 0.798*SQRT[(p)/(Kd*Ce)]

Where:

p = load per inch of length Kd = D = 0.857 Ce = 2*(1 .32 )/26E+06 = 7.00E+08 The 0.875 inch diameter bolts have an original length of 5". However, when broken, it is not realistic to assume that the full length of broken bolt will come out; therefore. a lergth of four inches will be assumed. Note that this length does not include the sucket head, which is too large in diameter to fit into the 0.981 inch gap.

m ,-ero2s. io 8-2

WESTINGHOUSE CLASS 3 Therefore:

6,c e-p=l ]1b/in

__ b,c4 psi = linear compressive stress in fractured bolt at the Sc =__

point where energy absorber begins to yield This is certainly not a realistic stress, since it is many times the yield ,

strength of 60000 psi. It simply indicates that the bolt will crush long

)

before the energy absorber.

As plastic deformation of the bolt continues after yield, the effective area will increase. To estimate if the bolt will completely yield or. dig into the

- stainless steel cladding and baseplate, assume a final effective diameter equal to the minor diameter of the bolt (0.7387 in.). The resultant stress between the fully embedded bolt and the surrounding surfaces is b,c,c-

, Sc1 = psi > 22000 psi = estimate of actual Sy.

, The contact stress is significantly greater than the yield strength of the contact surfaces; therefore, the bolt can be expected to fully dig into the surrounding surfaces before any yielding of the energy absorbers is realized.

As a further matter of interest, the even more unlikely prospect of numerous bolts lodging under the core support will be examined. Clearly, if a sufficient number of bolts were to be involved, eventually the energy absorber will begin to yield. However, the energy absorber is designed to undergo plastic deformation. The following is a description of that deformation.

X1 = Maximum expected plastic deformation caused by an indefinite number of bolts lodged under th6 core support

= 0.7387 e

e nu.-eano se 8-3

' WESTINGHOUSE-CLASS'3

'L1 = Nominal length of necked down section of.~ energy absorber

= 3.'375 in.

1

, 'EL.= percent elongation of the energy'. absorber

=.0.7387/3.375

=:0.219

- This is significantly less than the ultimate elongation of the energy absorber-

material; the energy absorber may still be considered functional.

e.l s

l~

t O

( ..

me.-onne io 8-4

WESTINGHOUSE CLASS 3  !

SECTION 9.0 EVALUATION OF WORST CONCEIVABLE DEGRADED CONDITION j 1

As discussed in section 7.1, a complete failure of the thermal shield supports which would allow the shield to either fall or become tilted on it's supports is not credible based on the design of the supports. Nevertheless for conservatism, an evaluation of the structural effects of the dropped thermal shield was performed in order to demonstrate that even in that limiting case, I the downward motion of the shield would be restrained by the lower radial supports for the reactor internals. In addition, an evaluation of the effects f

of either a tilted thermal shield and a dropped thermal shield on the flow distribution, pressure drop and thermal design flow was performed in order to ,

demonstrate that even in those incredible situations, no catastrophic effects would result. Although not evaluated because the lower radial support can withstand the impact loads, it should noted that even if the thermal shield were to somehow fail the lower radial supports, there would still be some, albeit reduced, level of flow into the core.

9.1 Hydraulic Evaluation

\

Two "what if" scenarios have been postulated that is, l

1) The thermal shield becomes tilted or

2) The thermal shield drops and comes to rest on the radial keys.

1 Evaluations have been performed to assess the impact of these two postulated I events on the hydraulics within the reactor pressure vessel system. These evaluations determined: a) the change in hydraulic resistance and pressure drop within the reactor vessel / internals system, b) whether the thermal design flow could be maintained and c) the effect on the core inlet flow maldistribution.

o 3589s-020289 10 g.}

WEST!NGHOUSE CLASS 3 Tilted Thermal Shield This evaluation assumes the thermal shield retains support at the support blocks but has maximum tilt. At locations where the gaps between the thermal shield and the core barrel and vessel are small, the flow will tend to be lower than at other azimuthal locations. A conservative evaluation was i performed by assuming one-dimensional (axial) flow in a parallel array of circumferential1y distributed flow channels. Eight parallel channels ,

distributed azimuthally were considered. This calculation is conservative because it does not permit fluid to flow azimuthally from regions of high  ;

axial velocity and pressure loss to regions of low axial velocity and pressure loss, and thus maximizes the overall pressure drop and maldistribution.

Figure 9.1-1 shows the model that was used to evaluate the hydraulic loss factor in a given one of these channels. A given total flow enters at the top and splits into a sub-channel one (core barrel-thermal shield) flow and a sub-channel two (thermal shield-vessel) flow. Entrance, frictional, and exit

, hydraulic losses are calculated for each sub-channel and are used to estimate the flow in each sub-channel and the overall flow resistance of the channel as

, a whole (i.e., the two sub-channels in parallel).

It should be noted that this analysis is complicated by the fact that the flow split between sub-channels in a given azimuthal channel is not known a priori. Consequently, the entrance and exit hydraulic loss factors for each sub-channel cannot be calculated a priori because the entrance and exit area ratios are not known. Thus, the flow split and the entrance / exit loss for each sub-channel must be determined iteratively. The procedure for doing this is outlined below:

a. Calculate hydraulic loss factors and frictional losses for both sub-channels, with assumed flow split.

b. Calculate inlet-to-outlet pressure drop for each sub channel. If they do not agree within a certain error band, adjust the flow split l1 accordingly. For example, if the calculated pressure drop in Channel 1 is higher than in Channel 2, reduce the flow in Channel 1, relative to Channel 2.

me -azo 2mo 92

WESTINGHOUSE CLASS 3 SEPARATION STREAMLINE

.- g ji' '

. W , ENTRANCE A LOSS

_ . - - - . _ 1 -_____

CHANNEL FRICTION J

/ TWO LOSS CHANNEL [

DNE i

i j i s LOSS

- . . - - . . . o -------

/

( SEPARATION STREAMLINE

- Figure 9.1-1.

ssa.-owne io 9-3

WESTINGHOUSE CLASS 3 l

c. Recalculate the hydraulic loss factors and frictional losses and repeat the procedure.

d. Repeat until convergence is obtained.

4 Using this procedure, the azimuthal flow distribution and the effective K-loss of the entire barrel-vessel annulus is calculated.

TheresultsofthesecalculctionsindicagthatthetotalK-lossofthe vessel-barrel annulus increases byl ]ue to the tilting of the thermal shield.

Asaresult,theincreaseinthetotalvesselpyssurejrgpfromthe _

as-designed to the tilted condition was determined to be ,.- The effect of 4- _ _ e ,c. e_

the tilted shield on the total vessel flow will be a reduction of+t .

GPM/ Loop. Thisflowreductioncanbeaccommodatej,sdncethereissuffkcient marginbetweenthethermaldesignflowof[' ~~GPM/ Loop and the best estimate vessel flow. Thus, the thermal design flow can be maintained even with the thermal shield in the maximum tilted condition.

From the above described hydraulic model, the maximum and minimum flow fractions were determined to be

. , bs C s t-respectively. However, these circumferential flow effects will result in an azimuthally uniform flow after an axial distance of only 25-35 inches. This estimate assumes that the only I mixing mechanism is that which exists is free expansion, i.e., turbulent  ;

stresses resulting from the velocity mismatches between the adjacent streams in the model. In addition, the powerful azimuthal diffusing effect of the vessel bottom curvature and the secondary core support structure may also level out any azimuthal flow non uniformities which may exist. It can therefore be concluded that the azimuthal flow distribution at the core inlet will be approximately the same with either a normal or a tilted thermal shield configuration. '

Dropped Thermal Shield This evaluation assumes the thermal shield has fallen onto to the radial

, keys. The primary effect of a dropped thermal shield on the hydrodynamics within the reactor vessel is that it places a high velocity region closer to 3589s-C2028910 g_4

r- 1 WESTINGHOUSE CLASS 3 the core than was previously the case.. With the thermal shield supported at~

l. b the support blocks, the velocity at the entrance into the lower plenum is{ 1,c,e.

ft/sec. For the dropped thermal shield case the annulus has the additional blockage due to the thermal shield and the velocity becomes[ ]ft/sec.

c , c, e As the flow enters the lower plenum and flows over the bottom head, two effects can result in significant radial inlet pressure gradients at the core inlet. One of these is the pressure recovery that occurs as the flow converges at the center of the vessel. The other is the centr.ifugal pressure gradient resulting from this flow as it passes over the curved bottom head.

Both of these effects increase the pressure difference between the center of

-3 the core and the core periphery, and both become higher as the flow velocity into the lower plenum increases.

The increase in hydraulic resistance in the lower plenum due to the dropped  :

thermal shield was estimated and the increase in the total vessel pressure drop from the as-designed to the dropped condition was determined to be[ ].

. The effect s,cs of the dropped shield on the total vessel flow will be a reduction of[ ]GPM/ Loop. Thisflowreductioncanbeaccommodatedsin,ceJhereis

, sufficientmarginbetweenthethermaldesignflowo( _

GPM/ Loop and the best estimate vessel flow. Thus, the thermal design flow can be maintained even with the thermal shield falling onto the radial keys.

- 6 C, e These calculations determined that the effect on the core inlet flow ma1 distribution due to a dropped 6,c, thermsi shield was that the inlet flow to the .

peripheral assemblies will be--

Lc d-

_ ower while the center assemblies will be] [_ 4 higher than the average. For the case when the thermal shield has not dropped the peripheral assemblies ar{ hhwand the center assemblies are[_1I11gh, on average. O' 3589s-020289 10 9-5

WESTINGHOUSE CLASS 3 Conclusions Although not considered probable, the effects on the hydraulics in the reactor pressure vessel system of two postulated events have been determined. The

, effect of a tilted or dropped thermal shield on overall reactor pressure vessel system pressure drop and flow rate was determined and is not expected to be significant. Therefore,thethermaldesignflowrateof[ 1GPM/ Loop can be maintained. bC' E For the dropped thermal shield case, it was determined that the core inlet flow to the erigheral assemblies will bel llo$e'r while the center assemblies willbe[1h 'er than the average. Core performance calculations have been performed and the results indicate that this effect can be accommodated during normal operating conditions without affecting core limits.

9.2 Thermal Shield Drop Structural Evaluation

, For a postulated thermal shield drop onto the radial keys, the structural adequacy of all impacted components was evaluated. This condition, through

, unlikely, is postulated during the event in which all six support blocks are assumed to fall off from the thermal shield and core barrel. Figure 9.2-1 shows a schematic sketch for the hypothetical drop onto the radial keys.

Typically for impact evaluation, the effective mass (Me) is that fraction of total structure mass determined based on the assumed static deformed shape of the structure due to the applied impact force. The equivalent mass is derived I

to maintain equality of kinetic energy with the real system.

M, = / m v(X) 2 dx where m = mass per unit length; and, f(X) = assumed deflected shape.

c The impact time is given by:

t d

= 2L/C 3589s-023289 10 g_g L - _ _ _ _

WESTINGHOUSE CLASS 3 9]" es l7 - .

[ /

N / IA

, THERMAL SHIELD N l / ,

l  % l /

/ s l s /

p p i N CORE BARREL BADIAL KEY ETATED INTO SECTION)

. i

/pj e

/ .

l ,

' ~~

RADIAL KEY (6 TOTAL) b A8 _

SECTION A.A

• Figure 9.2-1. Schematic Diagram for Hypothetical Drop on Lower Radial Restraints n n.-e x anio g,7

WESTINGHOUSE' CLASS 3' The force during-impact is given by:

P =' aCV, where ,

P = contact pressure (psi) a = mass density of missile (1b-sec2 /in4)

C = speed of sound .in missile, / E/a (in/sec)

V, = impact velocity (in/sec)

L = Length of the missile (in)

The ability to mobilize the effective mass of the static deformed' shape, (1st mode of structure), can be evaluated based on the response of the structure to a rectangular plus force time history. The response in each mode of vibration for a uniform beam is'given:

b 10 D(X)'n(x)dx X" = w2 m 10b (DLF)=B(DLF) 2 (x)dx where:

DLF = Dynamic Load Factor

'n(x) = n rmalized shape of mode n p(x) = load as function of x un = circular freqJenCy m = mass per unit length b = length of beam i 4

Figure 9.2-2 shows the Dynamic Load Factor (DLF) for a rectangular pulse load as a function of tddivided by the period of vibration (TI) of the structure. From this figure, it is seen that for t /T1 ratios greater than d

0.5, where Ty is the period of the first mode, the predominant response will be in the first mode, or the assumption for target mass based on the static deformed shape is valid.

l

, 35ess-02028910 gg L_ _ _ _ -

p ^

l.

WESTINGHOUSE CLASS 3

- .Using appropriate param,eters for San Onofre reactor internals and the thermal shield drop'of-- .

~~incliks through water, the impact load on the lower radial keys was determined using the principles of Energy conservation. Th s,c,e

-,- -- e impact force applied to the radial support was calculated to be4l .) __lbf.

' The structural evaluations of the radial keys, radial key to support saddle weldment, support saddle to lower support plate weldment, key weldments, core

l. barrel shell, and the core barrel flange were carried out. The stress results of these evaluations show that the structural adequacy of all the impacted components is maintained for this postulated condition.

4 e

I l

............ g., ,

WESTINGHOUSE CLASS 3 D

2.0 1.5 -

f E

o 1.0 -

U

.(

l' O.5 -

0 !I!!  !  ! ! ! !III I ! ! ! !I!!

5 10'1 2 5 1$ 2 5 101 Cy = sd/T Figure 9.2-2 Dynamic Load Factor me.-c2ano io 9-10 1 _ . _ _ _ _ _ _ _ _ _ _ _

l WESTINGHOUSE CLASS 3 SECTION 10.0 MONITORING PROGRAM l* 10.1 Overview j As shown above, the unit can operate safely with the postulated worst case degradation of the thermal shield supports. Detection of degradation during the next cycle of operation, should it occur, provides SCE with information to make decisions regarding the continued operation of the unit. l Internals vibration monitoring, using the neutron noise method and loose parts monitoring, will be performed during the next fuel cycle to detect degradation of the thermal shield supports. The noise signals from ex-core, power range flux detectors will be monito ed to detect changes in internals vibration.

Accelerometers will be mounted on the exterior surfaces of the reactor vessel  !

to detect loose parts that could be generated due to further degradation or further disengagement of presently degraded fasteners. The implementation plans for these programs are described below.

10.2 Internals Vibration Monitoring Internals vibrations will be monitored using the signals from the eight ex-core, power range detectors. The program will consist of three phases as described in ANSI /ASME Standard Part 5, OM-1987(1): baseline, surveillance I and diagnostic. The specific program to be implemented on SONGS 1 is 4 described below.

Baseline Phase Baseline data will be acquired at partial and full power during the initial power escalation of the next cycle. In addition to the ex-core detector signals, data from in-core moveable detectors and accelerometers mounted on the reactor vessel will be acquired at 100% power. The in core detector data and reactor vessel accelerometer data are acquired because these data might (1) ANSI /ASME Part 5, ON-1987 Operation and Maintenance of Nuclear Power Plants., " Inservice Monitoring of Core Support Barrel Axial Preload in Pressurized Water Reactors" me.-oro2" "

10-1

WESTINGHOUSE CLASS 3 provide additional information to support interpretation of the ex-core detector signais. Also as part of the baseline, data from the eight ex-core detectors will be reviewed after 1 month of operation, 2 months of operation l and approximately 3 month intervals thereafter.

.. 1 The initial data will be reduced and reviewed to determine the likely types of modes reflected in the signal content. These results will be interpreted using previous experience with neutron noise data, and correlated with  !

analytical model results. Based on the information from this work, the signal content and analytical model results for several postulated thermal shield support conditions, screening guidelines will be established for use in the surveillance phase of the program.

In the surveillance phase, the data will be reviewed for changes in frequency and amplitude content. However, presently available information can be used to consider the likelihood of detecting significant thermal shield degradation.

The likely present condition of the thermal shield supports has been inferred from inspection results and analytical studies. Detectable differences between the response of the thermal shield in the expected present condition and the responses with the postulated loss of support from all upper bolts in the blocks and loss of support from the remaining flexure are indicated based on analysis results, and available neutron noise information.

Analysis results indicate that failure of the remaining unbroken flexure would cause the beam mode natural frequencies of the thermal shield to change significantly. The natural frequency of this mode is calculated to be greater than Northeexpectedpresentcondition. For the worst case degraded condition, modal analysis results indicate that the thermal shield beam mode natural frequencies are abouti ~_l[z. Since the calculated vibration levels for this case are an order of magnitude greater than the levels of thermal shield vibration detected in other plants, this response should be i detectable. In addition, increases in the levels of the thermal shield n=2 shell modes are also expected to accompany a postulated degradation to the l worst case condition, further supporting detection.

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-WESTINGHOUSE CLASS 3 Note that a delay in achieving the response characteristics associated with the worst case condition could occur because~of vibration amplitude limitation

,=

at the keys. The fully degraded response would appear more clearly with further wear.

Analysis results indicate that less degraded conditions than the worst case might also be detectable. For failure of the unbroken flexure with no further degradation of the bolts from the expected present condition, the thermal shield beam mode natural frequencies are calculated to change from greater than 10 Hz to approximately 4 Hz. For this condition, however, the calculated thermal shield beam mode natural frequencies are close to the thermal shield n=2 shell mode natural frequencies.

Surveillance Phase Surveillance during operation will be conducted by periodically reviewing data from selected pairs of ex-core detectors. The data will be reviewed:

On an every-other-day basis for the first eight days at 100% power, on a once per week basis thereafter.

After an earthquake (part of SONGS system review guidelines after earthquakeof.25g)

After detection of increased loose parts activity If the responses exceed the screening guidelines established in the baseline phase-of the program, the diagnostic phase of the program will be entered.

Diagnostic Phase

~

The diagnostic phase of the program is to establish whether or not the observed changes are due to changes.in the vibration of the thermal shield which might be indicative of degradation of the thermal shield supports and to

establish actions. The actions could include, for example, additional data I acquisition, more detailed data reduction, or consideration to shut down to limit the potential for further damage.

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WESTINGHOUSE CLASS 3 10.3 Loose Parts Monitoring

~The thermal shield LPM program is intended to monitor the potential deposition of thermal shield fasteners or. fragments in the Reactor Vessel bottom. It is anticipated that sequential fastener failure (if it occurs) would result in an increasing number of loose parts collected at the RV bottom.

Hardware Implementation Currently four accelerometers are stud mounted and torqued on the RV seal ring flange. Their sensors are fed through four in-containment remote charge converters (preamplifiers) to a signal processing rack out of containment, populated with vibration monitoring type equipment. Due to access limitations, no accelerometers can be mounted on the RV bottom nor can impact testing be performed there.

Program l

o Verify performance of currently installed accelerometers using simulated impacts of various mass and energy levels on the RV proper 180 degrees from each installed sensor. Retain this data for future comparison. I l

Evaluate frequency content and amplitude. Be prepared to directly mount i alternate sensors to seal ring if performance is not adequate. l 1

o Upgrade processing equipment to allow audio monitoring and true-RMS )

amplitude trending. ]

o Implement startup data collection program integrated into nuclear noise program.

~

o Develop formal LPM program that incPades on going qualitative evaluation by shift personnel (STA/SRO), quantitative trending and alert procedures.

4 o Monitor during power escalation and at power Audio monitoring by an assigned operator on a once per shift basis.

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WESTINGHOUSE CLASS 3 Measure arid plot true rms levels on a weekly basis.

5 -

Tape record signals on a weekly basis for the first month, on a monthly basis thereafter. -

o If changes are detected by audio surveillance or in the graph of rms energies, perform a loose parts analysis on the current data and historical data.

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}

WESTINGHOUSE CLASS 3 l SECTION

11.0 CONCLUSION

S A. The 124* flexure is still intact based on the visual inspection performed

, of the flexure and its' attachments to the thermal shield. In addition, the observed condition of the lower support block assemblies is consistent with pre-inspection predictions based on the assumption that the flexure is functional. For the reasons outlined in Section 6.5, tnis flexure is expected to remain intact throughout the next 18 month cycle of operation.

B. The " hidden" bolts in the 0'and the 240* support block are clearly broken at this time. In addition, one of the top long visible bolts in the 240*

block is also clearly broken. Based on the experience at another plant and based on engineering analyses, it is concluded that the hidden bolts and the top long bolts in both of these support blocks are either broken or cracked or they can be expected to fail in the future.

C. Depending on the unknown amount of displacement limiter key wear that has

, occurred, it is also possible that the three top bolts in the 300* block are damaged or soon will be damaged. Although no visible evidence of damage to this block was observed, a conservative assessment of the current condition would assume this block is also degraded.

D. Based on the observations, analysis and the experience at the other plants which use this type of support, all the lower bolts and all the dowel pins are still intact and fully functional at this time.

E. The cause of the observed degradation is high cycle fatigue failures of the upper bolts due to flow induced vibration, probably resulting long term wear at the displacement limiter keys coupled with the degradation of the thermal shield flexures. Based on the analyses which were performed, it has been concluded that loads due to vortex shedding from

' the trailing edge of the thermal shield and loeds due to thermal effects

_ are not significant factors in the degradation of the support block assemblies.

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} WESTINGHOUSE CLASS 3 l

F.- Analyses which were performed prior to the inspection showed that even if significant limiter key wear has not occurred, degradation of the 240*

and 300* blocks should be expected. Further analyses showed that if the l limiter keys have high wear, the the 0* and 240* blocks should be

', degraded but the 300* block might not be. As a result it is concluded j that the vibration and structural analysis provides a reasonable basis i for evaluation of the behavior of the thermal shield support system.

G.I If the 0*, 240* and even the 300 are currently degraded such that the top bolts in all three blocks are broken and if the 124* flexure remains intact, no degradation of the remaining support blocks is expected for the next 18 month fuel cycle provided that the currontly degraded blocks do not become dislodged. As will be explained later, no viable mechanism exists for the blocks to become dislodged and the remaining flexure is expected to remain intact.

H. The remaining flexure at 124* is expected to remain intact throughout the

-, next cycle of operation since it haE operated sufficiently long that if high cycle fatigue were the cause of the other flexure failures, this one

, also should :.e failed by now. In addition, radial thermal loads on the flexure have now been reduced since the other flexures are broken.

Finally, analytical studies have indicated that flexure loads due to flow induced vibration should reduce if unexpected degradation should occur at additional support block locations.

I. Analysis indicates that in the unlikely event that the remaining flexure should fail, the beam mode natural frequencies of the thermal shield would change significantly. In addition, the calculated levels of vibration for this condition are substantially higher than the levels of vibration which hate been detected by using neutron noise monitoring in other plants. Thus while flexure failure is not expected, should failure occur; it is expected to be detectable using neutron noise monitoring.

J. Although analysis has been performed which indicates that significant

, further degradation of the thermal shield support system is not expected during the next fuel cycle, the consequences of an unexpected failure of m e .o m eeto 11-2

WESTINGHOUSE CLASS 3 1

all the bolts in all the support blocks plus failure of the remaining flexure have been evaluated. The results of this worst credible degraded case indicate that while vibratory amplitudes would increase, the thermal shield would not experience dynamic instability and the support system

, would still maintain it's function of supporting the shield even if it is postulated that the design basis earthquake occurs while the plant is running. In other words, even for this worst case, the shield would remain in place and loadings on the core barrel would not be expected to result in significant damage to the core barrel.

K. As the ultimate "what if", the effect of a complete failure of the thermal shield support system which results in a loss of support was also investigated. In this unrealistic , postulated situation, the thermal shield would fall onto the lower radial supports between the bottom of the core barrel and the reactor vessel. Analysis has been performed to demonstrate that the lower radial supports can withstand the resulting impact loads without losing structural integrity. In addition, because of

, the increase in flow velocity entering the lower plenum, this situation would result in a small decrease in reactor flow but this reduction is

, only a small fraction of the current margin between the best estimate flow and the flow used in thermal design and analysis. Therefore, it would not be necessary to reduce thermal design flow in order to evaluate the effect of this scenario on safety analysis results. There is however, an increase in the maldistribution of the core inlet flow and this would need to be considered in the evaluation. Core performance evaluations have shown that the estimated increase in core inlet flow maldistribution can be accommodated without impacting core limits for normal operating conditions.

3580s-020289 10 11-3