ML20211P336
| ML20211P336 | |
| Person / Time | |
|---|---|
| Site: | Trojan File:Portland General Electric icon.png |
| Issue date: | 02/28/1987 |
| From: | WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19292G835 | List: |
| References | |
| WCAP-11315, WCAP-11315-R01, WCAP-11315-R1, NUDOCS 8703020345 | |
| Download: ML20211P336 (32) | |
Text
dESTINGHOUSE CLASS 3 NCAP 11315 Revision 1 TUBESHEET REGION PLUGGING CRITERION FOR THE PORTLAND GENERAL ELECTRIC COMPANY TROJAN NUCLEAR STATION September 1986 Revised February 1987 NESTINGHOUSE ELECTRIC CORPORATION SERVICE TECHNOLOGY DIVISION P. O. BOX 3377 PITTSBURGH, PENNSYLVANIA 15230 i
0703020345 870220 PDR ADOCK 05000344 P
PDR I
ABSTRACT An evaluation was performed to develop a plugging criterion, known as the F' criterion, for determining whether or not repairing or plugging of partial 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 51A 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 0.91 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.
3801H:49/021287 2
CONTENTS TOPIC PAGE i
1.0 INTRODUCTION
5 l
2.0 EVALUATION 6
2.1 Determinatior, of Elastic Preload 8etween the 8
Tube and Tubesheet 2.1.1 Radial Preload Test Configuration Description 8
2.1.2 Preload Test Results Discussion and Analysis 9
2.1.3 Residual Radial Preload During Plant Operation 11 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 20 Hithin-Tubesheet Tube Cracks 2.4 Tube Integrity Under Postulated Limiting Conditions 21 2.5 Chemistry Considerations 21 2.5.1 Tubesheet Corrosion Testing 21 2.5.2 Tubesheet Corrosion Discussion 23 3.0
SUMMARY
24
4.0 REFERENCES
26 380lM:49/021287 3
TABLES Table 1.
Model 51A Preload Measurement Tests - TEST DATA 27 Table 2.
Model SlA Preload Measurement Tests - STRESS 30 ANALYSIS RESULTS Table 3.
Model 51A Tube Roll - PRELOAD ANALYSIS SU M RY 31 k
380lH:49/021287 4
l TUBESHEET REGION PLUGGING CRITERION FOR PARTIAL DEPTH HARDROLL EXPANDED TUBES 1
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 partial 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 Portland General Electric Company, Trojan Nuclear Station 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.
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 I
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, which is roll expanded into the tubesheet. This evaluation applies to the Trojan Westinghouse Model 51A steam generators and assesses the integrity of the tube bundle, for tube ECT indications occurring on the roll expanded length of tubing within the tubesheet, relative to:
- 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.
380lM:49/021287 5
I
- 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.
l 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 for which tube degradation of any extent does not necessitate remedial action, e.g., plugging or sleeving. 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 Trojan steam generators was found to be 0.91 inch based on preload analysis. 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 the tubesheet region.
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 roll expansion (RE) 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 i
same as for a nondegraded tube, instead of acting on the tube material.
- Thus, l
axially oriented linear indications, e.g., cracks, cannot lead to tube failure within the RE and may be considered on the basis of leakage effects only.
l 380lH:49/021287 6
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 developed 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 than 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.
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 the hardroll, no significant axial forces will be transmitted further along 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 51A 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.
3801H:49/021287 7
2.1 DETERMINATION OF ELASTIC PRELOAD BETHEEN 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 partial l
depth hardrolling operation. The data generated in these tests have 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 Hestinghouse Mode) SIA steam generators.
2.1.1 RADIAL PRELOAD TEST CONFIGURATION DESCRIPTION The test program was designed to simulate the interface of a tube to tubesheet partial depth hardroll for a h) del 51A steam generator.
The test configuration consisted of sir cylindrical collars, approximately [ l.c.e a
inches in length, [
]"'C inches in outside diameter, and (
l,c.e a
inch in inside diameter. A mill annealed, lnconel 600 (ASME SB-163), tubing specimen, approximately ( 1
'C inches long with a nominal (
8 J
c.e a
outside diameter before rolling, was hard rolled into each collar using a process which simulated actual tube installation conditions.
3801M:49/021287 8
l
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 diameter was selected to provide the same radial stiffness as the tubesheet.
The collars were fabricated from AISI 1018 carbon steel similar in mechanical l
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 10.
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 00 measurements 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 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 tubes and collars.
These consisted of the ID and OD of the tubes prict to and after rolling and removal from the collars as well as the inside and outside 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.
l Using the measured and calculated physical dimensions, an analysis of the tube deflections was performed to determine the amount of preload radial stress l
380lH:49/021287 9
present following the hardrolling. The analysis consisted of application of conventional thick tube equations to account for variation of structural 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 tile 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+C
/r (1) j 2
where, u - radial deflection r - radial position within the tube wall, and the constants, C) and C2 are 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 radii, and the applied internal and external pressures.
The difference between an analysis assuming plane stress and one assuming plane strain is manifested only in a change in the constant C '
2 The first constant, is the same for both conditions.
For materials having a Poisson's ratio of 0.3, the following relation holds for the second constant:
C (Plane Strain)'- 0.862
- C (Plane Stress)
(2) 2 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 l
l 380lH:49/021287 10
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 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 and 00 respectively, e.g., dUl/dPo is the ratio of the radial deflection at the ID due to an OD pressure. Thus, dUl/dPo was used to find the interface pressure and radial stress between the tube and the tubesheet as:
S
- - P, - - (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 3 c.e psi and [
a l.c.e psi respectively.
a deviation was found to be [
In order to determine a value to be used in the analysis, a tolerance factor a
a 3 c.e percent confidence to contain [
l.c.e percent of the for [
3,c.e useable data points, to a
population was calculated, considering the [
a a
l.c.e Thus, a [
3,c.e lower tolerance limit (LTL) for the be [
J,c.e psi.
a radial residual preload at room temperature is [
2.1.3 RESIDUAL RADIAL PRELOAD DURING PLANT OPERATION During plant operation the amount of preload will change depending on the l
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 380lH:49/021287 11
decreases with temperature. Both of these effects may result in the preload being reduced upon return to ambient temperatsrc cnditions, i.e., in the cold condition. However, as documented in Reference 5, for a similar investigation, tube pullout tests which were preceded by a very high thermal relaxation soak showed the analysis to be conservative.
The plant operating pressure influences the preload directly based on the application of the pressure load to the ID of the tube, thus increasing the i
(
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 bow increase of preload for primary-to-secondary pressure differential will occur at the bottom, interior part cf the tubesheet.
Since F* is measured from the bottom of the hardroll transi ion (BRT) 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 tightening or loosening during operation is considered for evaluation.
The central location case is not the most stringent case for normal operation and FLB. The most stringent case for Normal Operation and FLB involves a peripheral tube.
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 552*F is 7.80*10-6 in/in/'F.
That for the steam generator tubesheet is 7.18*10-6 in/in/'F. Thus, there is a net difference of 0.62*10-6 in/in/*F in the expansion property of the two materials. Considering a temperature difference of 492*F between ambient and operating j
conditions the increase in preload between the tube (t) and the tubesheet (ts) was calculated as:
SrT - (0.62E-6)*(492)*(Collar ID) / 2 / ((dU1/dPi)ts-(duo /dPo)t)
(4) 3801H:49/021287 12
This calculation was performed and tabulated in Table 2. The results indicate that the increase in preload radial stress due to thermal expansion is [
3"'C psi. It is to be noted that this 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 3,c.e psi. The internal pressure acting on the a
generator is [
wall of the tube will result in an increase of the radial preload on the order of the pressure value.
The increase was found as:
SrP = - Po - - Pq (duo /dP1) / ( (dui/dPi)ts - (duo /dPo)t)
(5)
In actuality, the increase in preload will be more dependent on the internal pressure of the tube since water at secondary side pressure would not be expected between the tube and the tubesheet.
i Results from the performance of this calculation are tabulated in Table 2. for normal operating conditions and summarized on the summary sheet for both normal and faulted conditions. The results l,c.e a
indicate that the increase in preload radial stress is [
l c.e psi for faulted a
psi for normal operating conditions and [
(FLB) operating conditions.
3.
Tubesheet Bow Tightening or Loosening - An analysis of the Model 51A tubesheet was performed to evaluate the increase of preload stress that would occur as a result of tubesheet bow for interior tubes.
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 bottom 380lH:49/021287 13
surface and then the presence of the holes was accounted for.
For the location where the increase of preload is a maximum, the radial l,c.e psi during normal a
preload stress would be increased by [
operation and [
]"'C'8 psi during faulted (FLB) operating conditions.
However, the interior tubes are not the limiting case for primary-to-secondary pressure differential. The limiting case involes peripheral tubes where tubesheet bowing has a negligible effect on tube-to-tubesheet preload. Therefore, the analyses will address only tubes in the peripheral region of the tubesheet. During
(
LOCA, the differential operating pressure is from secondary to l.c.e p3g a
primary. Thus, the radial preload will decrease by [
as the tubesheet bows downward. However, the action of the differential pressure is such that the tube is pushed toward the 4
tube-to-tubesheet weld. This case is of no consequence to the determination of F*
- Combining the room temperature hardroll preload with the thermal and l,c.e a
pressure effects results in a net operating preload of [
psi during normal operation and [
]"'C psi for faulted operation.
In addition to restraining the tube in i:he 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 Trojan steam generators is based on determining the length of hardroll necessary to equilibrate the applied loads during the maximum normal operating conditions or faulted conditions, whichever provides the largest value.
Thus, the applied 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.
380lH:49/021287 14
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, for a tube with internal degradation, e.g., cracks oriented at an angle to the axis of the tube, the internal pressure may also act on the flanks of the 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 tubosheet hole diameter has been used to determine the magnitude of the pressure forces acting on the tube.
The forces a
acting to remove the tube from the tubesheet are [
J.c.e pounds and
[
l,c.e pounds respectively for normal and faulted operating a
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 The analysis for the radial preload pressure between the tube and the tubesheet made no consideration of the effect of the expanion method change, if any at the hardroll transition to the explosive-expanded ("HEXTEX") length of tubing. As a limiting case, the assumption was made that, contrary to known results, the "HEXTEX" section was ignored.
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 distance from each assumed discontinuity the stiffness, and hence the radial preload, of the tube is reduced relative to that remote from the ends of the RE.
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 3801H:49/021287 15
end, the deflection of points away from the end relative to the end deflection is given by:
=e
- cosine ( A
- x )
(6) rx / u u
ro where, X - [
]"'C End Effect Constant.
x - Distance from the end of the tube.
For the radially preloaded tube, the distance for the end effects to become l
negligible is the location where the cosine term becomes zero. Thus, for the l
roll expanded Model 51A tubes the distance corresponds to the product of "A"
(
times "x" being equal to (pi/2) or [
J,c.e inch.
a The above equation can be integrated to find the average deflection over the affected length to be 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 a
more than [
l.c.e inch from the ends.
For example, for the normal a
l.c.e psi or [
a J c.e pounds per inch operating net preload of [
l.c.e inch from a
of length, the effective preload for a distance of [
l,c.e pounds per inch or [
l c.e pounds.
a a
the end is [
2.2.3 CALCULATION OF ENGAGEMENT DISTANCE REOUIRED. 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
[
l c.e reference 5.
For normal operation the radial preload is a
a l
[
l c.e psi or [
a 3 c.e pounds per inch of engagement. Thus, the I
axial friction resistance force is [
l c.e pounds per inch of a
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 3801H:49/021287 16
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 [
l.c.e pounds. Considering both ends of the presumed a
severed tube, i.e., the hardroll transition is considered one end, the axial resistance is [
l c.e pounds plus the resistance of the material between a
a the ends, i.e., the total length of engagement minus [
l.c.e inch. For example, a one inch length has an axial resistance of, I
j. c. e.
a g
l c.e a
Conversely, for the maximum normal pressure applied load of [
l,c.e pounds with a sa.fety factor of 3, the a
pounds, considered as [
length of hardroll required is given by, F* - [
f,c.e - 0.91 inch Similarly, the required. engagement length for faulted condition, can be found to be 0.72 inch using a safety factor of 1.43 (corresponding th a ASME Code safety factor of 1.0/0.7 for allowable stress for faulted conditions).
The calculation of the above values is summarized in Table 3.
The F* value thus determined for tha required length of hardroll engagement below the BRT, for normal operation is sufficient to resist tube pullout during both normal and postulated accident condition loadings.
Based on the results of the testin:1 and analysis, it is concluded that following the installation of a tube by the standard hardrolling process, a l
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.
The resistance-to-leakage of the HEXTEX section.is ignored for the purpose of this l
evaluation.
3801H:49/021287 17
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 cont.iderations and implicitly enveloping the leak before break consideration for a throughwall crack in the free span of a tube, is also applicable to a leak source within the tubesheet. In evaluating the primary to secondary j
leakage aspect of the F* criterion, the relationship between the tubeshtet 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 ~
QEERATING CONDITION LEAK CONSIDERATIONS j
In actuality, the hardrolled joint would be expected to be leak tight, i.e.,
the plant would not be expected ta experience leak sources emanating below F*.
Because of the presence of the tubesheet, tube indications are not expected to increase the likelihood that the plant would experience a significant number of leaks.
It could also be expected, that if primary to secondary leakage is detected in a steam generator it will not be in the tube region below F*.
Thus, no significant radiation exposure due to the need for personnel to look for tube 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 roll transition (RT) includes the consideration of postulated accident conditions in which the violation of the tube wall is tery 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 2xistence of significat circumferential cracking is considered to be of very low probability. Thus, consideration of whether or not a plant will come 3801H:49/021287 18
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.
l Axial cracks in the free span portion of the tube, with no superimposed thinning, would leak at rates compatible with the technical specification l
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 Hestinghouse 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 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.
The presence of the tubesheet 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.
The additional 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, the dilatation of the tube and bending of the side faces of the crack are suppressed.
Thus, a 0.5 inch crack located within the F*
region up to the top of the roll transition would be expected to leak, without considering the flow path between the tube and tubesheet, at a rate less than a similar crack in the free span, i.e., less than the Trojan units technical specification lin.it of 0.35 gpm.
Leakage would be expected to'be about equal to that from a 0.0625 inch free span crack. Additional resistance provided by the tube-to-tubesheet interface would reduce this amount even further, and in 380lH:49/021?87 19
the hardroll region the residual radial preload would be expected to eliminate it. This conclusion is supported by the results of the preload testing and a
analysis which demonstrated that a residual preload of about [
3,c.e psi exists between the tube and the tubesheet at normal operating conditions.
2.3.2
. POSTULATED ACCIDENT CONDITION LEAK CONSIDERATIONS For the postulated leak source within the RE, 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 BRT equal to F*, the FLB pressure differential results in an insignificant 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 l
assoc.iated with the increased differential pressure as well as the tightening l
contribution of the tubesheet bending.
Thus, for a circumferential indication within the RE 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 net radial preload of about
[
l,c.e psi would exist between the tube and tubesheet.
a For axial indications in a partial depth hardrolled tube below the BRT of the roll transition zone (which is assumed to remain in the tubesheet region), the tube end remains structurally intact and axial loads would be 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.
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 present has been to tolerate them with no 380lH:49/021287 20
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 I
circumferential or axial indications.
I The maximum secondary to primary differential pressure during a postulated l,c,0 psi. This value is significantly below the residual a
LOCA is [
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 an6 tubesheet as a result of localized primary to secondary leakage and subsequent effects on the friction coefficient assumed in the development of the F* Criterion is also discussed.
2.5.1 TUBESHEET CORROSION TESTING Corrosion testing performed by Hestinghouse 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 3801H:49/021287 21
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 P 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 shown 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 condit. ions (600 degrees F, 2000 psi).
During wet layup, the test solution was aerated but at operating conditions 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.
380lH:49/021287 22
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 tube 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.
3801H:49/021287 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 l
other structural components made of steel. The latter usage involves concentrations of 5 - 10 ppm boron, but, crevice flushing procedures have been 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 Hestinghouse is not aware o'f 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.
3.0
SUMMARY
On the basis of this evaluation, it is determined that tubes with eddy current indications in the tubesheet region below the F* pullout criterion shown in Table 3 can be left in service.
Tubes with circumferential1y oriented eddy current indications of pluggable magnitude and located a distance less than F*
below the bottom of the hardroll transition 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 commensurate with the requirements of RG 1.121 for indications in the free span of the tubes.
3801H:49/021287 24
For tubes with axial indications, the criterion which should be used to determine whether tube plugging or repairing is necessary should be based on
. leakage since the axial strength of a tube is not reduced by axial cracks.
Under these circumstances it has been demonstrated that significant leakage would not be expected to occur for through wall indications greater than
[
l.c.e inch below the bottom of the hardroll transition.
a i
NOTE.
The methodology for developing the F* criterion was first reported in a l
previous publication, reference 6, on the same subject.
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 smaller than those reported for P*.
380lH:49/021287 25
4.0 REFERENCES
1.
United States Nuclear Regulatory Commission, Regulatory Guide 1.121,
" Bases for Plugging Degraded PHR 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 Deformation," Transactions of the A.S.H.E., July,1943, pp. 489-496.
3.
Grimison, E.
D., and Lee, G.
H., " Experimental Investigation of Tube Expanding,", Transactions of the A.S.M.E., July,1943, pp. 497-505).
r 4.
ASME Boiler 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.
~ l, 5.
HCAP-11241, "Tubesheet Region Plugging Criterion for the Central Nuclear de ASCO, ASCO Nuclear Station Units 1 and 2 Steam Generators", October, 1986.
(Proprietary) 6.
HCAP-10949, "Tubesheet Region Plugging Criterion for Full Depth Hardroll Expanded Tubes," Hestinghouse Electric Corporation, September,1985.
(Proprietary) l 3801H:49/021287 26
Model 51A steam Generator Time Rott Pre Load Test TABLE 1.
TEST DATA Tos2 Location Cottar ID Pre Roll Cottar CD Pre Roll TLhe ID Sefore Rott TLee CD 8efore Rott do.
No.
a,C,E A Deg. 90 Dog.
Avg.
O Dog. 90 Dog.
Avg.
O Dog. 90 Dog.
Avg.
0 Dog. 90 Dog.
Avg.
I 1
2 3
4 5
0 6
Average 2
1 2
3 4
5 6
Average 3
1 2
3 4
5 6
Average 6
i L.
2 3
4 5
6 Avarege F
1 2
3 4
5 6
Average l
l 0
1 2
3 4
5 6
l Average Col. Avgs:
e 27
Model 51A Steam Generator TLhe Rott Pre Load Test TABLE 1.
TEST DATA (Cont.)
Test Location Pre Roll cotter CD Post ROLL Collar TLhe ID Post Roll Tube ID TLhe ID Post Roll No.
No.
Thickness Delta Growth Cotter Removed 8,C,e 0 Dog. 90 Deg.
Avg.
O Deg. 90 Dog.
Avg.
O 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
l Average l
7 1
2 3
4 5
6 Average 8
1 2
3 4
5 6
Average Col. Avgs:
28
Model 51A Steam Generator 7the Actt Pre Load Test TABLE 1.
TEST DATA (Cont.)
Tsst Location Tihe CD Post Rolt Post-Thick-Colter Radil Tube 10 No.
No.
Collar Removed Rott nesa Flex.
Ratio Spring-0 Dog. 90 Dog.
Avg.
Thick Red.Ie dJi/#1 (4) sack a,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 Col. Avgs:
Notes: 1. Alt measured dimensions are in inches.
- 2. The CD stress is calculated usire the measured ID springback.
- 3. The radif ratio is e ters that appears frequently in the analysis and is fomd as (OD*2+ID"23/(CD*2 ID"2).
29
Model $1A Steam Generator Tube Roll Pre Load Test TABLE 2.
STRESS ANALYSIS RESULTS Thermal Oper.
Test Location T @ e 10 T@e Tse 00 CD 00 Exp.
T@e Pressure Total Total No.
No.
Spring-Flex.
Flex.
Radial Hoop Axial Radial Flex.
Radial Radial vorv41ses Back dJi/&o dJo/dPo Stress Stress Stress Stress dJo/@ l Stress Stress Stress 8,C.
1 1
2 3
4 5
6 Average 2
1 2
3 4
5 6
i Average i
l 3
1 2
3 4
5 6
Average 6
1
~
2 3
4 5
6 Average 7
1 2
3 4
5 6
Average 8
1 2
3 4
5 6
Average Col. Avgs:
Notes: 1. The CD stress is calculated using the measured ID springback.
30
Modet 51A steam Generator Tthe Roll Pre Load Test T.3LE 3.
PRELCAD ANALT5!$ sumART Material Properties:
Ttee/Tthesheet Dimensions (Tested):
3,c,e Elastic Modatus 2.87E+07 pel Init. Avg. TLhe 00:
Poisson's Ratio:
0.30 Init. Avg. TLhe Thickness:
I6CO Expension:
7.80E 06 in/in/F Init. Avg. TLbesheet 10:
T/s Expension:
7.18E 06 in/in/F Actual Thiming:
Oper. Detta 7:
492.00 F Apparent Thiming:
Normel Delta Pt 1340.00 pel Faulted Delta P 2650.00 pel Additional Analysis Input:
Tthesheet Bow Stress Redtartion a,C,e coefficient of Friction:
Normult End Effects:
Faulted:
Mean Radius (Rotted):
Lower Tolerance Limit Factor:
Thickness (Rolled):
Lashda 95/95 LTL:
2.1577 (N = 36)
End Effect Leeeth:
Load Factor EVALUATION OF RdQUIRED ENGAGEMENT LENGTM a C,e Elastic Analysis:
RT Preload (LTL)
Thermal Expansion Preload Pressure Preload TLbesheet Bow Loss NET Preload NET Radial Force NET Axial Resistance Applied Load:
Analysis Load:
End Ef fect Resistance (2):
NET Analysis Load:
Length Required:
TOTAL Length RegJired:
0.91 in.
0.72 in.
=
sessnes==
NOTES:
- 1. 95/95 Lower Tolerance Limit Rotted Preload Used.
- 2. For NORMAL Operation a Safety Factor of 3 we.s used.
- 3. For FAULTED Conditiene a Safety Factor of 1.43 was used Corresponding to ASME Code use of 0.7 cn Ultimate Strength.
- 4. The Required Length Does NOT Include Eddy Current inspection Uncertainty for the Location of the Bottom of the Nordrott, or the Top of the Tthesheet, Relative to the Degradation.
- 5. Pretoed stresses used were for the most stringent case, i.e.,
cold teg, peripheral location. This minimizes the thermal expansion preload and eliminates the preload chas to tubesheet bowing.
31
____