Text

Nonproprietary Fracture Toughness Evaluation of Shell Side Lpcis Heat Exchanger for Dresden Station - Unit 2
	ML20067B492
Person / Time
Site:	Dresden
Issue date:	02/04/1991
From:	; AEA O'DONNELL, INC. (FORMERLY SMC O'DONNELL, INC.
To:	;
Shared Package
ML17202U972	List: ML20067B492;
References
	NUDOCS 9102080270
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NON FROPRIETARY VERSION FRACTURE TOUGHNESS EVALUATION OF THE SHELL SIDE OF THE LOW PRESSURE COOLANT INJECTION SYSTEM (LPCI)

HEAT EXCHANGER FOR DRESDEN STATION - UNIT 2 2042.04 400-002-00 NON PROPRIETARY VERSION Page 1 of 61 9102080270 910204 PDR ADOCK 05000237 P PDR

i NON PROPRIETARY VERSION FRACTURE TOUGHNESS EVALUATION OF THE SHELL SIDE OF THE LOW PRESSURE COOLANT INJECTION SYSTEM (LPCI)

HEAT EXCHAllGER FOR DRESDEN STATION - UNIT 2 Table of Contents P.HR 1.0 INTR 000CT10N.............................................. 4 2.0

SUMMARY

OF RESULTS........................................ 5 3.0 GENERAL DESCRIPT10N....................................... 6 3.1 LPCI-HX Design....................................... 6 3.2 LPCI-HX Materials.................................... 7 4.0 ANALYTICAL METH005........................................ 9 5.0 OVERALL AXISYMMETRIC ANALYSIS OF SHELL, TUBESHEET, TUBE BUNDLE AND HEAD........................... .......... 16 5.1 Development of the Axisymmetric Finite Element Model. 16 5.2 Base Case loads...................................... 20 5.3 Operational Loads.................................... 22 5.4 Axi symmetric Analysi s Resul ts . . . . . . . . . . . . . . . . . . . . . . . . 22 6.0 3-0!MENSIONAL FINITE ELEMENT ANALYSIS,.................... 49 6.1 Development of 3-D Finite Element Model . . . . . . . . . . . . . . 49 6.2 Base Case loads...................................... 49 6.3 Operational loads.................................... 50 6.4 3-D Analysis Results................................. 50 7.0 FRACTURE MECHANICS ANALYSIS.............................., 57 8.0 RrFERENCES... .......................................... 60 2042.04-400-002 00 NON PROPRIETARY VERSION Page 2 of 61

NON PR'J RIETARY VERSION List of ADoendices APPENDIX A AX! SYMMETRIC TUBESHEET FINITE ELEMENT MODEL APPENDIX sppendix A.1 Finite Element Model Material Properties Appendix A.2 Adjusted Material Properties for Stud Hole Region Appendix A.3 Tubesheet and Tube Data Appendix A.4 Stud Bolt Properties Appendix A.5 Calculation of Closure Stud Preload Appendix A.6 Calculation of Tube Temperature Appendix A.7 Axisymmetric Analysis Results APPENDIX B 3 0 FINITE ELEMENT MODEL APPENDIX Appendix B.1 3-0 Finite Element Model Material Pros erties Appendix B.2 Calculation of Blowoff Pressure Appendix B.3 3-0 Analysis Results APPENDIX C FRACTURE MECHANICS ANALYSIS APPEN0lX 9

c 2042.04-400-002-00 NON PROPRIETARY VERSION Page 3 of 61

NON PROPRIETARY VERSION FRACTURE TOUGHNESS EVALUATION OF THE SHELL SIDE OF THE LOW PRESSURE COOLANT INJECTION SYSTEM (LPCI)

HEAT EXCHANGER FOR DRESDEN STATION - UNIT 2

1.0 INTRODUCTION

This report presents the results from the fracture toughness evaluation of the shell side of the Low Pressure Coolant Injection (LPCI) system heat exchanger (HX) for Dresden Station Unit 2. This evaluation was performed to resolve the remaining open item for Commonwealth Edison Company's (Ceco) review of the Systematic Evaluation Program (SEP), Topic III-1, Classification of Structures. Comoonents and Systems which is to demonstrate that the shell sides of the LPCI-HXs have adequate fracture toughness. in NUREG-0577 , the USNRC calculates an average nil ductility temperature (NOT) of 40*F for plain carbon or mild steels, a group that includes A 212B, the LPCI-HX shell material. From this, the USNRC has concluded that material in this group will have adeqtate fracture toughness if its temperature is maintained above 77'F. A review of the LPCI system showed that the temperature of the LPCI-HX shell will be the same as the torus temperature. The torus temperature, as determined by a recorder chart from February 1989 dropped as low as Sl'F , Since this is less than 77'F, CECO needed to determine whether the shell material has adequate frccture toughness at this temperature. The task of resolving this issue was assigned to by CECO as a request to perform a fracture toughness analysis of the LPCI-HX's shell side.

To perform the analysis, computer assisted calculations including 2-0 and 3-O finite element analyses as well as hand calculations are employed. They are utilized in compliance with NUREG-0577 and with the ASME Boiler and Pressure Vessel Code, Section !!! Class C . The 2-D finite element analysis presented in Section 5.0 of this report is an overall axisymmetric analysis of the shell, tubesheet, tube bundle and head.

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i NON PROPRIETARY VERSION lts purpose is to simulate the effects from the interaction between sections of the HX (i.e., tube /tubesheet/shell interactions) and to calculate the maximum stress in the tubesheet caused by the structural (mechanical) and thermal load conditions that the HX is subjected to. Displacements from the 2-D finite element model were then applied to the detailed 3-D finite element model at nodes which represent common locations in the two models.

Tha 3-0 quarter symmetry model includes structural discontinuities in the HX that could not be modeled in the 2-D axisymmetric model because of their lack of axisymmetry to the vessel. The discontinuities analyzed are the shell side outlet nozzle and the vertical support bracket. By applying the displacements from the 2-D model to the locations common in the 3 D model and subjecting the 3-0 model to the same loading conditions that are applied to the 2-0 model, the maximum principal stress in the shell is calculated.

This principal stress is used along with the material toughness, KIC'

, to determine the critical crack size for failure to cccur, A cr. Using the maximum stress range and the crack growth rate equation , the number of cycles, N, required for various assumed initial crack sizes, Aj , to grow to the critical crack size, A

ct, is calculated. This provides the desired fracture toughness characteristics of the LPCI HX's shell side material, A-212B.

2.0

SUMMARY

OF RESULTS The LPCI-HX was analyzed by cycling it between steady-state operation and shutdown (assuming no flow and a worse case service temperature of 51*F throughout the vessel). The results from this analysis show that based upon.

an initial crack size of 0.010 inch (chosen as the maximum track size that could go undetected by inspection of the unit at the time of construction) and a cold shutdown temperature of 51*F, the number of cycles (one cycle defined as steady-state operation to shutdown) that would be required to propagate a crack of this size to the critical crack size of 5.25 inches is 2042.04-400-002-00 NON Pr.JPRIETARY VERSION Page 5 of 61

I NON PROPRIETARY VERSION approximately 1.5 x 108 Based solely from this fracture mechanics analysis and a system startup and shutdown cycle occurring on an average of once a month, the acceptable number of cycles is far beyond any anticipated service life of the i.PCI-HX as well as several orders of magnitude beyond the maximum conceivable plant lifetime including life extension. Therefore, the fracture toughness of the shell side of the LPCI-HX is more than adequate.

3.0 GENERAL DESCRIPTION 3.1 (PCI-HX Desian The Low Pressure Coolant Injection (LPCI) system heat exchanger is a two tube pass, split flow, shell and-tube heat exchanger with the top stationary head having its channel integral with the fixed top tubesheet and the pass partition and having a removable channel cover. The bottom stationary head has its channel integral with the fixed bottom tubesheet and also has a removable channel cover. See Figure 1. The LPCI-HX has an overall length of 28 feet, 2-5/8 inches and is mounted vertically via three support brackets. The cylindrical shell has a 61 inch I.D. and a 7/8 inch wall thickness.

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NON PROPRIETARY VERSION The two tubesheets have a thickness of 3-3/4 inches each and are both overlayed with 1/4 inch monel clad steel. The tubesheets are welded to both the shell and the head. The unit has 2512 tubes on a 1.00 inch rotated square pitch. Each tube is 22 feet in length, has a 3/4 inch 0.D. and an 18 BWG minimum wall thickness.

Each channel cover is attached using 60 3/8 inch bolts. The tube side inlet and outlet nozzles are 16 inch Schedule 40 pipe with a 1/2 inch wall thickness. The shell side inlet and outlet nozzles are 24 x 18 inch extra strong concentric reducers with a wall thickness of 1/2 inch throughout.

Reinforcement pads, each having a 7/8 inch thickness, are welded to the nozzles for added support Details of the LPCI HX are shown on the Heat Exchanger Specification Sheet and on the following PERFEX Corporation drawings:

Title Drawino No. Rev.

Outline Drawing (as built) K-5009 - 1 & 2 Sheet 1 4 Outline Drawing (as built) K-5009 - 1 & 2 Sheet 2 2 Orientation & Top Channel Assembly 4 5009-2 2 Shell & Bottom Channel Details 4-5009-3 1 and Assembly Tube Bundle Anembly & Details 4-5009-4 3 Gaskets 2-5004-1 0 3.2 LPCI-HX Materials The LPCI-HX is constructed primarily of SA-212, Gr. B. Because of the discontinuance of A212 (identical with ASME Specification SA-212) in 1966

, the specification for High Tensile Strength Carbon Silicon Steel Plates for Boilers and Other Pressure Vessels was replaced by A515 (identical with ASME Specifications SA 515 specification for Carbon Steel Plates for Pressure Vessels for Intermediate and Higher 2042.04-400-002-00 NON PROPRIETARY VERSION Page 7 of 61

NON PROPRIETARY VERSION Temperature Service . Therefore, SA-212. Gr. B material properties are utilized when available; otherwise, SA 515, Gr. 70 material properties are used.

The cylindrical shell, tubesheets, heads, support brackets and reinforcement pads are constructed of SA-212, Gr. B. The tubes are constructed of AL-6XN, a superaustenitic stainless steel which has been developed by Allegheny Ludlum Corporation . Both tubesheets are layered with a monel steel cladding. The inlet and outlet nozzles on the shell side and tube side are constructed of SA-106, Gr. B, a carbon steel used for high temperature service. The bolts used to attach the two covers to their corresponding channels are constructed of SA 193, Gr. 87.

The material properties necessary for the analysis are as follows:

Modulus of Coefficient of Thermal Elasticity Density Thermal Expansion Conductivity Material (osi) .(lh/in3) (in./in./*F) (BTV/in.-s *F)

SA-212, Gr. B 28.8 x 105 0.283 5.76 x 10

5.62 x 10 *

(SA-515, Gr. 70)

Al-6XN 27.0 x 108 0.291 8.50 x 10

1.83 x 10 4 Monel 400 26.0 x 105 0.319 7.74 x 10 8 4.44 x 10 4 (67 Ni - 30 Cu)

SA-106, Gr. B 29,0 x 105 0.283 6.31 x 10 8 6.82 x 10 4 SA-193, Gr. B7 29.2 x 10' O.283 6.31 x 10 5 5.56 x 10 *

(1 Cr - 1/5 Mo)

All material properties are taken from unless otherwise stated.

Material properties are evaluated at 165 F, the maximum shell side operating temperature . AL-6XN material properties, which are unavail-able at this temperature, are evaluated at the ambient temperature of 70*F.

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4.0 ANALYTICAL METHODS The LPCI-HX is analyzed by a combination of hand calculations and detailed finite element analyses. The following is a list of computer programs used in the analyses:

Computer Program Information Forms are given on the pages that follow.

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NON PROPRIETARY VERSION COMPUTER PROGRAM INFORMATION FORM I

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NON PROPRIETARY VERSION COMPUTER PROGRAM INFORMATION FORM (Con't) el i

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NON PROPRIETARY VERSION I

{ 5.0 OVERALL AX1 SYMMETRIC ANALYSIS OF SHELL. TVBESHEET. TVBE BUNDLE AND HEAD 5.1 Development of the Axisymmetric Finite Element Model A two-dimensional axisymmetric finite element model of the heat exchanger is shown in Figures 5.1 to 5.5. This model, hereafter referred to as the tubesheet model, includes the tubesheet, the head, one-half of the shell barrel, the tubes, the monel layer on the tubesheet, and the closure studs.

Since much similarity exits between K 5009-1 and K-5009-2, the two types of containment exchanger units (PERFEX Drawing Numbers K-5009 1 and 2 Sheet 1),

heat exchanger K-5009-1 is chosen to be representative of the LPCI HX for the analyses. The top half of the HX is analyzed because it contains discontinuities (shell side outlet nozzle, vertical support brackets, tubesheet to shell interface and tube side inlet and outlet nozzles) that are not all present in the lower half of the HX. Analyses that are performed using this model include structural (mechanical) and thermal load conditions that the heat exchanger is subjected to. The purpose of the tubesheet model is to apply these conditions to the overall heat exchanger in order to obtain the simulated effects from the interaction between sections of the HX (i.e., tube /tubesheet/shell interaction) and to determine the tubesheet stresses. The nozzles and the vertical support brackets are not included in the tubesheet model since they cannot be axisymmetrically represented in the model. Instead, these regions are analyzed using a 3 D finite element model described later in this report. Displacements from the 2-D model are applied to the 3-0 finite element model in order to simulate the overall heat exchanger corditions at the nozzle and vertical support bracket discontinuities of the HX.

The LPCI-HX is a two-tube pass, split flow, shell and-tube heat exchanger with a pass partition to separate flow in the top head. The pass partition could not be modeled in the tubesheet model because of its lack of axisymmetry. To show that its exclusion from the analysis would add 2042.04-400-002-00 NON PROPRIETARY VERSION Page 16 of 61 l

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( conservatism to the stresses in the tubesheet, two 3-D quarter symmetry models are used. The first model consists of the tubesheet with the attached partition plate and the second model consists of the tubesheet alone. A shell side pressure of 100 psi (an arbitrary value used for comparison purposes only) is applied to both models. The stresses in the tubesheet without the partition plate are approximately two times those in the model with the partition plate. See Figures 5.7 to 5.13. In addition, the exclusion of the partition plate would induce greater rotations into the shell which is also conservative. Therefore, it is concluded that by not including the partition plate in the analysis the stresses that would occur in the tubesheet and shell would be higher than those in reality. This is a conservative approach in the analysis.

The tubesheet model and the two 3-0 quarter symmetry models used in the aforementioned evaluation are analyzed using the a computer program

. The two 3-0 models are made with elastic quadrilateral shell elements . In the tubesheet model, the tubesheet, shell barrel, monel layer and head are represented by 2-0 isoparametric solid elements

. The tubes and closure stud elements are modeled using 2-D elastic beams .

General material properties for the components included in the tubesheet model are given in Section 3.2 of this report. The maximum range of temperature which the LPCI-HX is subjected to is 51*F-165*F. Where 51*F is the assumed lowest HX temperature since a temperature as low as 51*F was recorded in February 1989 and 165'F is the maximum shell side inlet temperature . Since thermal stresses are proportional to the product of the modulus of elasticity, E, and the coefficient of thermal expansion, o, the temperature that maximizes the product Ea is chosen. For this range of temperature, 165*F produces the maximum product of Ea, therefore, material properties are evaluated at 165'F. Since material properties for Al-6XN are unavailable at this 2042.04-400-002 00 NON PROPRIETARY VERSION Page 17 of 61

NON PROPRIETARY VERSION temperature, properties at ambient temperature (70*F) are used. Material properties utilized in the finite element analyses are summarized in Appendix A.I.

Two regions of the tubesheet model which require modifications of the material properties are the stud holes located in the channel flange and the channel cover. Modifications to the general material properties given in Section 3.2 are calculated in Appendix A.2. This set of adjusted properties is used for all load cases. The reduced stiffness in the annulus formed by the stud holes is accounted for by modifying the modulus of elasticity in the radial and axial directions by the ratio of the actual material in the annular stud hole region to the total material modeled in that same annular region. The circumferential modulus is set to 1.0 to simulate the lack of hoop stiffness in the hole region. The density of these regions is also modified by the area ratio described above.

The perforated tubesheets are modeled as equivalent solid plates following the guidelines outlined in Article A 8000 in Appendix A of the ASME Boiler and Pressure Vessel Code The effective boundary of the equivalent material of the tubesheet, R*, is based upon the irregular outer boundary of the unit cells

. The dimensions for this particular case are given in Appendiv 3 of this report. The calculated value of R* is 28.28 inches.

The tubes are modeled as 2-D beams whose properties represent an axisymmetric equivalent to the sum of the actual tubes swept out by the annulus centered on the beam.

A summary of the properties for each beam in the model is given in Appendix A.3. Also included are the initial strains for the beam elements to simulate the elongation or contraction of 2042.04-400-002-00 NON PROPRIETARY VERSION Page 18 of 61

l NON PROPRIETARY VERSION the tubes due to internal and external pressure. The pressure inside (tube side pressure) and outside (shell side pressure) the tubes cause the tubes to change length axially due to the Poisson ratio effect.

Constraint equations are used at the tube-tucesheet interfaces to provide

" built in" conditions for the tubes. This is accomplished by forcing the rotation of each beam element at the tube tubesheet interface nodes to be equal to the relative axial displacement of the two adjacent tubesheet nodes divided by the distance between the adjacent nodes (i.e., the beam element is forced to remain perpendicular to the tubesheet surface as the tubesheet rotates.)

The main closure studs joining the channel and the channel cover of the head are represented in the model by beam elements with "per radian" properties.

Calculations for these properties are given in Appendix A.4 .

Stud beam elements extend beyond the flange surface a distance equal tn one-half of the stud diameter. This additional length accounts for the flexibility of the stud and fastener and is calculated

. Rigid beam elements are used to connect the ends of the stud to the flanges via the node coupling scheme shown in Figure 5.2 and 5.3. By coupling the studs to the flanges in this manner, axial, radial and rotational displacements of the flanges are imposed upon the stud.

At the flange seating surface, compressive loads must exist during operation to prevent leaks. Therefore, in order to have a mathematically linear modei and to simulate the gasket's performance during operational loads, two nodes at the gasket seating surface are coupled together axially (see Figure 5.4). This will be the only connection between the channel and channel cover besides the beams used to represent the studs. Stud preload can then be adjusted to ensure that tensile loads do not develop at the seating surface. This coupling method also allows the impact of the closure joint slipping from bolt preload to be 2042.04-400-002-00 NON PROPRIETARY VERSION Page 19 of 61

NON PROPRliTARY VERSION included since relative radial motion between the components is not restrained at the seating surface.

5.2 Base Case loads The tubesheet finite element model described in Section 5.1 is analyzed for various base case loadings. Stress and displacement results from these base cases are then combined to provide results for the required steady-state operation and shutdown conditions. The base case loads and a description of each are listed below, in all cases, the displacement boundary conditions specified are no axial motion at the axial symmetry plane of the shell barrel and tubes, no rotation at the axial symmetry plane of the tubes and no radial motion of the tubes at the baffle plate locations. Figure 5.6 is a node plot showing the boundary constraints described above.

. Base Case 1: Preload on the Main Closure Studs An initial strain of 0.0011 in./in, is applied to the beam elements representing the studs which join the channel cover to the channel flange to simulate a preload of 25,100 lb/ bolt,

, in the beam elements. No other loads are applied.

Base Case 2: Tube Side Pressure = 185 psi A pressure of 185 psi is applied to the inner surface of the head and to the non-perforated tube side surface of the tubesheet. For the perforated region the equivalent tube side pressure is 123.23 psi

. Included in the model is the initial strain on the tube beam elements to simulate growth due to internal tube pressure and a uniform tubesheet temperature of 185'F. The uniform temperature coupled with the modified value of ALPX simulates the effect of pressure in the perforations .

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i l

l NON PROPRIE~ARY VERSION Base Case 3: Shell Side Pressure 155 psi A pressure of 155 psi is applied to the non-perforated, shell side surface of the tubesheet and to the inner surface of the shell. For the perforated region, the equivalent shell side pressure is 86.52 psi

. The uniform tubesheet temperature is I

not recuired for this load case since there is no pressure in the penetrations. The initial strain to simulate tube growth due to external pressure is included .

Base Case 4: Steady-State Operating Temperatures Temperatures are applied to the axisymmetric model as shown in Figures 5.14 and 5.15. The temperatures of 95'F on the tube side and 165'F on the shell side are conservatively chosen because of this being the maximum temperature extremes that the LPCI-HX is subjected to

. The tube temperature,122'F, is calculated .

These temperatures are applied to the model to simulate thermal strains caused by expansion due to the temperature differences. The reference temperature is set to 70*F for the thermal strain calculations.

Thermal strains are a function of the product of the coefficient of f thermal expansion and the difference between the actual temperature and the reference temperature. No other loads are applied.

l Base Case 5: No Flow Thermal Conditions (Shutdown)

To simulate shutdown thermal conditions, a uniform temperature of 51'F (the assumed worse case temperature, ,

that the entire vessel is subjected to) is applied throughout the vessel. The reference temperature is se', to 70'F for thermal strain calculations. This temperature allows for any thermal contractions caused by a temperature drop below the refere: ice temperature. No other loads are applied.

2 2042.04-400-002 00 NON PROPRIETARY VERSION Page 21 of 61

l l

NON PROPRIETARY VERSION 5.3 Operational Loads Two operational load cases are analyzed by combining '.>ase case loads that are applicable during specific operating conditions. The base case loads are combined by using a computer program .

Stresses and displacements from the individual base cases are mathematically combined in order to obtain the desired resulting operational load case.

The operational load cases and a description of each are listed below.

31gady State Operation

, Steady-state operation is a combination of the base case loads:

preload, tube side pressure, shell side pressure and steady-state operating temperatures. This case simulates the LPCI-HX being ;t full steady-state operating conditions.

No Flow Condition (Shutdown)

Shutdown is a combination of the base case loads: preload and no flow thermal conditions. This case simulates the LPCI-HX being shutdown with the temperature of the vessel dropping to 51'F.

5.4 Axisymmetric Analysis Results The maximum stress due to steady-state operation is 15,030 psi. This value, which is located in the tubesheet, is calculated

. Displacement plots for the two operational loads (steady-state and shutdown) are shown

. Radial and hoop stress contour plots are shown

, respectively.

Displacement conditions from the 2-D model are next applied to the 3-D model at common nodal locations in the two models, herein referred to as the cut boundaries, to simulate these conditions at the modeled discontinuities described in Section 6.0 of this report.

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I c ..'Ff Iu. ilW "

2 FEB 20 1990 o 08:02:52

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FIGllRE 5.2 TUDESitEET fill!IE ELEMEliT MODEL GE0 METRY

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EAT EXCHANGER - 2D 'AXISYMMETRIC TIGURE 5.4a TUBESilIET FIN 11E ELEMENT H0 DEL GEOMETRY ,

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g- ,.

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. . _ _ _ _ . _ - _ . . - _ - - - ~ , . . _ _ - - ~ . .

T

[REi < -liiF-

'o T FEB 13 1990

? 1 14:03: 40 PREP 7 ELEf1EleTS 6 TYPE iRJ1 2 T = 95*f e .

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C liEAT EXCilAtr.ER TIGURE 5 14 STEADY-STATE OPERATING TEt4PERATURES _

-_ _ _ _ _1 l

y [ REP 7 -liTP=

"g FEB 13 1990

? 14808:29 8 PREP 7 ELEr.ErITS TYPE tRJr1 o _

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l NON PROPRIETARY VERSION 6.0 3-01 MENS 10NAL ELEMENT ANALYSIS 6.1 Development of 3-D Finite Element Model A three-dimensional finite element model of the region of interest is shown in Figures 6.1 to 6.5. This model uses isoparametric solid elements to represent the quarter symmetry region. Cut boundary nodes on the 3 D model have the displacements from the 2 0 model applied at their common locations (cut boundary locations). See Figure 6.3.

l .

1 General material properties for the components included in the 3 D finite element model are given in Section 3.2 of this report. Material properties utilized for the 3 0 finite element analysis are summarized in Appendix B.1.

6.2 [Lise Case loads The 3 D finite element model is analyzed for the same base case loads that the 2 D finite element model is analyzed for. For the shell side pressure case, internal pressure is applied to the inside of the nozzle as well as the shell. A blev:of f load is applied to the nozzle safe end as an outward acting pressure with a magnitude of 629.2 psi .

4 For the tube side pressure case, pressure is applied only to the tubesheet rim on the tube side, The steady-state thermal case has the same temperatures applied as the 2 0 model except fcr the addition of the support bracket temperature of 70*F (acting as a cooling fin). The thermal gradients can be seen in Figure 6.6.

For all cases, displacement boundary conditions include symmetry constraints at x = 0 and z = 0 planes by fixing all nodes on these two planes in the e-direction (see Figure 6.4). Also, bolt attachments are simulated at their corresponding locations in the vertical support bracket by fixing nodes at 2042.04-400 002 00 NON PROPRIETARY VERSION Page 49 of 61

l NON PROPRIETARY VERSION

)

these locations in the x and z directions (see Figure 6.5). Displacements from the 2-D finite element model for each base load case are then applied to the 3 D finite element model at the cut boundary loc $.tions (see Figure 6.3).

6.3 ODerational loads The operational loads analyzed in the 3 D finite element analysis are the same as those analyzed in the 2 D finite element analysis. The steady-state operating load includes tube side pressure, shell side pressure and steady-state operating temperatures along with steady-state operation displacements

from the 2 D model. The preload effects are accounted for in the displacements from the 2-D model.

The no flow shutdown load includes the assumed worse case shutdown temperature of 51*F. The preload effects are again accounted for in the i displacements from the 2 D model.

6.4 F D Analysis Results The maximum principal stress due to steady state operation is 16,711 psi.

This value is approximately the same at two places in the vessel. They are the nozzle to shell intersection and the bracket to shell intersection

. This stress is the maximum stress

$ the shell side is subjected to.

The principal stresses due to shutdown condition:;, at the same locations of the maximum stresses due steady state operation. are approximately 0 psi.

This stress is used as the minimum stress the shell side is subjected to

. Therefore, cycling occurs between the maximum and minimum stresses for the fracture mechanics analysis.

2042.04-400-002-00 NON PROPRIETARY VERSION Page 50 of 61

- ~

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S:

@ pair- - I m= -

2 FEB 28 1990

1. Monel Cladding 11s28 48 g PREP 7 ELEl1EtiTS 8 Tutiesheet P!AT F4Ur1 7 Rim N DA- -

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l NON PROPRIETARY VERSION 7.0 FRACTURE MECHANICS ANALYSIS The stress results from the 3 D finite element analysis are used for the fracture mechanics because they are higher stresses than those from the 2 D finite element analysis. The cyclic stress range from the 3 D analysis is conservatively chosen to be from 0 psi (shutdown conditions) to 16,700 psi (steady-state operation).

The material toughness, K!C, of 76 ksi /Tii. used to calculate the critical crack size at which failure would occur, Acr, was obtained from an

+ approximate K IC vs. temperature plot using dynamic A 2128 data taken from

. Because only two data points are given, it is reasonable to assume these two points define the upper and lower shelf values of the ductile brittle transition region

. To r5tain temperature adjusted data, the two data points were shif ted to the right by 17'F. This moves the upper shelf datum point at 60'F to the lowest allowable service temperature, 77'F,

. The straight line connecting the two adjusted transition boundary points was used to determine the conservative estimate of KIC - 76 ksi /TE at 51'F, the minimum expected vessel temperature.

The two data points could also have been shifted to the right by 37'F to bring the NDT, defined below, to the lowest allowable service temperature. This was not done because 'Se low cyclic strain rates between shutdovin and steady state operation of LPCI-hX are consistent with the use of K lC fracture toughness values. tstimating K IC .t ting dynamic K data embodies substantial conservatism since the temperature shift ID between KIC sad KID is approximately 158'F, via the Barsom shift

, . This shift is calculated based upon the yield stress of 38 ksi for A-2128.

2042.04-400-002 00 NON PROPRIETARY VERSION Page 57 of 61

1 NON PROPRIETARY VERSION To relate the material toughness to the nil ductility temperature (NDT) of 40*F ,K IC was calculated

. The straight line between this value, 57 ksi /T5" and the adjusted upper bound of the transition region, 98 ksi /Tir. could also be used to conservatively avaiuate K at 51*F as indicated . It was judged that IC this K IC value of 69 ksi Jin. is unnecessarily conservative, so the 76 ksi

/iii. was used in the fracture mechanics evaluation. However, it is apparent from the extremely long calculated fatigue life discussed below that the 69 ksi /In, material toughness would also have been iguate.

For the calculated maximum stress from the 3-0 analyses,16.7 ksi, the critical crack size at which failure would occur is 5.25 inches

. Using the maximum stress range (0.0 - 16.7 ksi) and the crack growth rate equation , the number of cycles, N, required for various assumed initial crack sizes, Aj , to grow to the critical size, A p, is calculated .

The results, based u. son an initial crack size of .010 inch (chosen as the maximum crack site that could go undetected by inspection of the unit at the time of its construction) and a cold HX shutdown temperature of 51*F, show that the number of cycles that would be required to propagate a crack of this size to the critical crack size is approximately 1.5 x 105 Since the critical crack size. is greater than the wall thickness an additional calcul; tion was performed which showed that the number of cycles that wor be requ! red to propagate the initial crack through the wall is approximately 1.4 x 108, 2042.04 400 002-00 NON PROPPIETARY VERSION Page 58 of 61

NON PROPRIETARY VERSION

. From ti. e results it can be seen that the crack should begin to leak approximately 96,000 cycles before it would fail. '

Based solely from this fracture mechanics analysis and a system startup and shutdown cycle occurring on an average of once a month, the acceptable number of cycles is far beyond any anticipated service life of the LPCI HX as well as several orders of magnitude beyond the maximum conceivable piant-lim :ae including life extension. Therefore, it is concluded that the frat.ure toughness cf the shell side of the LPCI HX is more than adequate, I

2042.04-400-002-00 NON PROPRIETARY VERSION Page 59 of 61

NON PROPRIETARY VERION

8.0 REFERENCES

1 I

2042.04-400-002-00 NON PROPRIETARY VERSION Page 60 of 61

-l

\

NON PROPRIETARY VER10N r

8.0 REFERENCES

(Con't) i l

f l

l I

l 2042.04-400 002-00 NON PROPRIETARY VERSION Page 61 of 61

A -

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-8PPENDIX A AXISYMMETRIC TVBESHEET FINITE ELEMENT MODEL APPENDIX 1

1 1

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l 1

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Accendix A.1 f Finite Element Model Material Properties

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- - - _ - ~ - . .

l l

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1 p i Accendix A,2 Adjusted Material Properties for Stud Hole Regions

_ , - . g.w-,,, ,ey e

-. . . . -- . _ ... . . . - - - - - . - . - . . - . . . . ~ . .-- - ..

ADoendix A,3 1

h Tubesheet and Tube Data 2

l t

l l

I l

l f

t

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Anoendix A 4 Stud Bolt Properties

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Accendix A,5 Calculation of Closure Stud Preload i

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Accendix A.6 Calculation of Tube Temperature L

i n

. Apoendix A.7 Axisymmetric Analysis Results 4

m APPENDIX B 3-0 FINITE ELEMENT MODEL APPENDIX I

s l.

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. . . . . . _ . .~ ..- - _ _ - . . _ _ . . . . _ - . . .. - . - . ..

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i j Accendix B.2 Calculation of Blowoff Pressure i

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Accendix B.3 3-0 Analysis Results

APPENDIX C FRACTURE MECHANICS ANALYSIS APPENDIX

Attachment A Fracture Toughness Evaluation of the Shell Side of the LPCI Heat Exchanger for Dresden Station Unit 2 t

(Propietary Report - Report No. 2042,04-400-001-00)

ENLD701/14

ML20067B492

Text

SUMMARY

1.0 INTRODUCTION

SUMMARY

.= = =

8.0 REFERENCES

8.0 REFERENCES

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