ML20151B600

Core Spray Line Crack Analysis for Brunswick Steam Electric Plant,Unit 2

ML20151B600
Person / Time
Site:	Brunswick
Issue date:	03/31/1988
From:	Cornwell K, Stevens G, Tran P GENERAL ELECTRIC CO.
To:
Shared Package
ML20151B599	List: ML20151B596 ML20151B600
References
DRF-E21-00094, DRF-E21-94, EAS-14-0388, EAS-14-388, NUDOCS 8804110052
Download: ML20151B600 (40)

Text

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O EAS-14-0388 DRF-E21-00094 March 1988 O-CORE SPRAY LINE CRACK ANALYSIS c', FOR BRUNSWICK STEAM ELECTRIC PLANT UNIT 2 O Prepared by:

K.F. Cornwell G.L. Stevens P.T. Tran O

Reviewed by:

4, l.E. Nithols , 'Jr. , Southofn Region Licensing Services Manager

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Approved by:

g S. Ranganath, Manager i

Structural Analysis Services i

Approved by:

G.L. Sozzi, Manager g Plant Perfortnance Engineering

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IMPORTANT NOTICE REGARDING CONTENTS OF THIS REPORT O

PLEASE READ CAREFULLY g The only undertakings of General Electric Company respectin5 information in this document are contained in the contract between the customer and General Electric Company, as identified in the purchase order for this report and nothing contained in this document shall be construed as changing the contract. The use of this information by anyone other than the customer or for any purpose other than that for which it is intended, is not authorized; and with respect to any unauthorized use, General Electric Company makes no representation or warranty, and assumes no liability as to the completeness, accuracy, or usefulness of the information contained in this document.

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- EAS-14-0388 g TABLE OF CONTENTS ZA&R O

1. INTRODUCTION AND

SUMMARY

1-1 1.1 Crack Leakage Estimate 1-1 1.2 Structural Analysis 1-2 1.3 Lost Parts Analysis 12 1.4 Effect on LOCA Analysis 1-2 1.5 Conclusions 1-3

2. CRACK LEAKAGE ESTIMATE 2-1 C

2.1 Current Leakage Rate 21 2.3 Maximum Expected Leakage Rate 2-2 2.2 Conclusions 2-3

3. CORE SPRAY PIPE STRUCTURAL INTECRITY 3-1 3.1 Potential Cause of Cracking and Likelihood of Crack Arrest 3-1 3.2 Structural Integrity 3-3 3.2.1 Summary ,

3-3 3.2.2 Crack Arrest Evaluation 3-4 3.2.3 Allowable Flaw Size Determination 35 3.2.3.1 Analysis and Results 35 0 3.3 Summary and Conclusions 37

4. LOST PART ANALYSIS 4-1 4.1 Introduction 4-1 4.2 Loose Piece Description 41 4.3 Safety Concerns 41 4.4 Evaluation 4-2 4.4.1 General Description 42 4.4.2 Postulated Loose Pieces 43 ii O

., EAS-14-0388 g- 4.4.2.1 Core Spray Pipe 43 4.4.2.2 Small Pieces 4-3 4.5 Conclusions 45

5. LOSS OF COOLANT ACCIDENT ANALYSIS WITH CRACKS IN CORE 0

SPRAY PIPING 5-1 5.1 Limiting Break Size and Single Failure Analysis 52 5.2 Analysis Results 5-3 g 5.3 Conclusions 54

6. REFERENCES 6-1 g APPENDIX: STRUCTURAL ANALYSIS OF THE BRUNSWICK UNIT 2 CORE SPRAY PIPE A1 O

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EAS-14 0388 g 1. INTRODUCTION AND

SUMMARY

During the current refueling and maintenance outage, the vessel in service inspection identified a crack indication (Figure 1-1) on the north core spray line (header) at Brunswick Steam Electric Plant (BSEP)

Unit 2. The indication was identified using an underwater camera during the inspection in response to IE Bulletin 80-13 (Reference 1). The crack indication is ic,cated outside the shroud where the piping and junction box meet in the heat affected zone (HAZ) of the weld. Additional inspections were performed to confirm whether the indication was indeed a crack, and if so, if it was through wall. Subsequent liquid dye penetrant and ultrasonic testing (PT and UT examinations) revealed the following additional information:

a) The crack is through wall, b) The crack is approximately 3.5 inches in length along the inside diameter of the pipe.

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c) The crack is approximately 1.75 inches in length along the outside diameter of the pipe.

GE Nuclear Energy has performed an evaluation to address the safety

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significance of the through wall crack. The technical basis to support the continued structural integrity of the core spray line for all normal and injection conditions is provided. A discussion of the possible consequences q'

of potential loose pieces from a cracked pipe is also presented. Finally, ,

the consequences of a postulated Loss-of Coolant Accident (LOCA) with a l crack in the core spray piping are discussed.

1.1 CRACK LEAKAGE ESTIMATE G

A bounding calculation to estimate the leakage through the crack, presented in Section 2, demonstrated that the total leakage is well within the margin inherent in the core spray system design and performance l~ evaluations. The results indicate that the leakage through the crack is less than 7 gpm.

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1,2 STRUCTURAL ANALYSIS The structural analy ;is , presented in Section 3, concludes that the integrity of the core spray piping will be maintained for all conditions of O

operation during the next operating cycle (Cycle 8). In addition, potential causes of cracking are discussed, and based on the information available, it is expected that the most likely cause is Intergranular Stress Corrosion Cracking (ICSCC) due to a combination of cold working during fabrication and O

the oxidizing water environment in the core spray line.

1.3 LOST PART ANALYSIS O

Because continued sparger structural integrity was demonstrated, lost parts (loose pieces) are not expected. Nevertheless, a lost parts analysis has been performed and is presented in Section 4. It is concluded that the probability of unacceptable flow blockage of a fuel assembly and O unacceptable control rod interference due to lost parts is negligible. The potential for corrosion or other chemical reaction with reactor materials does not exist because thu piping material is designed for in-vessel use.

It is also shown that loose pieces are not expected to cause damage to the O other reactor pressure vessel internals.

1.4 EFFECT ON LOCA ANALYSIS D Section 5 presents the results of the LOCA analyses. The results show that the inherent conservatisms present in current LOCA analyses more than offset the small amount of leakage through the crack. The impact on ca'.culated peak cladding temperature is shown to be insignificant. It is O concluded that no change to the present Maximum Average Planar Linear Heat Generation Rate (MAPLHGR) for BSEP Unit 2 is required.

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• o EAS-14 0388 O 1.5 c0NCLUSIONS A detailed evaluation of the BSEP Unit 2 core spray pipe crack has been performed. This cvaluation included structural, lost parts and LOCA C analyses to determine the impact on plant operation with the crack in the core spray piping. Based on the analysis, it is concluded that BSEP Unit 2 can safely operate in this condition during the next fuel cycle, and that no operational changes or restrictions are required during that period.

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EAS-14-0388 g 2.0 CRACK I.EAKAGE ESTIMATE There are no direct measurements of leakage from the crack during the opration of the core spray system. However, from previous analyses and tests performed for cracks observed in other BWRs, it is possible to establish an upper bound leakage for the crack identified at BSEP Unit 2.

The significance of previous crack occurrences at other BWRs has been assessed by both visual inspections and air-bubble tests. Based upon these inspections and tests the upper bound leakage was estimated to be less than half the leakage through the 1/4 inch vent hole present in the T box. (The vent hole is part of the ori Einal piping design and is included to allow the release of any non condensibles which could collect in the core spray O

piping). The video from the BSEP inspection indicated that the crack in the BSEP pipe is significantly shorter in length than those observed at other BWRs. Consequently, it is conservative to assume that the maximum leakage from the BSEP Unit 2 crack is approximately one half of the total vent hole leakage.

2.1 CURRENT LEAKAGE RATE The vent hole is a 1/4 inch hole present in the T box. The leakage rate through the vent hole is estimated assuming incompressible Bernoulli flow through the hole:

Q - CA / 2g, AP/p (1) where, C -

flow coefficient (assumed to be 0.6 for an abrupt contraction) b A - area p = mass density of fluid AP - pressure difference across the pipe / vent ,

The flow rate through the vent hole was determined utilizing a differential pressure of 125 psig across the core spray line. This 21 C

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EAS-14 0388 D

corresponds to the differential pressure expected during rated core spray flow conditions. Utilizing the equation above the estimated leakage rate through the vent hole during a LC';A was determined to be less than 13 gpm.

Therefore, during core spray injection phase of a LOCA, the total leakage D

through the crack is expected to be less than 7 gpm (approximately one half of the vent hole leakage).

2.2 MAXIMUM ESTIMATED CRACK LEAKAGE D

In order to estimate the maximum leakage expected through the crack, the configuration for a 180 degree through-wall crack was used. This configuration was considered to be the upper bound based on the crack arrest D

results of Section 3.0. A crack width of 0.01 inch was conservatively assumed based on the results of Linear Elastic Fracture Mechanics (LEIM) methods which showed the crack opening to be < 0.01 inch under the applied loads described in section 3.0. Using the method of Section 2.1 for these C'

loads and the 180 degree through wall crack configuration, the leakage was determined to be 20 gpm.

It was also estimated that the current crack size is expected to grow

' less than 0.5 inch during the next 18 month cycle. This result is based on the use of crack growth rates at high conductivity (= 1.0 S/cm) for low carbon stainless steel, and considers crack growth from both ends of the crack. Since conductivity is expected to be much lower (=0.15pS/cm) due to

@ the addition of Hydrogen Water Chemistry, this crack growth result is considered conservative. Thus, even after 18 months of additional operation, the crack length is expected to be less than 90 degrees of the circumference. Therefore, the leakage estimate of 20 gpm for a 180 degree C crack length is conservative during the next cycle.

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2.3 CONCLUSION

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The estimated leak rate though vent hole is 13 gpm and the maximum leak rate through a 180-degree through wall crack is less than 20 ggm during core spray operation. The core spray system included a design allowance of

' O' apprcximately 100 gpm to allow for leakage through the vent holes and thermal sleeve between the T-box and vessel nozzle. There is no leakage from the thermal sleeve area since the original thermal sleeve design installed in the BSEP Unit 2 has been replaced by a welded design.

Therefore, during a LOCA the combined leakage through the crack, and vent hole is expected to be well within the core spray system design leakage allowance.

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3.0 CORE SPRAY PIPE STRUCTURAL INTEGRITY ,

O The structural integrity aspects of the core ~ spray piping were reviewed to assess:a) the potential crack mechanism, and b) the impact the crack could have on the structural integrity of the piping. Structural analyses were performed to determine the potential sources of stress in the piping, g the potential causes of cracking, and the likelihood of crack propagation.

Although there is currently not enough information to definitively determine the made of cracking, it is expected that the crack is due to an y IGSCC mechanism. The results of the assessment are discussed below.

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3.1 POTENTIAL CAUSE OF CRACKING AND LIKELIHOOD OF CRACK ARREST The core spray line in BSEP Unit 2 where the crack is located is made of Type 316 L grade stainless steel. The low carbon (L grade) stainless steel does not undergo weld sensitization and therefore the veld heat affected zone (HAZ) is unlikely to experience ICSCC. Other plants which have seen crackin5 have had core spray piping made of Type 304 stainless steel which can sensitize, leading to IGSCC in the weld HAZ. If the crack is in fact due to ICSCC, the most likely cause is local cold work on either the outside or inside of the pipe. Potential sources of cold work for this piping include grinding, on either the inner or outer circumference of the pipe, or bending of the piping assembly during installation.

A review of the video tape indicates that no grindin5 was performed on the outside pipe surface, since all the typical characteristics of the weld L,

are evident. It is likely that some grindi.ng was performed on the inside of the pipe (next to the T-box). Grinding in this region is expected because of the requirement for a PT exam on the inner and outer veld surfaces. In order to obtain an effective PT result on a weld root area, grinding is l

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commonly required. Current practice would require that the surface be polished following any grinding, to remove any cold work induced by the grinding. However, at the time the core spray line was replaced in BSEP Unit 2, no such requirement was in place.

O Cold working could also have been caused by the handling of the piping assembly during the fit-up and instaliation process. This piping was heat treated following fabrication, which would correct any cold work which might have occurred during the fabrication process. (Improper heat treatment of O

the piping is highly unlikely to result in sensitization due to the very low carbon content of the 316L material.) However, handling of the piping during installation could have introduced some cold work if the piping assembly was bent during fit-up. If bending had occurred, the cold work n"

would most likely be induced on both the outside and inside of the pipe because of the relatively small pipe wall thickness (approximately one quarter of an inch).

Such cold work induced IGSCC is concistent with a previous crack occurrence in 316L material at another WR . This incident involved a safe end thermal sleeve crack which was verified as cold work induced IGSCC due to grinding, based on a metallurgical analysis of a boat sample.

O An additional contributing factor to IGSCC may be the environmental condition inside the core spray line. The core spray sparger is located in this upper plenum directly above the core. Radiolysis of water during the n"

boiling process can liberate gaseous hydrogen and oxygen which can be carried by the steam into the core spray sparger nozzles. The high point in the core spray line is the T box where ene crack is located. Consequently, any non condensibles generated in the core can travel through the core spray U line and be released from the line through the vent hole in the T box. It is difficult to assess the impact : hat the high concentration of non-condensibles have in aggravating the oxidizing potential and increasing the susceptibility to IGSCC. Peroxides in the core spray line could also increase the IGSCC susceptibility significantly.

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To further substantiate IGSCC as the most likely cause of cracking, another potential crack mechanism, thermal fatigue, was evaluated. The postulated thermal fatigue mechanism could be due to the turbulent mixing of the colder feedvater with the downcomer flow. This mixing could cause rapid O

thermal cycling in the core spray line provided a sufficient temperature difference existed. If this occurred, the most likely crack location would be in the weld region at the junction between the T-box and the piping, where the residual stresses are highest. The potential for this type of

1. thermal fatigue cracking, however, was judged to be very low for several reasons. First, if thermal cycling had occurred it is expected that a more uniform cracking of the piping would have occurred, instead of the isolated crack which was identified. Secondly, thermal fatigue cracks would be b

expected to have initiated from the outside of the piping, since the thermal stresses attenuate rapidly through the thickness of the pipe. The PT and UT examinations revealed that the crack was longer on the inner diameter of the piping, which indicates that the crack was probably initiated from the O inside surface. Based on the information above, it is unlikely that the crack was induced because of thermal fatigue.

3.2 STRUCTURAL INTEGRITY 3.2.1 Summary With the exception of weld residual stresses, all identified stresses

'O expected during normal reactor operation were found to be small. Based upon a review of these stresses, it is concluded that the structural integrity of the piping with the crack will be maintained during core spray injection.

The stresses considered include those due to downcomer flow impingement M loads, seismic loading, pressure, weight and thermally induced loads.

Although the normal operating loads by themselves do not result in stresses which are sufficient to cause IGSCC initiation, the addition of the Il weld residual stresses coupled with local cold work could result in exceeding the initiation threshold. Once initiated, the normal operating 33 C

O EAS-14-0388 load stresses could cause subsequent growth of cold work induced cracks.

O The improved water conductivity and the planned introduction of Hydrogen Water Chemistry will aid in mitigating future crack growth.

In order to determine the integrity of the core spray line with the O crack, a crack arrest evaluation was performed. The stresses due to pipe restraint and the fabrication residual stresses were also included in this evaluation. Because the applied normal loading is predominantly displacement controlled, the stresses relax as the crack grows and the compliance (or flexibility) of the pipe increases. The results of the analysis showed that when the crack reaches 180' of the circumference, the compliance is reduced sufficiently to relieve almost all of the displacement controlled stresses. Consequently, the crack growth is expected to 9 negligible or at virtual arrest prior to reaching 180 degrees. (The current

. through wall affected area is less than 90 degrees of the piping circumference.)

3.2.2 Crack Arrest Evaluation Stresses in the core spray piping due to bracket restraint are governed by the applied displacement and the compliance of the pipe. Since the  :

T.) displacement is fixed, the compliance change with crack growth could lead to crack arrest. This is comparable to crack arrest in a bolt loaded  ;

vedge opening loading (WOL) specimen in stress corrosion tests.

., Figure 3 1 shows the variation of compliance with crack length for a pipe subjected to bending. The compliance was determined using the relationship between the strain energy release rate, C; and the compliance change per unit area of crack extension de/dA (Reference 2). For the cracks 3 in the core spray line, L/d is expected to be in the range of 0 < L/D < 40.

Figure 3-1 shows that the compliance of the pipe increases by a factor of ten when more than 30% of the pipe is cracked. Therefore, for the given initial displacement, the stress in the core spray line and the applied stress intensity factor would decrease by a factor of ten when more than 30%

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, of the pipe circumference is cracked. Clearly, when the crack length ,

exceeds this value, the restraint stresses become negligible and crack arrest is expected. Therefore, crack arrest is expected before the crack grows to 180 degrees.

'O 3.2.3 Allowable Flaw Size Determination Even though the cracks are expected to self arrest at 180' under the sustained displacement controlled loadin5 as discussed in Section 3.2.2, an P'

evaluation was performed to determine the maximum allowable circumferential through wall flaw size in the core spray pipe. This analysis provides an assessment of the safety margin in the pipe due to primary loads such as deadweight, pressure, flow impingement and seismic.

- The acceptable through wall flaw size of the core spray line is determined utilizing the net section collapse formulation of Reference 3.

To apply this methodology, primary membrane stresses in the longitudinal direction and primary tending stresses were determined for the T-box region of the pipe. A finite element model of the core spray pipe was developed to ,

obtain the stresses due to deedweight, seismic, and reactor vessel downcomer flow impingement on the pipe at the location of interest. The resulting stresses were then combined with the stresses due to pressure and core spray flow loads in order to get the total stresses acting on the pipe. Stresses due to thermal mismatch were not included since they would impose second order effects. By applying these resulting primary stresses, utilizing tha  !

[ methods of Reference 3, it was shown that the core spray pipe can tolerate a crack up to 235' through wall at the T-box location without an incipient failure.

J C 3.2.3.1 Analysis and Results

A finite element model of the core spray line configuration was constructed using the ANSYS computer code (Reference 4). A sketch of the 35

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finite element model is shown in Figure 3-2. The following boundary O

conditions were applied to the model:

Nodes 1, 49, 53: completely fixed Nodes 13, 37  : fixed in vessel radial direction to O'

account for bolted vessel clamps.

Loads due to the weight of the pipe (including captured water in the pipe) were applied to the model along with vertical and horizontal seismic

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loads and reactor vessel downcomer flow impingement loads. Calculations of these loads are given in the Appendix. The largest resulting stresses in the region of the T-box (nodes 24 26) were used from the finite element n

model results and were combined with the stresses due to pressure and core spray flow loads. The resulting total stresses are shown in Table 3.1.

Note that loads due to thermal mismatch of the core spray line and reactor vessel need not be included as they are secondary in nature.

TABLE 3.1 RESULTING PRIMARY STRESSES AT TEE BOX kEGION Membrane Stress, P, 927 psi 9 Bending Stress, P b 2,243 psi L'

The stresses of Table 3.1 are utilized to determine the acceptable through wall flaw size based on the methods of Reference 3. The acceptable flaw size is determined by requiring a suitable design margin on the critical flaw conditions. The critical flaw size is determined by using O limit load concepts. It is assumed that the pipe with a circumferential crack is at the point of incipient failure when the net section at the crack develops a plastic hinge. Plastic flow is assumed to occur at a critical stress level, af, called the flow stress of the material. For ASME Code analysis, og may be taken as equivalent to 3S,. This results in considerable simplification of the analysis.

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O EAS-14 0388 Consider a circumferential crack of length, 1 - 2Ra , and constraint

'm depth, d, located as shown in Figure 3 3. In order to determine the point at which collapse occurs, it is necessary to apply the equations of equilfbrium assuming that the cracked section i ahaves like a hinge. For this condition, the assumed stress state at the cracked section is shown in O

Figure 3-3 where the maximum stress is the flow stress of the material, ag.

Equilibrium of longitudinal forces and moments about the axis gives the following equations:

O (For neutral axis located such that a + p < x)

A - [(x - ad/t) - (P,/ag)w)/2 o"

Pb ~ (2# f /x) (2 sin B - d/t sin a) where, t - pipe thickness, inches.

a - crack half angle as shown in Figure 3-3.

G $ - angle that defines the location of the neutral axis.

Using the stresses of Table 3.1 and a d/t ratio of 1.0 (through-wall y flaw), the allowable through-wall crack for which failure by collapse might i

occur is 235*.

O 3.3

SUMMARY

AND CONCLUSIONS The potential sources of stress in the piping resulting from fabrication, installation, normal operation, and operation during postulated 3 1.o s s of Coolant Accidents were reviewed. Potential causes of cracking, thermal fatigue and ICSCC, and the likelihood of crack propagation were also evaluated. It is expected that the crack was caused by IGSCC due to local cold work in the piping.

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O EAS-14 0388 Because of the predominant secondary stressas, the crack can be G expected to arrest prior to reaching 180*. An assessment was made to determine the critical flaw size of the core spray pipe by treating stresses associated with the design loadings as primary stresses and performing a net section collapse evaluation. The results of this evaluation confirm that a O through wall crack of up to 235' around the circumference would not cause pipe failure. Therefore, it is concluded that the structural integrity of the piping with a crack will be maintained for all conditions of operation for the next operating cycle.

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O EASo14 0388 N 4 14ST PARTS ANALYSIS

4.1 INTRODUCTION

Based on the structural analysis given in Section 3, it is expected Cl that the BSEP Unit 2 core spray pipe will not break and consequently, will not result in loose pieces in the reactor. However, an evaluation of the possible consequences of a potential loose piece is presented in this section.

O 4.2 LOOSE PIECE DESCRIPTION Since a piece has not been lost, it cannot be uniquely described. Two S different types of loose pieces are postulated:

1) a section of core spray pipe; and, 0 2) a small piece of the core spray pipe.

4.3 SATETY CONCERNS O The follouing safety concerns are addressed in this analysis:

1. Potential for corrosion or other chemical reaction with reactor materials, s]

2. Potential for fuel bundle flow blockage and subsequent fuel damage.

O 3. Potential for interference with control rod operation.

4 Potential for dataage to other reactor internals.

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.O EAS-14-0388 4.4 EVALUATION O

The above safety concerns for the postulated loose pieces are addressed in this section. The effect of these concerns on safe reactor operation is also addressed.

O 4.4.1 Ceneral Descrio';12D Since the core spray pipe with the crack is in the annular region of the reactor pressure vessel, this evaluation assumes that any potential loose piece generated from the core spray pipe will most likely sink into the downcomer region.

O For a loose part to reach, and potentially block the inlet of a fuel

- assembly (Figure 4.1), it would have to be carried into the lower plenum.

To accomplish this, it would have to be carried by the recirculation flow through the jet pump nozzle into the lower plenum, then make a 180' turn and O be carried upward to the fuel assembly inlet orifices.

For a piece of the core spray pipe to reach a control rod it must first inigrate to the lower plenum, pass through the fuel inlet orifice, and C' traverse the fuel bundle. Then, it must either fall through the restrictive passage between two fuel channels, or fall through an opening between the peripheral bundles and the core shroud. Both of these potential paths are unlikely.

O The core spray pipe is fabricated from both Type-304 and 316L grade stainless steel and all parts of the core spray pipe are designed for in reactor service. Consequently, there is no postulated loose part that C will cause any corrosion or other chemical reaction with any reactor material.

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O EAS 14 0388 g 4.4.2 Postulated Loose Pieces .

4.4.2.1 Core Spray Pipe The core spray pipe is 5 inch Schedule 40 pipe. In order to generate a loose piece of pipe, a minimum of two through vall cracks would have to propagate 360* around the pipe.

If a Pi P6 segment were postulated to break off, it would sink into the O

downcomer region. Since it cannot fit through the jet pump, it cannot enter the lower plenum, and therefore will not cause any flow blockage at the fuel inlet orifice. Since it is too large to fit between fuel channels, it can not cause any interference with control rod operations. Nevertheless, due to the slow propagation rate of potential cracks, and based on previous experience with cracks in core spray spargers, it is judged that a piece of the piping will not break off and become loose.

G 4.4.2.2 Small Pieces In order to generate small pieces of the core spray pipe, both longitudinal and circumferential through wall cracking must occur. A small piece could then sink, be carried into the downeomer annulus, pass through the jet pump and enter the lower plenum. A piece that entered the lower plenum would probably be driven by the jet pump flow to the bottom of the reactor pressure vessel where it would be expected to remain. However, a small piece < 0.4 inches could be carried by the flow up to the fuel inlet orifices. The orifice sizes in BSEP Unit 2 vary from approximately 1.2 to 2.1 inches in diameter.

Given the dimensions, the piece would pass through the inlet orifices and be trapped at the lower tie plate grid and cause some bundle flow blockage. However, the flow blockage is much less than that required to initiate critical boiling transition in the bundle. Multiples pieces migrating to the same bundle may result in critical flow blockage, but the probability for such an occurrence is extremely low.

4-3 O

m -- - y, y - . - - ." , _ . _ , -. _

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'D EAS 14 0388 "g

It is also very unlikely that a small piece could lift and migrate from the lower plenum through the fuel bundle and fall into the control rod guide tube. In order to do this, the piece would have to be so small that it could pass through all the bundle spacers and cut through the top of the

'O bundle. Such a small piece would not present any potential for control rod i interference.

Figure 4-3 shows a typical unit cell of four fuel assemblies and one g

control rod. The control rod moves in the gap between the fuel channels.

There is a small possibility that a piece small enough to fit in the gap between the channel vall and control blade could sink and pass through the

, cavity between the control blade and the fuel support casting and migrate G,

into the control rod guide tube. Should this happen the piece will most likely come to rest on the top of the velocity limiter where it is expected to remain and move only with the movement of the velocity limiter as the control rod is inserted or withdrawn. If the piece is small enough to pass between the velocity limiter and the guide tube wall it will most likely sink and come to rest at the bottom of the guide tube. Due to the hardware geometry of the control blade drive mechanism it is highly unlikely that any piece would be small enough to migrate into the control blade drive system.

lD Thus, any potential small piece which migrates to the control rod guide tubo I

is not expected to pose any concern for potential interference with control

! rod operation.

One of the licensing bases of the reactor is that with the highest worth control rod fully withdrawn the reactor can be brought to cold

shutdown. Thus, unacceptable control rod interference would require multiple precisely sized pieces interfering simultaneously with control rods

)q' that are in close proximity to each other. The probability of this is

{ judged to be insignificant.

I i

44 O

l

O EAS 14 0388

4.5 CONCLUSION

S The core spray pipe at BSEP Unit 2 is expected to remain intact; therefore, it is hi 5hly unlikely that pieces of the core spray pipe will O break off. From the above evaluation it is concluded that the probability for unacceptable corrosion or other chemical reaction due to loose pieces is zero. The potential for unacceptable flow blockage or other damage to the fuel assemblies is negligible. The potential for unacceptable control rod O

interference is negligibly small. Therefore, it is concluded that there is no safety concern posed by any postulated loose parts.

0 9

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EAS-14-0388 i

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CORE SPRAY SP ARGE R UPPE R PLE NUM If il

, f) ANNULUS ,

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' e'- n LOWIM Tit PLATE ll

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LOWER PLENUM ,

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.-l FIGURE 4-1 l LOOSE PIECE POTENTIAL UPWARD FLOW PATH l 4-6 t.

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TUBE 1 .

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FIGURE 4-2 ORIFICED FUEL SUPPORT C.

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11. LOWER

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13. FUEL PELLETS 14.END PLUG i I, lki fi

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GENER AL h ELECTRIC .

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O EAS 14 0388

'7 5.0 LOSS OF COOIANT ACCIDENT ANALYSIS WITH CRACKS IN CORE SPRAY PIPING The estimated leakage through the crack is expected to be less than 7 gpm. This is insignificant when compared to the rated flow assumed in the

~

reload licensing analyses of 4625 gpm per core spray system. The core spray system includes a design allowance of approximately 100 gpm to allow for leakage through the vent hole and the thermal sleeve originally located between the T box and vessel nozzle. The original thermal sleeve design installed in the BSEP Unit 2 has been replaced by a welded design and hence there is no leakage. Therefore, during a LOCA the combined leakage through the through wall crack and the vent hole (20 gpm) will be well within the margin (100 gpm) inherently assumed in the core spray system. Thus, there should be no impact on either core spray or ECCS performance. Considering

- the rated flow of a core spray system is actually 4725 gpm (the 100 gpm i

leakage allowance is conservatively assumed in the licensing analyses), the offects are negligible.

From 'setion 3, .

the crack growth is expected to arrest prior to reaching 180 degrees. The maximum potential leakage through a 180 degree crack is also expecte2 to be 20 gpm. This maximum leakage combined with the vent hole leakage is 33 6;7 which is well within the 100 gpm design allowance. Fevertheless, this section describes the LOCA analyses which were perforacd to conservativr.ly bound the consequences of continued operation with the crack.

J This section describes the methods used to evaluate the MAPillGR i requirements to meet 10CFR50.46 for the BSEP Unit 2 fuel cycle 8 with cracks in the core spray piping. The potential effect of the cracks on the 3 limiting break size and single failure is discussed in Section 5.1. The results of the LOCA evaluations are given in Section 5.2, and the  ;

conclusions are presented in Section 5.3.

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=

9 EAS 14 0388 5.1 LIMITING BREAK SIZE AND SINGLE FAILURE ANALYSIS O

For BSEP Unit 2, there are no ' single failures for any location (other than a core spray ine break) that can result in less than one core spray system injecting water into the upper plenum above the reactor core. For a

@ core spray line break, there are always at least three low pressure ECCS pumps injecting water into the reactor vessel, thereby ensuring that- this break is not a limiting event. For medium and large break sizes (which depressurize relatively quickly), the most limiting failures are those that

,O result in the least number of ECCS pumps remaining operable (i.e., injecting water into the reactor vessel).

The only two single failure candidates that are potentially limiting for recirculation pipe breaks aro:

A. Diesel Generator Failure. This postulated f ai?,ure leaves 1 core spray (LPCS) + 1 Low Pressure Coolant Inj ection (LPCI) + High Pressure Core Injection + the Automatic Depressurization System operable; B. LPCI Injection Valve Failure. This leaves 2 core spray (LPCS) +

r.

'~'

HPCI + the ADS operable.

Since the High Pressure Coolant Inj ection (HPCI) is steam turbine powered, it is not a significant contributor to mitigating medium to large f5 breaks which depressurize rapidly. Also, since the function of the Automatic Depressurization System (ADS) is to depressurize the reactor as a backup to the HPCI, it contributes little toward mitigating medium and large break LOCAs. Therefore, failure candidates A and B are similar and each rasult in a dependence on only two ECCS pumps.

The plant specific LOCA analysis based on the SAFE and REFLOOD codes (Reference 6), indicates that failure candidate B (LPCI Injection Valve failure) is by far the most limiting. This is primarily due to the 52 0

. e .

O' EAS 14 0388 conservative modeling of counter current flow limiting (CCFL) at the fuel

'O assembly upper tie plates. The coda calculation limits the coolant delivery or downflow from the core spray systems to the fuel bun'dles , and further delays core reflooding by neglecting the ref:ocding contribution of the water held back in the upper plenum upon CCFL breakdown.

10 Both single failure candidates (A and B) were re examined for large breaks to determine whether there would be a change in the limiting case assuming an additional 100 gpm flow reduction in the core spray system with

(

a crack. The limiting single failure, break size, and location were found not to change. This is due to the conservative treatment of CCFL that neglects the core spray water held back in the upper plenum, which is approximately 20% of the total core spray flow.

C

- 5.2 ANALYSIS RESULTS It should be noted that during postulated 'LOCA the' core spray pumps O

would be operating close to their run out flow capability (=6500 gpm). The combined effects of a high flow and the significantly lower developed pump head (or pressure in the core spray line) would further support that the leakage is negligible.

O In order to substantiate that the leakage expected from the crack at BSEP Unit 2 will have a negligible impact on the current LOCA response, a bounding ECCS evaluation was performed. This evaluation entailed

@ reanalyzing the limiting Design Basis Accident (DBA) LOCA event assuming an i

additional 100 gpm reduction in the core spray flow rate. This represents twice the design allowable value, or a reduction in the core spray flow rate of approximately 30 times the estimated leakage for the crack.

.' O The results of this evaluation demonstrated that the effect on the limiting LOCA response is negligible. Thus, it is concluded that the current MAPMGRs for BSEP Unit 2 (Cycle 8) remain identical to those documented in the Cycle 8 reload evaluation. Consequently, no change to the BSEP Unit 2 MAPMGR Technical Specifications are required.

53

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Q EAS-14 0388 l

5.3 CONCLUSION

S LOCA analyses were performed assuming an additional 100 gpm flow p leakage for the core spray system with the crack in the piping outside the core shroud. The results of the analyses demonstrate that the present r61oad licensing analysis conservatively bounds the effects of leakage through the crack, and that no MAPLHCR reduction is required.

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. EAS 14 0388 C'

6. REFERENCES

1. USNRC IE Bulletin No. 80 13, Cracking in Core Spray Sparger, May 12, 1980.

2. E. Kiss, J.D. Heald, D.A. Hale, "Low Cycle Fatigue of Prototype Piping," GEAP-10135, January 1970.

3. Ranganath, S. and Mehta, H.S., "Engineering Methods for the Assessment of Ductile Fracture Margin in Nuclear Power Plant Piping,"

Elastic Plastic Fracture: Second Symposium, Volume II - Fracture

.O Resistance Curves and Engineering Applications, ASTM STP 803, C.F. Shih

- and J.P. Gudas, Eds., American Society for Testing and Materials,1983, pp. II 309 - 11 330.

4. Gabriel J. DeSalvo, Ph.D. and John A. Swanson, Ph.D., "ANSYS Engineering Analysis System Ussr's Manual," Revision 4.1, Swanson Analysis System, Inc., Houston. ?A, March 1, 1983.

.C-NEDO 10174, Rev.

5. 1 "Consequences of a Postulated Flow Blockage Incident in a Boiling Water Reactor," General Electric Company, October 1977.

C

6. SER, 0.D. Parr (NRC) to C. G. Sherwood (GE), "Review of General Electric Topical Report NEDO 20566, Amendment 3," June 13, 1978 G

O 61 C,

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e EAS 14 0388 Q

APPENDIX O

STRUCTURAL ANALYSIS OF THE BRUNSWICK UNIT 2 CORE SPRAY PIPE O

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Ot F.AS 14 0386 Q

The stress results given in Table 3 1 of Section 3.2.3.1 of this report are developed in this Appendix. The stresses were determined by appif ing dead weight, seismic and flow impingement loads to the sinite element model p developed for the core spray line (see Figure 3 2). The calculation of these loads is given here. Also included are the calculations of the stresses due to pressure and flow loads as well as the total combined primary stresses given in Table 3 1, O

Vaimht of linen An equivalent density was input to the ANSYS finite element model to

include both the weights of the pipe and captured water. This equivalent density is calculated below

Metal density - 0.2879 lb/in 3 g

Water density - 62.4 lb/ft3 - 0.0361 lb/in3 Pipe size - 5 inch schedule 40 OD - 5.563", t - 0.258", ID - 5.047" Metal area - (w/4) (5.563 2 5.0472 ) - 4.3 in2 Water area - (w/4) (5.0472 ) - 20.0 in 2 Metal weight - (0.2879 lb/in 3) (4.3 in 2 ) - 1.238 lb/in Water weight - (0.0361 lb/in 3 ) (20.0 in 2 ) - 0.722 lb/in Adjusted density - (total weight)/(metal area)

~

- (1.238 + 0.722)/4.3

- 0.456 lb/in3 - 0.0012 slugs /in 3 A1 u

• *e C) [

. EAS.14 0388  ;

6 Imoinnement Leads (90' Deflection of Flow)?

2 c F - PA - pV DL/g D - 5.563"/12 - 0.464 ft

(7  !

Assume downconer flow, V - 5 ft/sec (conservative)

For water, p - 62.4 lb/ft3 O F/L - pV2 D/g - (62.4)($2 )(0.464)/32.2 - 22.5 lb/ft - 1.87 lb/in The nodes of the 'inite element model are spaced 5' apart. Thus, the following load will be applied to all nodes comprising the horizontal arms

'G of the core spray line (nodes 10 40);

a  !

~

Node spacing - Rf - (106.0") (5') (w/180') - 9.25" [

Load per node - (1.87 lb/in) (9.25") - 17.1 lb {

4 .

1.s l Seismie Leads! [

F i

l From the Brunswick FSAR, the vertical acceleration for a Design Basis t

,Q Earthquake (DBE) is 0.16 g and the horizontal acceleration is small and negligible. In order to provide conservative and bounding results, seismic f

i coefficients for similar plants were also considered. Based on a review of those coefficients, the following were selected for use in this analysis:

50 E

! Vertical - 0.16 g 1

l Horizontal - 1.70 g I  !

ij The following accelerations were therefore applied to the finite i element model:

Total vertical acceleration - Weight + Sefamic f

- 1.0 g + 0.16 g -

- 1.16 g - 447.8 in/see l  ;

I r

Total horizontal acceleration - 1.20 g - 463.2 in/sec 2  !

x, A'2 l

,a o ,

O EAS 14 0388

C This acceleration was applied in both the X and Y directions of the model such that the resultant was 463.2 in/see2. Thus, the horizontal acceteration applied to both directions of the model was

O 463.2//2 - 327.5 in/secA Pressure / Flow Loadst C -

Rated flov - 4625 gpm through line i

3 1

Q - (4625 gpm) (1 min /60 see) (1 ft3 /7.48 gal) - 10.31 fc /see G.

F = pQV - pQ(Q/A) - (62.4) (10.31 2 )/[(w/4)(5.047/12)2] - 1,483 lb l

AP - 113 psi @ 4625 gpm

C Stresses Due to Loadst l

Pressure:

ap - (113) (w/4) (5.0472) / ((w/4) (5.563 2 5.0472 ))

- 526 esi Flow Load: '

3 ap - F/A - 1,483/((w/4) (5.562 2 5.0472 ))

- 345 est

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l,

)

A3 L

4 a. * .  :

4 (T t EAS.14 0388

• l 0

Impingement, weight and seismic:

Stress due to the above 1cadin* was determined using the finite element model of the internal cere spray piping. From the ANSYS results, the maximum of the stresses at nodes 24 26 were used since they are in the area of the cracks. The maximum' stresses are given below:

DIR - Axial Stress - 55.9 psi

'g) aBEND - Bending Stress - 2239.5 psi aTOR - Torsional Stress - 123.0 psi

' TAU - Shear Stress - 187.5 psi

?;

Combining all of the primary stresses, the following values are obtained:

Primary Membrane - P,- aF + 'p + 'DIR

- 345 + 526 + 55.9 .

- 927 nat, i

4 i Primary Bending -Pb ~ /# BEND + ' TOR'

- / 2239.52 + 123.02 - 2.243 esi ,

i The shear stress ' TAU' is small and its effect is negligible so it is ,'

not included.

I i

i s

A.4 w - - . -