ML20210D304

Rev 2 to Tubesheet Region Plugging Criterion for Alabama Power Co Farley Nuclear Station Unit 2 Steam Generators

ML20210D304
Person / Time
Site:	Farley
Issue date:	04/30/1987
From:	WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19292H207	List: ML20210D277 ML20210D297 ML20210D304
References
TAC-65287, WCAP-11314, WCAP-11314-R02, WCAP-11314-R2, NUDOCS 8705070115
Download: ML20210D304 (40)

Text

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WESTINGHOUSE CLASS 3 si n TUBESHEET REGION PLUGGING CRITERION FOR THE ALABAMA POWER COMPANY FARLEY NUCLEAR STATION UNIT 2 STEAM GENERATORS September 1986 Revised April 1987 WESTINGHOUSE ELECTRIC CORPORATION SERVICE TECHNOLOGY DIVISION P. O. BOX 3377 PITTSBURGH, PENNSYLVANIA 15230 Work Performed Under Charge No. BH60 - 44307 (1986) 8705070115 870504 PDR ADOCK 05000364 P PDR 430sM;a9/ 4:057

ABSTRACT An evaluation was performed to develop a plugging criterion, known as the F*

criterion, for determining whether or not repairing or plugging of full depth hardroll expanded steam generator tubes is necessary for degradation that has been detected in the region of the tube located within the tubesheet. The evaluation consisted of analysis and testing programs aimed at quantifying the residual radial preload of Westinghouse Model 51 steam generator tubes hardrolled into the tubesheet. Analysis was performed to determine the length of hardroll engagement required to resist tube pullout forces during normal and faulted plant operation. It was postulated that the radial preload would be sufficient to significantly restrict leakage during normal and operating conditions. On this basis an F* criterion value of 1.54 inches was established as sufficient for continued plant operation regardless of the

extent of tube degradation below F*. The evaluation also demonstrates that application of the F* criterion for tube degradation within the tubesheet affords a level of plant protection commensurate with that provided by RG 1.121 for degradation located outside of the tubesheet region. A conservative eddy current elevation measurement uncertainty value is evaluated and included in an operational F* distance.

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CONTENTS

, TOPIC PAGE

1.0 INTRODUCTION

5 2.0 EVALUATION 7 2.1 Determination of Elastic Preload Between the 8 Tube and the Tubesheet 2.1.1 Radial Preload Test Configuration Description 9 2.1.2 Preload Test Results Discussion and Analysis 9 2.1.3 Residual Radial Preload During Plant Operation 12 2.2 Engagement Distance Determination 14 2.2.1 Applied Loads 15 2.2.2 End Effects 15 2.2.3 Calculation of Engagement Distance Required, F* 16 2.3 Limitation of Primary to Secondary Leakage 18 2.3.1 Operating Condition Leak Considerations 18 2.3.2 Postulated Accident Condition Leak 20 Considerations 2.3.3 Operating Plant Leakage Experience for Within 21 Tubesheet Tube Cracks 2.4 Tube Integrity Under Postulated Limiting Conditions 21

2.5 Chemistry Considerations 21 2.5.1 Tubesheet Corrosion Testing 22 2.5.2 Tubesheet Corrosion Discussion 23 2.6 Determination of Eddy Current Elevation 24 Measurement Uncertainty 3.0

SUMMARY

25

4.0 REFERENCES

27 1

1 4303M:49/042 i' 3

TABLES TOPIC PAGE Table 1. Model 51 Preload Measurement Tests - TEST DATA 28 Table 2. Model 51 Preload Measurement Tests - STRESS 31 RESULTS Table 3. Model 51 Tube Roll - PRELOAD ANALYSIS

SUMMARY

32 APPENDICES Appendix I - Disposition of Tubes with Indications 33 Above F' Appendix II - Tube Wall Temperatures of Plugged Tubes 37 430SM:49/042437  :

TUBESHEET REGION PLUGGING CRITERION FOR FULL DEPTH HARDROLL EXPANDED TUBES

1.0 INTRODUCTION

The purpose of this report is to document the development of a criterion to be used in determining whether or not repairing or plugging of full depth hardroll expanded steam generator tubes is necessary for degradation which has been detected in that portion of the tube which is within the tubesheet.

Existing Alabama Power Company, Farley Nuclear Station, Unit 2 Technical

, Specification tube repairing / plugging criteria apply throughout the tube length, but do not take into account the reinforcing effect of the tubesheet on the external surface of the tube. The presence of the tubesheet will constrain the tube and will complement its integrity in that region by essentially precluding tube deformation beyond its expanded outside diameter.

The resistance to both tube rupture and tube collapse is significantly strengthened by the tubesheet. In addition, the proximity of the tubesheet significantly affects the leak behavior of through wall tube cracks in this region, i.e., no significant leakage relative to plant technical specification allowables is to be expected. Based on these considerations, the use of an alternate criterion for establishing plugging margin is justified.

This evaluation forms the basis for the development of a criterion for obviating the need to repair a tube (by sleeving) or to remove a tube from service (by plugging) due to detection of Indications, e.g., by eddy current

, testing (ECT), in a region extending over most of the length of tubing within the tubesheet. This evaluation applies to the Farley Unit 2 Westinghouse Model 51 steam generators and assesses the integrity of the tube bundle, for tube ECT indications occurring on the length of tubing within the tubesheet, relative to:

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1) Maintenance of tube integrity for all loadings associated with normal plant conditions, including startup, operation in power range, hot standby and cooldown, as well as all anticipated transients.

2) Maintenance of tube integrity under postulated limiting conditions of primary to secondary and secondary to primary differential pressure, e.g., feedline break (FLB).

3) Limitation of primary to secondary leakage consistent with accident analysis assumptions.

The result of the evaluation is the identification of a distance, designated F* (and identified as the F* criterion), below the bottom of the roll transition, or the top of the tubesheet, whichever is lower in elevation, for which tube degradation of any extent does not necessitate remedial action, e.g., plugging or sleeving. This portion of the tube is referred to as the F*

region. The F* criterion provides for sufficient engagement of the tube to tubesheet hardroll such that pullout forces that could be developed during normal or accident operating conditions would be successfully resisted by the elastic preload between the tube and tubesheet. The necessary engagement length applicable to the Farley Unit 2 steam generators was found to be 1.54 inches based on preload analysis. Added to the the necessary engagement length to specify the operational value of F* is an allowance for eddy current elevation measurement uncertainty. The operational value of F* 15 the value to be included in the Technical Specifications. A conservative eddy current uncertainty value of 0.25 inch is included in the operational F* distance of 1.79 inches for J. M. Farley Nuclear Plant Unit 2, see Section 2.5.

! Application of the F* criterion provides a level of protection for tube degradation in the tubesheet region commensurate with that afforded by Regulatory Guide (RG) 1.121, reference 1, for degradation located outside tne i

tubesheet region.

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2.0 EVALUATION Tube rupture in the conventional sense, i.e., characterized by an axially oriented " fishmouth" opening in the side of the tube, is not possible within the tubesheet. The reason for this is that the tubesheet material prevents the wall of the tube from expanding outward in response to the internally acting pressure forces. The forces which would normally act to cause crack extension are transmitted into the walls of the tubesheet, the same as for a nondegraded tube, instead of acting on the tube material. Thus, axially oriented linear indications, e.g., cracks, cannot lead to tube failure within the tubesheet and may be considered on the basis of leakage effects only.

Likewise, a circumferentially oriented tube rupture is resisted because the tube is not free to deform in bending within the tubesheet. When degradation has occurred such that the remaining tube cross sectional area does not present a uniform resistance to axial loading, bending stresses are develooed which may significantly accelerate failure. When bending forces are resisted by lateral support loads, provided by the tubesheet, the acceleration mechanism is mitigated and a tube separation mode similar to that which would occur in a simple tensile results. Such a separation mode, however, requires the application of significantly higher loads tnan for the unsupported case.

In order to evaluate the applicability of any developed criterion for indications within the tubesheet some postulated type of degradation must necessarily be considered. For this evaluation it was postulated that a circumferential severance of a tube could occur, contrary to existing plant operating experience. However, implicit in assuming a circumferential severance to occur, is the consideration that degradation of any extent could be demonstrated to be tolerable below the location determined acceptable for the postulated condition.

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When the tubes have been hardrolled into the tubesheet, any axial loads developed by pressure and/or mechanical forces acting on the tubes are resisted by frictional forces developed by the elastic preload that exists

! between the tube and the tubesheet. For some specific length of engagement of l the hardroll, no significant axial forces will be transmitted further along 43G34:49/G4Ms7 l

the tube, and that length of tubing, i.e., F*, will be sufficient to anchor the tube in the tubesheet. In order to determine the value of F* for application in Model 51 steam generators a testing program was conducted to measure the elastic preload of the tubes in the tubesheet.

The presence of the elastic preload also presents a significant resistance to flow of primary to secondary or secondary to primary water for degradation which has progressed fully through the thickness of the tube. In effect, no leakage would be expected if a sufficient length of hardroll is present. This has been demonstrated in high pressure fossil boilers where hardrolling of tube-to-tubesheet joints is the only mechanism resisting flow, and in steam generator sleeve to tube joints made by the Westinghouse hybrid expansion joint (HEJ) process.

2.1 DETERMINATION OF ELASTIC PRELOAD BETWEEN THE TUBE AND TUBESHEET Tubes are installed in the steam generator tubesheet by a hardrolling process which expands the tube to bring the outside surface into intimate contact with the tubesheet hole. The roll process and roll torque are specified to result in a metal to metal interference fit between the tube and the tubesheet.

A test program was conducted by Westinghouse to quantify the degree of interference fit between the tube and the tubesheet provided by the full depth hardrolling, also referred to as roll expansion (RE), operation. The data generated in these tests has been analyzed to determine the length of hardroll required to preclude axial tube forces from being transmitted further along the tube, i.e., to establish the F* criterion. The amount of interference was determined by installing tube specimens in collars specifically designed to simulate the tubesheet radial stiffness. A hardroll process representative of that used during steam generator manufacture was used in order to obtain specimens which would exhibit installed preload characteristics like the tubes in the tubesheet.

Once the hardrolling was completed, the test collars were removed from the tube specimens and the springback of the tube was measured. The amount of springback was used in an analysis to determine the magnitude of the interference fit, which is, representative of the residual tube to tubesheet radial load in Westinghouse Model 51 steam generators.

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2.1.1 RADIAL PRELOAD TEST CONFIGURATION DESCRlPTTON The test program was designed to simulate the interface of a tube-to-tubesheet full depth hardroll for a model 51 steam generator. The test configuration consisted of six cylindrical collars, approximately [ l a,c.e inches in length, [ l a,c.e inches in outside diameter, and [ l a,c.e inch in inside diameter. The RE process was designed to provide the same preload regardless of hole diameter up to a hole diameter of approximately 0.935 in.

The preload was determined by tube thinning which, in turn, was determined by roll expander motor stalling torque. A single nominal stalling torque value was used for all tube-to-tubesheet joints. A mill annealed, Inconel 600 (ASME 58-163), tubing specimen, approximately [ 3a ,c.e inches long with a nominal

[ 3a ,c.e outside diameter before rolling, was hard rolled into each collar using a process which simulated actual tube installation conditions.

The design of the collars was based on the results of performing finite element analysis of a section of the steam generator tubesheet to determine radial stiffness and flexibility. The inside diameter of the collar was chosen to match the size of holes drilled in the tubesheet. The outside I diameter was selected to provide the same radial stiffness as the tubesheet.

The collars were fabricated from AISI 1018 carbon steel similar in mechanical properties to the actual tubesheet material. The collar assembly was clamped in a vise during the rolling process and for the post roll measurements of the tube ID. Following the taking of all post roll measurements, the collars were saw cut to within a small distance from the tube wall. The collars were then split for removal from the tube and tube ID and OD measurements were repeated.

Two end boundary conditions were imposed on the tube specimen during rolling.

The end was restrained from axial motion in order to perform a tack roll at l

the bottom end, and was allowed to expand freely during the final roll.

2.1.2 PRELOAD TEST RESULTS DISCUSSION AND ANALYSIS All measurements taken during the test program are tabulated in Table 1. The data recorded was employed to determine the interfacial conditions of the

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tubes and collars. These consisted of the ID and OD of the tubes prior to and

, after rolling and removal from the collars as well as the inside and outside i dimensions of each collar before and after tube rolling. Two orthogonal measurements were taken at six axial locations within the collars and tubes.

All measured dimensions given in Table 1 are in inch units. The remainder of the data of particular interest was calculated from these specific .

dimensions. The calculated dimensions included wall thickness, change in wall thickness for both rolling and removal of the tubes from the collars, and percent of spring back.

! Using the measured and calculated physical dimensions, an analysis of the tube deflections was performed to determine the amount of preload radial stress present following the hardrolling. The analysis consisted of application of

conventional thick tube equations to account for variation of structural i parameters through the wall thickness. However, traditional application of

cylinder analysis considers the tube to be in a state of plane stress. For these tests the results implied that the tubes were in a state of plane strain elastically. This is in agreement with historical findings that theoretical values for radial residual preload are below those actually measured, and that axial frictional stress between the tube and the tubesheet increases the

} residual pressure. In a plane stress analysis such stress is taken to be zero, references 2 and 3. Based on this information the classical equations relating tube deformation and stress to applied pressure were modified to reflect plane strain assumptions.

The standard analysis of thick walled cylinders results in an equation for the radial deflection of the tube as:

u - C;

• r + C2 /r (1) ,

where, u = radial deflection r - radial position within tne tube wall, i

and the constants, Cj and C2are found from the boundary conditions to be functions of the elastic modulus of the material, Poisson's ratio for the material, the inside and outside radli, and the applied internal and external 430aM:49/C4:437 ';

.,,__,,-r--- _ - -._. _ , , , _ . _ - - - - _ - . _ - _ _ _ _ - - - - _ - - _ , - - . - _ . . - - - - - - - . -_--

pressures. The difference between an analysis assuming plane stress and one assuming plane strain is manifested only in a change in the constant C2 '

For materials having a Poisson's ratio of 0.3, the following relation holds:

C2 (Plane Strain) = 0.862

• C 2(Plane Stress) (2)

The effect on the calculated residual pressure is that plane strain results are higher than plane stress results by slightly less than 10 percent.

Comparing this effect with the results reported in reference 2 indicated that better agreement with test values is achieved. It is to be noted that the

residual radial pressure at the tube to tubesheet interface is the compressive radial stress at the OD of the tube.

By substituting the expressions for the constants into equation (1) the deflection at any radial location within the tube wall as a function of the f

internal and external pressure (radial stress at the ID and 00) is found. This expression was differentiated to obtain flexibility values for the tube deflection at the ID ard 00 respectively, e.g., dU1/dPo is the ratio of the radial deflection at the ID due to an 00 pressure. Thus, dUi/dPo was used to find the interface pressure and radial stress between the tube and the tubesheet as:

-I S --P g = - (ID Radial Springback) / (dUl/dPo) (3) ro The calculated radial residual stress for each specimen at each location is tabulated in Table 2. The mean residual radial stress and the standard deviation was found to be [ l a.c.e psi and [ l a,c.e psi respectively.

In order to determine a value to be used in the analysis, a tolerance factor for [ l a.c.e percent confidence to contain [ l a,c.e percent of the population was calculated, considering the [ Ja ,c.e useable data points, to be [ l a,c.e . Thus, a [ l a,c.e lower tolerance limit (LTL) for the radial residual preload at room temperature is [ l a,c e 39, j

l i

t-1 4308M:49/042487

,------.--..-r.--- - - - , - . - . - . - - . - - - . - - - - - . - , - - - - - - , - , - , - - - , - - - - - - - . - . , - , . , . - , . - - , -,e.. . - - - - - - - - - - - . , - - - , . . . - . .

2.1.3 RESIDUAL RADIAL PRELOAD CURING PLANT OPERATION During plant operation the amount of preload will change depending on the pressure and temperature conditions experienced by the tube. The room temperature preload stresses, i.e., radial, circumferential and axial, are such that the material is nearly in the yield state if a comparison is made to ASME Code, reference 4, minimum material properties. Since the coefficient of \

thermal expansion of the tube is greater than that of the tubesheet, heatup of the plant will result in an increase in the preload and could result in some yielding of the tube. In addition, the yield strength of the tube material decreases with temperature. Both of these effects may result in the preload being reduced upon return to ambient temperature conditions, i.e., in the cold condition. However, as documented in reference 5 for a similar investigation tube pull out test preceded by a very high thermal relaxation soak showed the analysis to be conservative.

Tne plant operating pressure influences the preload directly based on the application of the pressure load to the ID of the tube, thus increasing the amount of interface loading. The pressure also acts indirectly to decrease the amount of interface loading by causing the tubesheet to bow upward. This bow results in a dilatation of the tubesheet holes, thus, reducing the amount of tube to tubesheet preload. Each of these effects may be quantitatively treated.

The maximum amount of tubesheet bcw loss of preload will occur at the too of the tubesheet. Since F* is measured from the bottom of the hardroll transition (BRT) or the top of the tubesheet, whichever is lower in elevaticn, and leakage is to be restricted by the F* region of the tube, the potential for the tube section within the F* region to experience a net loosening during operation is considered for evaluation. The effects of the three identified mechanisms affecting the preload are considered as follows:

1. Thermal Expansion Tightening - The mean coefficient of thermal expansion for the Inconel tubing between ambient conditions and 543*F is 7.79*10 in/in/*F. That for tne steam generator tubesheet is 12:5M:4 M A:037  :

7.16*10-6 in/in/*F. Thus, there is a net difference of 0.63*10-6 in/in/*F in the expansion property of the two materials. Considering a temperature difference of approximately 483*F between ambient and cold leg operating conditions the increase in preload between the tube (t) and the tubesheet (ts) was calculated as:

Sri - (0.63E-6)*(483)*(Collar ID) /2/ ((dUi/dP1)ts- (duo /dPo)t (4)

This calculation was also performed and tabulated in Table 2. The results indicate that the increase in preload radial stress due to thermal expansion is [ ]"'C psi. It is to be noted that this 4

value applies for both normal operating and faulted conditions.

2. Internal Pressure Tightening - The maximum normal operating differential pressure from the primary to secondary side of the steam generator is [ l a.c.e psi during a loss of load transient. The internal pressure acting on the wall of the tube will result in an i

increase of the radial preload on the order of the pressure value.

The increase was found as:

SrP = Po - - Pg (duo /dPI) /( (dVi/dP1)ts - (duo /dPo)t' (5)

In actuality, the increase in preload will be more dependent on the I internal pressure of the tube since water at secondary side pressure would not be expected between the tube and the tubesheet. This i

differential pressure, along with the associated temperature and tubesheet bowing effects, was the most stringent normal condition in terms of determining the F* length. Increased differential pressure not only increases radial preload directly, it also increases l tubesheet bending which reduces preload.

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Results from the performance of this calculation are tabulated in Table 2 for normal operating conditions and summarized in Table 3, j the summary sheet for both normal and faulted conditions. The results indicate that the increase in preload radial stress is [ l a,c,e psi for normal operating conditions and [ l a,c.e psi for faulted (FLB) operating conditions.

3. Tubesheet Bow Loosening - An analysis of the Model 51 tubesheet was performed to evaluate the loss of preload stress that would occur as a result of tubesheet bow. The analysis was based on performing finite element analysis of the tubesheet and SG shell using equivalent perforated plate properties for the tubesheet, reference 3. Boundary conditions from the results were then applied to a smaller, but more detailed model, in order to obtain results for the tubesheet holes. Basically the deflection of the tubesheet was used to find the stresses active on the top surface and then the presence of the holes was accounted for. For the location where the loss of preload is a maximum, the radial preload stress would be reduced by [ l a.c.e psi during normal operation and

[ l a,c.e psi during faulted (FLB) operating conditions. During LOCA the differential operating pressure is from secondary to primary. Thus, the radial preload will increase by [ l a,c.e p3g as the tubesheet bows downward.

Combining the room temperature hardroll preload with the thermal and pressure effects results in a net operating preload of [ l a,c.e psi during normal operation and [ l a,c.e psi for faulted operation. In addition to restraining the tube in the tubesheet, this preload should effectively retard leakage from indications in the tubesheet region of the tubes.

2.2 ENGAGEMENT DISTANCE DETERMINATION The calculation of the value of F* recommended for application to the Farley 2 steam generators is based on determining the length of hardroll necessary to equilibrate the applied loads during the maximum normal ocerating conditions 4308M:49/042257

3 or faulted conditions, whichever provides the largest value. Thus, the appliad loads are equilibrated to the load carrying ability of the hardrolled tube for both of the above conditions. In performing the analysis, ,

consideration is made of the potential for the ends of the hardroll at the hardroll transition and the assumed severed condition to have a reduced load carrying capability.

2.2.1 APPLIED LOADS The applied loads to the tubes which could result in pullout from the tubesheet during all normal and postulated accident conditions are predominantly axial and due to the internal to external pressure differences.

For a tube which has not been degraded, the axial pressure load is given by the product of the pressure with the internal cross-sectional area. However,

{oratubewithinternaldegradation,e.g.,cracksorientedatanangletothe axis of the tube, the internal pressure may also act on the flanks of the l degradation. Thus, for a tube which is conservatively postulated to be severed at some location within the tubesheet, the total force acting to remove the tube from the tubesheet is given by the product of the pressure and the cross-sectional area of the tubesheet hole. The force resulting from the pressure and internal area acts to pull the tube from the tubesheet and the force acting on the end of the tube tends to push the tube from the tubesheet. For this analysis, the tubesheet hole diameter has been used to determine the magnitude of the pressure forces acting on the tube. The forces I

acting to remove the tube from the tubesheet are [ l a.c.e pounds and

[ ]#'C pounds respectively for normal and faulted operating conditions. Any other forces such as fluid drag forces in the U-bends and vertical seismic forces are negligible by comparison.

/- 2.2.2 END EFFECTS I

l The analysis for the radial preload pressure between the tube and the tubesheet made no consideration of the effect of the material discontinuity at the hardroll transition to the unexpanded length of tubing. In addition, for a tube which is postulated to be severed within the tubesheet there is a

material discontinuity at the location where the tube is severed. For a small l

l 4308M:a3/C42a37 '5

distance from each discontinuity the stiffness, and hence the radial preload, of the tube is reduced relative to that remote from the ends. The analysis of end effects in thin cylinders is based on the analysis of a beam on an elastic foundation. For a tube with a given radial deflection at the end, the deflection of points away from the end relative to the end deflection is given by:

u rx /u ro -e

• cosine ( A

• x ) W where, A = [ ]a,c.e End effect constant x - Distance from the end of the tube.

For the radially preloaded tube, the distance for the end effects to become negligible is the location where the cosine term becomes zero. Thus, for the roll expanded Model 51 tubes the distance corresponds to the product of "A" times "x" being equal to (pi/2) or [ l a,c,e inch.

The above equation can be integrated to find the average deflection over the affected length to ue 0.384 of the end deflection. This means that on the average the stiffness of the material over the affected length is 0.616 of the stiffness of the material remote from the ends. Therefore, the effective preload for the affected end lengths is 61.6 percent of the preload at regions more than [ l a,c.e inch from the ends. For example, for the normal operating net preload of [ l a,c.e psi or [ ]a,c.e pounds per inch of length, the effective prelcad for a distance of [ Ja .c.e inch from the end is [ ]d'C pounds per inch or [ 3a ,c.e ounds.

2.2.3 CALCULATION OF ENGAGEMENT DISTANCE REQUIRED, F*

The calculation of the required engagement distance is based on determining the length for preload frictional forces to equilibrate the applied operating loads. The axial friction force was found as the product of the radial preload force and the coefficient of friction between the tube and the tubesheet. The value assumed for the coefficient of friction was a

[ 3 ,c.e , reference 5. For normal operation the radial preload is

[ l .c.e psi or [

a a 3 ,c.e pounds per inch of engagement. Thus, the 430BM:a^d:al:37 '6

axial friction resistance force is [ l a,c.e ounds per inch of engagement. It is to be noted that this value applies away from the ends of the tube. For any given engagement length, the total axial resistance is the sum of that provided by the two ends plus that provided by the length minus the two end lengths. From the preceding section the axial resistance of each end is [ ]a,c e pounds. Considering both ends of the presumed severed tube, i .e., the hardroll transition is considered one end, the axial a

resistance is [ l c.e pounds plus the resistance of the material between the ends, i.e., the total length of engagement minus [ ]d' inch. For example, a one inch length has an axial resistance of, g ya .c.e Conversely, for the maximum normal pressure applied load of [ 3d '

pounds, considered as [ l a.c.e pounds with a safety factor of 3, the length of hardroll required is given by, F* = [ l a,c e = 1.53 inch Similarly, the required engagement length #or faulted conditions can be found to be 1.54 inch using a safety factor of 1.43 (corresponding to an ASME Code safety factor of 1.0/0.7 for allowable stress for faulted conditions).

The calculation of the above values is summarized on Table 3. The F* value thus determined for the required length of hardroll engagement below the BRT or the top of the tubesheet, whichever is lower in elevation, for faulted operation is sufficient to resist tube pullout during both normal and postulated accident condition loadings.

Based on the results of the testing and analysis, it is concluded that following the installation of a tube by the standard hardrolling process, a residual radial preload stress exists due to the plastic deformation of the tube and tubesheet interface. This residual stress is expected to restrain the tube in the tubesheet while providing a leak limiting seal condition.

43CD1:49/v v e;

2.3 LIMITATION OF PRIMARY TO SECONDARY LEAKAGE The allowable amount of primary to secondary leakage in a steam generator during normal plant operation is limited by plant technical specifications, generally to 0.35 gpm. This limit, based on plant radiological release considerations and implicitly enveloping the leak before break consideration for a through wall crack in the free span of a tube, is also applicable to a leak source within the tubesheet. In evaluating the primary to secondary leakage aspect of the F* criterion, the relationship between the tubesheet region leak rate at postulated FLB conditions is assessed relative to that at normal plant operating conditions. The analysis was performed by assuming the existence of a leak path, however, no actual leak path would be expected due to the hardrolling of the tubes into the tubesheet.

2.3.1 OPERATING CONDITION LEAK CONSIDERATIONS In actuality, the hardrolled joint would be expected to be leak tight, i.e.,

the plant would not be expected to experience leak sources emanating below F*

Since the presence of the tubesheet tube indications is not expected to r increase the likelihood that the plant would experience a significant number of leaks, it could also be expected, that if a primary to secondary leak is detected in a steam generator it is not in the tube region below F* Thus, no significant radiation exposure due to the need for personnel to look for tube i tubesheet leaks should be anticipated, i.e., the use of the F* criterion is consistent with ALARA considerations. As an additional benefit relative to ALARA considerations, precluding the need to install plugs below the F*

criterion would result in a significant reduction of unnecessary radiation exposure to installing personnel.

The issue of leakage within the F* region up to the top of the mechanical roll transition (RT) assuming the as-manufactured position of the roll transition is below the secondary side of the tubesheet includes the consideration of postulated accident conditions in which the violation of the tube wall is very extensive, i.e., that no material is required at all below F*. Based on operating plant and laboratory experience the expected configuration of any cracks, should they occur, is axial. The existence of significant circumferential cracking is considered to be of very low prcbability. Thus, 4305M:49/C4:467 i

consideration of whether or not a plant will come off-line to search for leaks a significant number of times should be based on the type of degradation that might be expected to occur, i.e., axial cracks. Axial cracks have been found both in plant operation and in laboratory experiments to be short, about 0.5 inch in length, and tight. From field experience, once the cracks have grown so that the crack front is out of the skiproll or transition areas, they arrest.

Axial cracks in the free span portion of the tube, with no superimposed thinning, would leak at rates compatible with the technical specification acceptable leak rate. For a crack within the F* region of the tubesheet, expected leakage would be significantly less. Leakage through cracks in tubes has been investigated experimentally within Westinghouse for a significant number of tube wall thicknesses and thinning lengths, reference 6. In general, the amount of leakage through a crack for a particular size tube has been found to be approximately proportional to the fourth power of the crack l("

length. Analyses have also been performed which show, on an approximate basis for both elastic and elastic plastic crack behavior, that the expected dependency of the crack opening area for an unrestrained tube is on the order of the fourth power, e.g., see NUREG CR-3464 The amount of leakage through a crack will be proportional to the area of the opening, thus, the analytic results substantiate the test results. This conclusion is supported by the results of the preload testing and analysis which demonstrated that a residual preload of about [ l ac.e psi exists between the tube and the tubesheet at normal operating conditions.

The presence of the tubetheet will preclude deformation of the tube wall adjacent to the crack, i.e., the crack flanks, and the crack opening area may be considered to be directly proportional to the length. Thc add 1&ional dependency, i.e., fourth power relative to first power, is due to the dilatation of the unconstrained tube in the vicinity of the crack and the bending of the side faces or flanks of the crack. For a tube crack located within the tubesheet and below the bottom of the roll transition, the dilatation of the tube and bending of the side faces of the crack are suppressed. Thus, a crack located within this region would be expected to leak, without considering the axial ficw resistance caused by the i

a308M:43/:a:a37 '3

tube-to-tubesheet interface, at a rate less than a similar crack in the free span, i.e., less than the Farley units technical specification limit of 0.35 gpm. Additional resistance provided by the tube to tubesheet interface would reduce this amount even further, and in the hardroll region the residual radial preload would be expected to eliminate it.

2.3.2 POSTULATED ACCIDENT CONDITION LEAK CONSIDERATIONS For the postulated leak source within the tubesheet, increasing the tube differential pressure increases the driving head for the leak and increases --

the tube to tubesheet loading. For an initial location of a leak source below _

the top of the tubesheet equal to F*, and without considering hardroll effects, the FLB pressure differential results in approximately a 10 percent increase in tiie leak rate relative to that which could be associated with normal plant operation. This small effect is reduced by the increased tube to tubesheet loading associated with the increased differential pressure. Thus, for a circumferential indication within the tubesheet region which is left in service in accordance with the pullout criterion (F*), the existing technical specification limit is consistent with accident analysis assumptions. For postulated accident conditions the preload testing and analysis showed that a residual radial preload of about [ 3a ,c.e psi would exist between the tube and tubesheet.

For axial indications in a full depth hardrolled tube below the bottom of the roll transition zone (which is assumed to remain in the tubesheet region), the tube end remains structurally intact and axial loads would b? resisted by the remaining hardrolled region of the tube. For this case, the leak rate due to FLB differential pressure would be bounded by the leak rate for a free span leak source with the same crack length, which is the basis for the accident analysis assumptions.

l Q 4325M:43i'02137

. 2[

2.3.3 OPERATING PLANT LEAKAGE EXPERIENCE FOR WITHIN-TUBESHEET TUBE CRACKS ,

A significant number of within-tubesheet tube indications, have been reported for some non-domestic steam generator units. The present attitude toward operation with these indications has been to tolerate them with no remedial action relative to plugging or sleeving. No significant number of shutdowns occurring due to leaks through these indications have been reported.

2.4 TUBE INTEGRITY UNDER POSTULATED LIMITING CONDITIONS The final aspect of the evaluation is to demonstrate tube integrity under the postulated loss of coolant accident (LOCA) condition of secondary to primary differential pressure. A review of tube collapse strength characteristics indicates that the constraint provided to the tube by the tubesheet gives a significant margin between tube collapse strength and the limiting secondary to primary differential pressure condition, even in the presence of circumferential or axial indications.

The maximum secondary to primary differential pressure during a postulated LOCA is C l a,c.e psi. This value is significantly below the residual radial preload between the tubes and the tubesheet. Therefore, no significant secondary to primary leakage would be expected to occur. In addition, loading on the tubes is axially toward the tubesheet and could not contribute to pullout.

2.5 CHEMISTRY CONSIDERATIONS The concern that boric acid attack of the tubesheet due to the presence of a through wall flaw within the hardroll region of the tubesheet may result in loss of contact pressure assumed in the development of the F* Criterion is addressed below. In addition, the potential for the existence of a lubricated interface between the tube and tubesheet as a result of localized primary to secondary leakage and subsequent effects on the friction coefficient assumed i

in the develcpment of the F* criterion is also discussed.

43C8M: 13/CO2J37 2'

2.5.1 TUBESHEET CORPOSION TESTING Corrosion testing performed by Westinghouse specifically addressed the question of corrosion rates of tubesheet material exposed to reactor coolant.

The corrosion specimens were assembled by bolting a steel (A336) coupon to an Inconel Alloy 600 coupon. The coupon dimensions were 3 inches x 3/4 inch x 1/8 inch and were bolted on both ends. A torque wrench was used to tighten the bolts to a load of 3 foot-pounds. The performance of A508 in testing of this nature is expected to be quite comparable to the performance of A336 (Gr F-1) steel (a material used for tubesheet construction prior to A508).

The arguments used in supporting the F* case relative to corrosion of tubesheet material due to minute quantities of primary coolant contacting the carbon steel were so conservative and had such margin that minor differences in material composition or strength would not change the conclusion.

The specimens were tested under three types of conditions:

1. Het-layup conditions

2. Het-layup and operating conditions

3. Operating conditions only The wet-layup condition was used to simulate shutdown conditions at high boric acid concentrations. The specimens were exposed to a fully aerated 2000 ppm boron (as boric acid) solution at 140 degrees F. Exposure periods were 2, 4, 6, and 8 weeks. Test solutions were refreshed weekly.

While lithium hydroxide is normally added to the reactor coolant as a corrosion inhibitor, it was not added in these tests in order to provide a more severe test environment. Previous testing by Westinghouse has shcan that the presence of lithium hydroxide reduces corrosion of Inconel Alloy 600 and steel in a borated solution at operating temperatures.

Another set of specimens were used to simulate startup conditions with some operational exposure. The specimens were exposed to a 2000 parts per million boron (as boric acid) solution for one week in the wet-layup condition (140 degrees F), and 4 weeks at operating conditions (600 degrees F 2000 psi).

During wet layup, the test solution was aerated but at operating conditions J308M:49/C4:457  ::

the solution was deaerated. The high temperature testing was performed in an Inconel autoclave. Removal of oxygen was attained by heating the solution in the autoclave to 250 degrees F and then degassing. This method of removing the oxygen results in oxygen concentrations of less than 100 parts per billion.

Additional specimens were exposed under operating conditions only for 4 weeks in the autoclave as described above.

High temperature exposure to reactor coolant chemistry resulted in steel corrosion rates of about 1 mil per year. This rate was higher than would be anticipated in a steam generator since no attempt was made to completely remove the oxygen from the autoclave during heatup. Even with this amount of corrosion, the rate was still a factor of nine less than the corrosion rate observed during the low temperature exposure. This differential corrosion rate observed between high and low temperature exposure was expected because of the decreasing acidity of the boric acid at high temperatures and the corrosive effect of the high oxygen at low temperatures.

These corrosion tests are considered to be very conservative since they were conducted at maximum boric acid concentrations, in the absence of lithium hydroxide, with no special precaution to deaerate the solutions, and they were of short duration. The latter point is very significant since parabolic corrosion rates are expected in these types of tests, which leads one to overestimate actual corrosion rates when working with data from tests of short duration.

Also note that the ratio of solution to surface area is high in these tests compared to the scenario of concern, i.e., corrosion caused by reactor coolant leakage through a tube wall into the region between the tuce and the tubesheet.

2.5.2 TUBESHEET CORROSION DISCUSSION At low temperatures, e.g., less than 140 degrees F, aerated boric acid solutions comparable in strength to primary coolant concentrations can produce corrosion of carbon steels. Deaerated solutions are much less aggressive and deaerated solutions at reactor coolant temperatures produce very low corrosion rates due to the fact that boric acid is a very much weaker acid at high temperature, e.g., 610 degrees F, than at 70 degrees F.

4306": 4'/C4:487 23

In the event that a crack occurred within the hardroll region of the tubesheet, as the amount of leakage would be expected to be insufficient to be noticed by leak detection techniques and is largely retained in the crevice, then a very small volume of primary fluid would be involved. Any oxygen present in this very small volume would quickly be consumed by surface reactions, i.e., any corrosion that would occur would tend to cause existing crevices to narrow due to oxide expansion and, without a mode for replenishment, would represent a very benign corrosion condition. In any event the high temperature corrosion rate of the carbon steel in this very local region would be extremely low (significantly less than 1 mil per year).

Contrast the proposed concern for corrosion relative to F* with the fact that Westinghouse has qualified boric acid for use on the secondary side of steam generators where it is in contact with the full surface of the tubesheet and other structural components made of steel. The latter usage involves concentrations of 5 - 10 ppm boron, but, crevice flushing procedures have been k conducted using concentrations of 1000 to 2000 ppm boron on the secondary side (at approximately 275 degrees F where boric acid is more aggressive than at 610 degrees F).

Relative to the lubricating effects of boron, the presence of boric acid in water may change the wetting characteristics (surface tension) of the water but Westinghouse is not aware of any significant lubricating effect. In fact, any corrosion that would occur would result in oxides that would occupy more space than the parent metals, thus reducing crevice volume or possibly even merging the respective oxides.

2.6 DETERMINATION OF EDDY CURRENT ELEVATION MEASUREMENT UNCERTAINTY In the application of the F* alternate tube plugging criterion, it is required that the distance from the BRT or the tubesheet top, whichever is lower in elevation, to the top of the degraded area of the tube be determined. This distance can be measured using conventional eddy current testing techniques.

During the inspection of the steam generators, eddy current data is collected in both the differential and absolute modes. The tube degradation is typically measured employing the differential mode, while 1ccation J E sit: 4?i04:437 _a I

l

___-_ - ~

measurements are normally taken from the absolute mode data. The absolute mode is used for length measurements because the top and the bottom of the roll transition being observed can be more easily determined. The F*

distance, plus the ECT uncertainty distance, 0.25 inch, is recAoned downward from the BRT or the tubesheet top, whichever is lower in elevation.

The physical phenomenon which is responsible for the require' ment of assigning a plus or minus value (uncertainty) to any measured value is the field spread of the eddy currents in front of and behind the eddy current coll. Recent studies conducted at the Westinghouse NDE Center have shown [

ga,b,c.e Therefore the conservative measurement uncertainty value of 0.25 inch applied to the P alternate plugging criteria can be justified by previous expertence and laboratory testing based on evaluations of addy current data recorded by Westinghouse.

3.0

SUMMARY

On the basis of this evaluation, it is d.etermined that tubes with eddy current Indications in the tubesheet region below the F* pullout criterton shown in Table 3, plus a conservative allowance for eddy current elevation measurement uncertainty, can be left in service. Tubes with circumferentially oriented eddy current indications of pluggable magnitude and located a distance less than F* below the bottom of the hardroll transition or the top of the tubesheet, whichever is lower in elevation, should be removed from service by

' plugging or repaired in accordance with the plant technical specification plugging limit. The conservativeness of the F* criterion was demonstrated by preload testing and analysis comensurate with the requirements of RG 1.121 for indications in the free span of the tubes.

4308M:49/042487 25

In addition, it has been determined, see Appendix II, that there is no need to stabilize tubes which are removed from service due to eddy current indications within the F* region.

1 e

4 NOTE. The methodology for developing the F* criterion was first reported in a

{

previous publication, reference 6, on the same subject. The difference being that the previously developed criterion, known as P*, was based on the available clearance for tube motion before it would be impeded by a neighboring tube or some other physical feature of the tube bundle. The values reported herein for F* are slightly larger than those reported for P*.

l' 1

4308M:43/04:487 :i

4.0 REFERENCES

1. United States Nuclear Regulatory Commission, Regulatory Guide 1.121,

" Bases for Plugging Degraded PWR Steam Generator Tubes," August, 1976.

2. Goodier, J. N., and Schoessow, G. J., "The Holding Power and Hydraulic Tightness of Expanded Tube Joints: Analysis of the Stress and y Deformation," Transactions of the A.S.M.E., July, 1943, pp. 489-496. $

3. Grimison, E. O., and Lee, G. H., " Experimental Investigation of Tube Expanding,", Transactions of the A.S.M.E. , July,1943, pp. 497-505) .

4. ASME Boller and Pressure Vessel Code,Section III, " Rules for Construction of Nuclear Power Plant Components," The American Society of Mechanical Engineers, New York, New York, 1983. ,

5. WCAP-11241, "Tubesheet Region Plugging Criterion for the Central Nuclear de ASCO, ASCO Nuclear Station Unit I and 2 Steam Generator", October, 1986. (Proprietary)

6. WCAP-10949, "Tubesheet Region Plugging Criterion for Full Depth Hardroll Expanded Tubes," Westinghouse Electric Corporation, September, 1985.

(Proprietary) 430Dh 49 / 04:457  ;'

Model 51 steen Generator fee Roll Pre Load Test TABLE 1. TEST DATA fest Location Coller ID Pre Rott Coller CD Pre Roll f ee to Before Roll im m gefore Rett bo. No. 3,C E 20es. 90 Dog. Avg. 0 Dog. 90 Deg. Avg. O Deg. 90 Dog. Ave. O Deg. 90 Deg. Avg 1 1 2

3 -

4 5

6 Average 2 1 2

3 4

5 6

Average 3 1 2

3 4

5 6

Average 6 1 2

3 4 ,

5 6

Average 7 1 2

3 4

5 6

Average 8 1 l

2 3

4 m 5

6 Average Col. Avest e

28

model 51 steen Generator Tube molt Pre Load Test TABLE 1. TEST DATA (Cont.)

7est Location Pre Rott Cotter 3 Post. ROLL Colter Tube 10 Post toll fthe 10 fthe ID Post tall No. me. Thicknese Dette Growth Cotter temoved a,C,e 0 Dog. 90 Dog. Ave. 0 Dog. 90 Dog. Ave. 0 Dog. 90 Deg. Avg.

1 1 2

3 4

5 6

Average 2 1 2

3 4

5 6 .

Average 3 1 2

3 4

5 6

Average 6 1 2

3 4

5 6

Average  !.

i I

7 1 2

3 4

5 6

Average 8 1 2

3 4

5 6

Average Col. Avgtt 29

l Modet 51 steen Generator T h Rott Pre Load feet TASLE 1. TEST DATA (Cont.)

feet Location T4 e 00 Post Rott Poet. Thick- Coller Radil T4e to No. No. Cotter Removed Roll nees Fles. Ratio Spring.

O Dog. 90 Dog. Avg. Thick Red.% cRJi/@f (4) Back 3,C,e 1 1 2

3 4

5 6

Average 2 1 2

3 4

5 6

Average 3 1 2

3 4

5 6

Average 6 1 2

3 4

5 6

Average 7 1 2

3 4

5 6

Average 8 1 2

3 4

5 6

Average Cat. Avge Notes: 1. ALL secoured dimensione are in inches.

2. The m strees le calculated using the measured ID opringbeck.

3. The redif retto le a terie that nopeere fremently in the enetyef e and f e fomd as (m*2+ID*2)/(cc"2 ID"2).

30

Model 51 Steen Generator Tube Roll Pre Load Test TASLE 2. STRESS ANALYSIS RESULTS Thermal Oper.

Tcst Location Tube ID T@e T@e W @ 2 Exp. T@e Pressure Total Total N3. No. Spring- Flen. Flen. Radial Noop Aalel Redfel Flem. Redfel Redfel vormises

_Bect dJi/#o eo/#o Stress Stress Stress Stress dJo/ # 1 Stress Stress Stress a .C .s

~

1 1 2

3 4

5 6

Average 2 1 2

3 4

5 6

Average 3 1 2

3 4

5 6

Average 6 1 2

3 4

5 6

Average 7 1 2

3 4

5 6

Average 8 1 2

3 4

5 6

Average Col. Avgs:

Notes: 1. The CD stress is calculated using the measured 10 springbeck.

1 I

31 4

i

- , . . , , - y_., _ _

m_ _ - _ . _ _ . __ , _ - , _ _ . - _ _ . - . - - . _ - . _ _ __ - . - - - _ _ . _ _ . .- _

I Model 51 Steam Generator Tube Roll Pre Load Test TA8LE 3. PRELCAD ANALYSIS

SUMMARY

Meterial Properties: 7 4 e/Tubesheet Dimensions (Tested): a,c,e Etestic Modulus: 2.89E+07 psi Init. Avg. Tube 00:

Poisson's Ratio: 0.30 Init. Avg. f e e thickness:

1600 Expansion: 7.79E 06 in/in/F Init. Avg. Tubesheet 10:

T/S Empenelon: 7.16E 06 in/in/F Actual Thinning:

Oper. Dette T: 483.00 F Apparent Thinning:

Normel Dette P: 1457.00 pel Faulted Dette P: 2650.00 poi Additional Analysis input:

Tubesheet low Stress Ree -*8-- ,a,c.e Coefficient of Friction:

Normel: End Effects:

Faulted:

Mean Radius (Rolled):

Lower Tolerance Limit Factor: Thickness (Rolled):

Leede 95/95 LTL: 2.1577 (N = 36) End Effect Length:

Load Factor EVALUAfl0N OF REQUIRED ENGAGEMENT LENGTN a,C,e p

TQtAL Length Required: 1,53 in. 1.54 in.

h0fts: 1. 95/95 Lower Tolerance Limit Rolled Preload used.

2. For NCAMAL Coerotion a Safety Factor of 3 was Used.

3. For FAULTED Conditione e Safety Factor of 1.43 was Used corresponding to ASME Code use of 0.7 on Ultimate strength.

4. The Required Length Does NOT Include Eddy Current inspection Uncertainty for the Location of the Bottom of the Nordroll, or the Top of the Tubesheet, Relative to the Degradation.

t 32

APPENDIX I - DISPOSITION OF TUBES WITH INDICATIONS ABOVE F*

Complementary to the criterion for leaving a tube in service with axial or circumferential indications within the tubesheet is a criterion for determining the need to stabilize tubes which are removed from service due to circumferential indications within the tubesheet. As was previously stated, ECT indications located above the F* criterion are to be dispositioned in accordance with the plant technical specification plugging limit which is based on USNRC RG 1.121, which does not distinguish between circumferential and axial cracks. Moreover, RG 1.121 is concerned with the depth of penetration of tube wall degradation, i.e., when the plugging limit is l

reached, the tube is either plugged or sleeved. RG 1.121 does not require stabilization of plugged tubes.

The kinetics of stress corrosion cracking of mill annealed Inconel 600 in primary water is highly temperature dependent. High temperatures accelerate rates of cracking. Laboratory measurements of Arrhenius relation type activation energies typically range from 30 to 75 kcal per mole. Field

experience with row I U-bends in domestic steam generators and roll transitions in foreign units indicate an activation energy of 85 kcal per mole.

4 Conditions in tubes leading to lower tube metal temperatures greatly retard the kinetics of any subsequent cracking even if applied or residual stresses are maintained. Below an assumed temperature, Thot, of 620 degrees F, cracking is retarded by a factor of 4 at 600 degrees F, a factor of 15.5 at 580 degrees F, and a factor of 64 at 560 degrees F. Moreover, the presence of hydrogen in primary water is another important consideration relative to the kinetics of cracking of Inconel 600. Laboratory measurements show that standard concentrations of hydrogen in primary water accelerates cracking by approximately a factor of 2 to 5 compared to control tests in the absence of hydrogen.

i l

430$M:49/042487 33

i for use in a materials evaluation, in determining whether a tube plugged for an eddy current indication above the F' criterion should be stabilized due to the potential for continued growth of an ID stress corrosion crack, tube temperatures within and above the tubesheet region were assessed, refer to Appendix III. A plugged tube was postulated to exist in a variety of l environments that would influence tube temperature, including the buildup of I sludge around the tube, as the sludge may act as an insulator and alter the heat conduction patterns and surface metal temperature of the tube.

For conservatism, active tubes adjacent to the plugged tubes, and the tubesheet itself except at the secondary surface are assumed to be at primary fluid temperature. For this tubesheet temperature condition, several sludge deposition cases were hypothesized:

I CASE 1 considers no sludge buildup adjacent to the tube and the tube does not have a through wall penetration prior to plugging. Certain conditions in tubes (such as wet walls prior to plugging) may lead to the presence of superheated steam existing within the tube. Limited data on l Inconel 600 at high temperatures is consistent with general observations on aluminum and steel alloys in low temperature water vapor. At low

. superheat, i.e., high relative humidity, the cracking response in water vapor is essentially equivalent to that in the liquid phase at the same I temperature, while at high superheat, i .e. , low relative humidi ty, the cracking kinetics are much reduced. A plugged tube is essentially dry on

, its ID when plugged; therefore, although the ID temperature of the tube l In the region within the tubesheet would most likely be eaulvalent to Thot, the ratio of the vapor pressure of any water trapped in the tube during plugging to the pressure of saturated water vapor would be low, l.e., high superheat, thus greatly reducing the cracking kinetics. Also, as previously discussed, the lack of the presence of hydrogen in a plugged tube significantly retards further cracking. Therefore, combining the above two effects, the probability a plugged tube with degradation that has not progressed through wall would continue to degrade is small in this environment and would not require t

4308M:49/04:437 3:

l stabilization. It is noted that this case is really independent of

! whether or not sludge is postulated to be present, i.e., the temperature inside the tube in the tubesheet region will be near Thot regardless of the presence of sludge.

4

, CASE 2 considers no sludge buildup either adjacent to a plugged tube or in a plugged tube. A through wall indication is postulated and the tube is filled with water due to the ingress of secondary side water through the penetration. The water contained in the tube in the tubesheet area boils. This rapid heat transfer mechanism maintains the tube inner diameter metal surface at or slightly above T sat for the portion of the tube in the tubesheet. The relatively low secondary side temperature will significantly inhibit the continuation and/or initiation of stress corrosion cracking. Therefore, considering both the effects of the reduced secondary side temperature and the lack of the presence of any hydrogen concentration on continued stress corrosion cracking, the probability of a plugged tube with a through wall penetration continuing to degrade is very small and would not require remedial action other than plugging or sleeving.

CASE 3 considers the effect of sludge buildup on the tubesheet adjacent to a plugged tube with a through wall penetration. Since the sludge acts as a poor conductor, the mechanisms for cooling the tube are not as efficient as for the previous two cases. If the secondary side water ingress remains primarily in liquid form with some localized boiling at the tube wall, and the sludge pile depth is less than about 4 inches, the temperature on the inner diameter of the tube will probably be slightly above T sat.

As discussed previously, certain conditions in plugged tubes may lead to the presence of superheated steam rather than liquid water the through wall degraded tube. It was also stated that at low superheat, the cracking response in water vapor is essentially equivalent

) to that in the liquid phase at the same temperature. At sludge depths j greater than about 8 inches, the tube metal temperature in the tubesheet j approaches plant hot leg temperature. The effect of low superheat and

! higher temperatures could result in additional crack growth. However, i

I

4308M;49/042487 35

l the above two scenarios are not expected to occur. In steam generators with a flow distribution baffle, sludge buildup to a height of 8 inches

'is precluded by geometry constraints. Moreover, as the postulated crack would most likely limit the ingress of the secondary side water in the through wall degraded tube, the most likely scenario would be that the tube is essentially dry on the inside and the ratio of the vapor pressure of the water to its saturation pressure is relatively low, thereby greatly reducing the crack kinetics. With the lack of the presence of any hydrogen concentrations, the potential for additional crack growth would be significantly reduced; therefore, tube stabilization is not

]

required.

An extension to CASE 3 could be postulated such that the throughwall penetration is of such size as to initially admit water into the plugged tube. The water subsequently boils and the internal pressure prevents any further water from entering the tube. In this case the steam would be at the secondary side pressure, i.e., high superheat conditions, and for reasons cited in consideration of case 1, further crack growth would not be expected.

i i

4303M;49/042487 36

APPENDIX II - TUBE HALL TEMPERATURES OF PLUGGED TUBES To assess whether further degradation due to postulated PHSCC can occur in a plugged tube and to disposition tubes with indications above the F* criterion i as to whether they should be stabilized when plugged, the metal temperature of the tube inside diameter at elevations above the F* criterion, but, below the

top of the tubesheet, was evaluated. Active tubes adjacent to the plugged

{

tube, and the tubesheet itself except at the secondary surface, are at the primary fluid temperature. For a tubesheet temperature condition equivalent to Thot, five different sludge deposition cases were hypothesized.

i j 1. An intact tube without sludge deposition on the tubesheet f 2. A perforated tube without sludge on the tubesheet

3. An intact tube with sludge deposition on the tubesheet.

4. A perforated tube with sludge deposition on the tubesheet.

5. A perforated tube with/without sludge deposition on the tubesheet without secondary water ingress.

An intact tube is defined as a plugged tube with no through-wall penetration f.e., no secondary water comes in contact with the tube inner wall, while a perforated tube is defined as a tube with a through-wall penetration i.e.,

secondary water comes in contact with the tube inner wall.

I.

1.1 Intact Tube Without Sludge Deposition With the exception of a shallow layer at the tubesheet surface, the l tubesheet metal temperature adjacent to active tubes just below the top surface of the tubesheet is expected to be at primary coolant inlet temperature (i.e. Thot) for the hot leg side of the tube bundle.

I i

l 4308M:43/042 37 37

.1

, Therefore, the outer wall temperature can be as high as T hot for a full

. depth hardroll expanded tube. For an intact tube, the inner wall of the tube is essentially dry. The inner wall maximum tube temperature along I

the length of the tubesheet would approach T hot

• 1.2 A Perforated Tube Without Sludge Deposition Once a plugged tube is perforated, secondary water can ingress into the primary side of the inactive tube. The water contained in that portion of the tube within the tubesheet boils. This rapid heat transfer mechanism keeps the inner tube wall temperature at approximately Tsat + 5'F (allowing for a localized wall superheat effect).

1.3 Intact Tube With Sludge Deposition i

l With sludge accumulation on the top of the tubesheet, the whole depth of the tubesheet is expected to be at Thot. As is the case without sludge j deposition, the inner wall of the inactive intact tube would be j essentially dry with a maximum temperature of T hot anticipated along the length of the tubesheet. .

i 1.4 A Perforated Tube With Sludge Deposition l

Similar to the case of a perforated tube with sludge deposition, secondary water can ingress into the primary side of the inactive tube.

The heat transfer mechanism for cooling the tube inner wall metal temperature would be the same as with the case of no sludge deposition on l the tubesheet. The inner wall temperature would be at approximately

{ Tsat + 5'F because of the boiling occuring inside the tube.

1.5 A Perforated Tube Without Communication i A situation could develop such that only a limited amount of secondary water would initially leak into a tube with a through-wall penetration, i

I 4308M:49/042487 38 i

I It can be postulated that the small amount of water ingressing into the

] tube inner diameter could evapcrate and form superheated steam within the

]'

depth of the tubesheet or the tubesheet plus the height of the sludge.

This case would be similar to an intact tube as the superheated steam would prevent the water from entering into the primary side of the tube.

The inner wall of the tube would essentially be in a dry condition and the maximum inner wall metal temperature would be T hot' i.

In summary, the inner wall temperature for the perforated tube within the depth of the tubesheet, both with or without sludge deposition, is essentially at Tsat + 5'F when there is water communication due to a through wall penetration. The inner wail temperature for a perforated tube without water

) communication could be as high as Thot. Finally, the inner wall temperature i

for an intact tube with or without sludge deposition could be as high as T

hot" t

i I

I i

l i

I i

I I

i I

I i

j i

i 1

43C8M

4)/042487 39 y e , , w- w y wwvvu =- ew m v w, -r-- e--v- w-vvww,v-v--**-v-^-~w-w---v-r---~+-+=--

m- ----v,-re---------e-e,_v me- --+,r--- .c 4

--,----r. -