ML051300624

Draft ECR #02-01121, Rev. 2, Inspection Acceptance Criteria and Leakage Assessment Methodology for TMI OTSG Kinetic Expansion Examinations.

ML051300624
Person / Time
Site:	Three Mile Island
Issue date:	05/03/2005
From:	AmerGen Energy Co
To:	Office of Nuclear Reactor Regulation
References
ECR #02-01121, Rev 2
Download: ML051300624 (110)

Text

er~ieSM An Exelon Company ECR #02-01121, Rev. 2 DRAFT I INSPECTION ACCEPTANCE CRITERIA AND LEAKAGE ASSESSMENT METHODOLOGY FOR TMI OTSG KINETIC EXPANSION EXAMINATIONS ORIGINATOR(S):

DATE APPROVALS:

INDEPENDENT REVIEWER DATE ENGINEERING MANAGER DATE I

ECR # 02-01121, Rev. 2 l

INSPECTION ACCEPTANCE CRITERIA AND LEAKAGE ASSESSMENT METHODOLOGY FOR TMI OTSG KINETIC EXPANSION EXAMINATIONS TABLE OF CONTENTS SECTION PAGE 1.0 PURPOSE .................................. 4 2.0 SYSTEM PERFORMANCE/KINETIC EXPANSION STRUCTURAL INTEGRITY ANALYSIS ............................... 4 3.0 MATERIAL CONDITION ASSESSMENT .............................. 22 4.0 BASIS FOR DISPOSITION OF INDICATIONS AND NDE PROCESS VARIABILITY .................................. 30 5.0 LEAKAGE ASSESSMENT METHODOLOGY .................................. 40 6.0 INSPECTION CRITERIA AND LEAKAGE ASSESSMENT

SUMMARY

......... 75

7.0 REFERENCES

......... 76 LIST OF FIGURES FIGURE PAGE Figure 1: 17" Kinetic Expansion (Typical) .13 Figure 2: "Defect Free" Concept .40 Figure 3: Leakage Evaluation Methodology Overview .44 Figure 4: Three Mile Island Unit I RETRAN Two Loop Model .46 Nodalization Diagram Figure 5: Break Nodalization .48 Figure 6: RPV Nodalization .49 Figure 7: Main Feedwater Flow Rates .51 Figure 8: Emergency Feedwater Flow Rates .52 Figure 9: Reactor Power .54 Figure 10: RCS Faulted Loop Temperatures .55 Figure 11: Steam Generator Pressure Response .56 Figure 12: Pressurizer Response .57 Figure 13: GOTHIC Model For Shell Cooldown Analysis .59 Figure 14: MSLB Temperature Response .60 Figure 15: MSLB Pressure Response .61 Figure 16: GPUN Tube Loads .64 2

ECR#02-01121,Rev.2

LIST OF FIGURES (continued)

FIGURE PAGE Figure 17: FTI Tube Loads .......................................................... 64 Figure 18: Tube Load Comparison ............................ .............................. 65 Figure 19: Summary of Finite Element Results for 17" Expansions .................. 11 Figure 20: Comparison of Table I Values to Model Results (17" Expansions)..... 12 Figure 21: Comparison of Table I Values to Model Results (22" Expansions)..... 19 Figure 22: B&W Rolled Sleeve .......................................................... 21 Figure 23: Growth of ID Kinetic Expansion VOLs, TMI-1, SG A ........ ............ 23 Figure 24: Distribution of IR15 Kinetic Indications ............. ....................... 36 LIST OF TABLES TABLE PAGE Table A: Summary of Finite Element Modeling Results for 17" Expansions .. 111.....

Table B: [deleted] .......................................................... 29 Table C: OTSG Machined Flaws in 1997 Study ................ ....................... 32 Table D: Historical Growth Summary for Kinetic Expansion Volumetric IDIGA Indications........................................................................ 24 Table E: Growth of Kinetic Expansion Circumferential Indications ....... ......... 24 Table 1: Inspection Acceptance Criteria for OTSG Kinetic Expansion Region (Required Expansion Length) .................................... .............. 79 Table 2: Inspection Acceptance Criteria for OTSG Kinetic Expansion Region (Flaw Dispositioning Criteria) . ................................................. 80 Table 3: [deleted]............................................................................. 81 Table 4: Leakage Assessment Evaluation Data ....................................................... 82 Table 5: In Situ Pressure Test Data Summaries .......................................... 85 Table of Acronyms .......................................................... 89 3

ECR 02-0112 1,Rev. 2

1.0 PURPOSE TMI-l 's Once Through Steam Generator (OTSG) tubes were repaired in 1982 - 1985 by forming new tube-to-tubesheet joints within the upper tubesheets using a kinetic expansion process. In 1997 GPU Nuclear (the prior owner of TMI-1) developed inspection criteria for use during ECT inspection of the kinetically expanded regions and these criteria were submitted to the NRC (References 25 and 26). In 2002 a single AmerGen document (Revision 0 of this ECR) was created to update those two 1997 submittals. This 2005 Revision 2 of this ECR, like the 2004 revision, was provided to incorporate additional information. Data from examinations of the kinetic expansions in the 1997 to 2003 outages is incorporated. In addition, this revision makes significant changes to the kinetic expansion criteria that further increase its conservatism:

- a 100% scope is implemented so that each in-service kinetic expansion is examined during each refueling outage,

- circumferentially-oriented flaw indications are removed from pressure boundary service,

- newly-identified flaws are removed from pressure-boundary service, and

- revisions to the leakage assessment methodology result in a more conservative (i.e.,

greater volume) estimate of accident-induced kinetic expansion leakage.

These inspection criteria identify the minimum required length of defect-free kinetically expanded tube that must be present, and provide acceptance criteria for any flaws that may be encountered, in order to ensure that the design capability of the joints is maintained. These criteria also ensure that margin is provided in depth against unacceptable performance of the joints (to prevent joint slipping, parting of the tube, or unacceptable accident-induced leakage.)

The purpose of this document is also to provide a summary of the conservative methods that are used to inspect and disposition the kinetically expanded joints. An assessment of the material condition of the joint is presented as regards the benefit of the residual stresses from formation of the joint in mitigating stress corrosion cracking. It is also shown that NDE performance characteristics for the several forms of potential damage in the joint are applied conservatively.

This document also provides the inspection methodology, inspection scope, acceptance criteria, and reporting requirements to be implemented during the kinetic expansion examinations.

This document is only applicable to the kinetically expanded tubing within the upper tubesheets of the TMI-I steam generators. The inspection criteria and leakage assessment methodology described herein are not applicable to unexpanded tubing within the TMI-I upper tubesheets, or to the transitions between the unexpanded and kinetically-expanded tubing. (Other documents describe examinations of unexpanded tubing within the TMI-I upper tubesheets and disposition of those examination results. For example, TMI-l ECR TM 01-00328 (referenced in the plant's Technical Specification 4.19) describes examination requirements and acceptance criteria for unexpanded tubing within the TMI-1 upper tubesheets).

2.0 SYSTEM PERFORMANCE/KINETIC EXPANSION STRUCTURAL INTEGRITY ANALYSIS The design basis performance for the kinetically repaired TMI-I OTSG tubes is that, as a result of a Main Steam Line Break (MSLB), no tube shall break or separate from the tubesheet (Reference 27). In the following analysis, this performance requirement was practically applied as first, a condition that the tube is not permitted to part within the kinetically expanded joint (or at any other location). In addition, the repaired tube is expected to sustain a design basis axial load of 3140 lbs. with no slippage (Reference 28).

4 ECR # 02-01121, Rev. 2 l

For the kinetic expansion areas within the upper tubesheet, it is necessary to consider only the axial load applied through the tube to the joint as a result of the MSLB. The axial tube loads that occur during normal operations, for example those resulting from a normal cooldown transient, are much lower and will not exceed about 35% of the faulted condition. Since the kinetically expanded tubing is captured within the steam generator upper tubesheets, applied bending loads are very low in magnitude, and bending stresses do not develop within the joint because no rotation can occur.

MSLB is the design-basis accident for the kinetic expansions since it represents a hypothetical accident where tube stresses are relatively high, and the potential exists for offsite dose consequences from tube leakage resulting from significant primary-to-secondary pressure drop.

Other transients, such as the Small Break and Large Break Loss of Coolant Accidents (SBLOCA, LBLOCA) may result in relatively high tube stresses, but breaks of the primary system do not result in large primary-to-secondary pressure differentials. Primary-to-secondary pressure differential can be negative (i.e. pressure in the secondary system is greater than pressure in the primary system) during some LOCA events. FeedWater Line Breaks (FWLB) result in comparatively lower tube stresses than the MSLB and LOCA events.

As is described in more detail below, steam generator tubes will yield if subjected to significant axial loads. Tubes with lower yield strengths will begin to yield at lower loads than tubes with higher yield strengths. Tubes with a range of yield strengths are present in steam generator tube bundles. (For the TMI-I kinetic expansions, initial testing revealed that expansions formed with low-strength tubing were limiting in terms of joint pull strength and leakage.) Each of the MSLB, SBLOCA, and LBLOCA events, when conservatively analyzed, may impart axial tensile loads that cause steam generator tubes to begin to yield. Since the MSLB event has a relatively high primary-to-secondary pressure differential along with the relatively high axial loads, it is the design basis accident for the kinetic expansions.

Note that, at the time of this writing, AmerGen is working with the other B&W plant owners to revise BAW-2374, the LBLOCA topical, to address the LBLOCA transient for all aspects of the steam generators' design and maintenance. Final resolution of the kinetic expansions with respect to LBLOCA loads will follow the industry resolution of BAW-2374 issues.

2.1 Finite Element Modeling/Benchmarking In order to evaluate the behavior of kinetically-expanded joints with hypothetical flaws, and under the theoretical conditions where the tubesheet may bow, a finite element analysis model was developed in 1997. The analysis model of the tube-to-tubesheet joint consisted of a tube, the tubesheet, and a contact element representing the interference/connection between the tube and tubesheet. (Reference 24).

The analysis model had the additional feature that tube material behavior in both the elastic and plastic regions was modeled using actual tube stress-versus-strain data. Also, tube internal pressure could be included in the analysis model. Finally, the effect of tubesheet bow was captured. (The tubesheet may bow slightly due to the combined effects of axial tube load and primary-to-secondary pressure differences.)

The effect of the drilled holes in the tubesheet upon tubesheet stiffness was conservatively included in the finite element structural model so that the amount of tubesheet bow was not underestimated. The effect of the drilled holes on the tubesheet stiffness was directly modeled 5

ECR # 02-01121, Rev. 2 l

and was captured in the structural analysis. Several independent solutions were integrated in the present structural analysis. The actual bending stiffness of the tubesheet, including the drilled holes, was addressed in the original design analysis performed to determine tube minimum required wall thickness.

Maximum tubesheet displacement, under load, was identified using a finite element structural analysis model. For the purpose of calculating the kinetic expansion joint pull-out resistance, a conventional, closed form, solution for a solid plate was used to identify the displacements through-out the tubesheet based upon the maximum displacement obtained using the finite element solution for the drilled tubesheet. No error is introduced by this method with respect to computing tubesheet strain, which is the key variable in determining tubesheet hole dilation and constriction.

Test results available from the original 1980's kinetic expansion qualification program (Reference 29) were used as the basis for benchmarking the finite element analysis model results.

The benchmark process used qualification program tubes with high yield strength (57 ksi) and wall thickness slightly larger than design minimum tube wall (0.038" vs. 0.034"). [The resulting repair criteria assume that all tubes have minimum yield strengths and the minimum tube wall.]

Qualification test results were available for expansion lengths equal to four, six and eight inches.

High yield strength tube material was exclusively used for only the 4" and 8" expansions. Test results indicated that the joint's capacity to resist slip was the same for the 6" expansion as it was for the 8" expansion data.

The original 1980's qualification program's tube pullout test results at room temperature were used to benchmark the 1997 finite element model. More than eighty (80) tubes were pull tested.

The majority of these tests were performed at room temperature, which is conservative since the kinetic expansion joints are tightened with increasing temperature. (The Inconel-600 tube's coefficient of thermal expansion is greater than that of the alloy steel tubesheet.) As described in Reference 29, during the original qualification some pullout tests were also performed at elevated temperatures. These elevated temperature pullout tests confirmed the conservatism of testing joint interference at room temperature.

Surface condition variabilities were addressed during the original qualification of the kinetic expansion joints in the 1980's. (Reference 29) For example, pull testing on both uncorroded and corroded tubesheet blocks was performed. (The uncorroded blocks had lower pullout loads.)

All of the kinetically-expanded tubes in the TMI-I generators were kinetically expanded twice to ensure that the proper joint expansion was attained.

The 1997 finite element model was a conservative model that was based on the conservative testing and implementation of the kinetic expansion process implemented in the 1980's. For example, all repair criteria determined by the model assumed that the tubing had minimum wall thickness (0.034") and minimum yield strength (41 ksi). The 1980's testing confirmed that higher yield strength tubing resulted in a strongerjoint The 1997 model also neglected the effects of temperatures above room temperature upon the contact pressure of the joint. As described above, contact pressure (i.e., joint "tightness") increases with temperature because of the tubing's higher coefficient of expansion than that of the tubesheets.

The 1980's testing evaluated differences in diametral clearance between the tubing outside diameter and the tubesheet bore inside diameter before the kinetic expansions were installed.

(The design tolerances of the steam generator allowed this annulus diametral gap to be from 0.003" to 0.0 16".) The 1980's study concluded that the size of the annulus before the tubes were 6

ECR # 02-01121, Rev.2

expanded was insignificant with respect to the strength of the kinetic expansion joints.(Reference 24).

The finite element model parameters that describe the performance of the expansion are the contact interference between the tube and the tubesheet that was achieved by the kinetic expansion, and the coefficient of friction. Use of a contact interference dimension equal to 0.0003" in the model produced the best agreement with the joints' original qualification test results when using a coefficient of friction equal to 0.2. The analysis model results accurately matched the minimum test results obtained for the 4" expansion and underpredicted the performance of the 6" and 8" expansions. The same contact interference and coefficient of friction were used throughout the analysis reflecting the assumption that the kinetic expansion was equally effective over the range of expansion lengths. No parameter adjustments were made to produce results matching the pullout capacities for the 6" and 8" expansions as accurately as that obtained for the 4" expansion, to more accurately represent the shorter expansion. The analysis results are conservative for the longer expansions as a consequence of not adjusting the expansion parameters. For the conditions of the original testing, the pullout resistance of the 4" expansion is predicted by the model to be 3260 lbs. where the minimum test data result was 3100 lbs., 4030 lbs. for the 6" expansion where the minimum test data result was 5000 lbs., and 4110 lbs. for the 8" expansion where the minimum test data result was 5000 lbs.

The 1997 finite element analysis incorporated the possibility that tubesheet bore dilations from tubesheet bow during an MSLB could adversely affect the joints. No "bowing" of the tubesheet mockup blocks was measured during the original 1980's pull testing. Since no tubesheet bow was expected or factored into the original 1980's pull testing, finite element analysis was used to address tubesheet bow effects. Tubesheet bow, which reduces the assumed pullout resistance of the kinetic expansion joints, was incorporated into the finite element model after the model was benchmarked against the pull testing results. The finite element model was benchmarked against the 1980's pull tests results using common conditions; later, tubesheet bow effects were calculated. Reference 24 provides detailed information regarding the benchmarking process that was utilized.

Bowing or flexure of the tubesheet mockup blocks would not have been expected to occur during the conditions of these original tests. (In addition, it is conservative to assume that no dilations occurred during these original tests. The original tests determined the pullout resistance for joints of measured lengths. For example, a 6" long joint may have had a pull strength of 5000 lbs. Had dilations occurred, a reduced effective length of joint would have been present, since some length would have been affected by the bow. Therefore, the pullout resistance, in lbs. force per unit length of joint, would have been greater -since the effective length would have been reduced. Had tubesheet bore constrictions occurred during the original testing, the opposite result would have occurred -the joint strengths could have been increased. However, as described above, the original tests were room temperature tests with small tubesheet blocks in which no significant tubesheet bore dilation or bore constriction occurred.)

2.2 Finite Element Model Results The key performance features of the kinetically expanded joint are shown in Reference 24, which documented the finite element analysis. Figure 3-2 of Reference 24 shows the finite element analysis model results for a 6" expansion using high yield strength tube material. [The 6" expansion of the analysis model actually contained 5.5 inches of expanded tubing and a 0.5" expansion transition. The 0.5" transition does not contribute to the pullout strength of the kinetic expansion joint since the transition tubing is not in contact with the tubesheet. Actual 7

ECR # 02-01121, Rev. 2

profilometry data from a qualification test block indicated that a typical kinetic expansion has a transition of 0.5" length.] The residual contact pressure is shown in Figure 3-2 of Reference 24 as a function of distance above the transition region for both the condition of no applied load (dashed line) and the condition when slip begins (solid line). As described above, the effective length of the expansion is less than 6" because of the transition, which gradually tapers away from the tubesheet. (The analysis model ignored the fact that 17" and 22" long kinetic expansions were actually installed in the steam generator tubes.) Without applied load, the joint's residual contact pressure reaches a plateau a short distance away from the transition at a pressure equal to about 3300 psi.

The residual contact pressure abruptly decreases near the end of the expansion because of the effect of the free edge. The free edge is more flexible than the interior portion of the expansion so that the reaction at the edge is less for the same interference. The influence length of the effect of the free edge is determined by analysis to be approximately 0.25", which is reasonable in that this dimension is about three times the "decay length" of 0.08" based on widely used approximations of the structural influence of local discontinuities in thin tubes such as OTSG tubes (decay length = 0.78 it , where R is the tube inner radius and t is the tube minimum wall thickness). An axial flaw, like the end of an expansion, also changes the local stiffness of the tube, and a change in the local stiffness influences the contact pressure of the kinetic expansion joint. The influence of a flaw on contact pressure decays outboard of the physical dimensions of a flaw (as depicted in Figure 3-12 of Reference 24). This is evident from shell theory with respect to displacement and moment reactions due to local changes in stiffness. The edge effect extends more than 0.125" on each side of a flaw (i.e. 0.25" total influence), but partial joint contact pressure is maintained at the edges. (A step change from full contact pressure to zero contact pressure does not occur.)

Under slip load conditions, the model demonstrated that residual contact pressure redistributes due to Poisson contraction of the tube wall. The reduction of residual contact pressure is less with increasing distance above the transition. This is because the tube reaction decreases with increasing distance above the transition due to the increasing total contribution of the friction reaction. The pullout capacity of the joint is the product of the total residual contact pressure, the contact area, and the coefficient of friction.

The design basis MSLB load for the OTSG tubes of 3140 lbs. was determined by assuming that all tubes remain fully elastic (Reference 17). In order to create a conservative finite element analysis model it was necessary to adjust the model to reflect that many of the 1980's pull testing results were obtained using tubes of high yield strength and greater wall thickness (for consideration of the minimum yield strength and nominal wall thickness tubes that may be present in the steam generators).

The tubes in the OTSG having the lower bound yield strength (41 ksi per Reference 29) are expected to be in the plastic range for the design basis MSLB load. The 3140 lb. load corresponds to an axial membrane stress equal to 49.5 ksi and a design basis tube strain of 0.16%. A stress-strain curve for the lower bound yield strength material was developed by conservatively adjusting actual tube material stress-strain data from a TMI-I OTSG material heat. Using the design basis tube strain (0.16%) and the stress-strain curve for the lower bound yield strength material the maximum axial load that must be considered was 2400 lbs. The design basis load is caused almost entirely by an applied thermal displacement since the OTSG shell is at a higher temperature than the OTSG tubes after a MSLB. Using the site-specific stress/strain curve over both elastic and plastic stresses reduced the range of uncertainty in this analysis involving elastic/plastic material behavior.

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The analysis model results indicated very little increase in pullout capacity for expansion lengths greater than 4". This is because the low yield strength tubing begins to yield at a load equal to 2400 lbs. Poisson contraction of the tube wall relieves the contact interference between the tube and tubesheet, particularly after the tube begins to yield. As an axial load is applied to a tube, Poisson contraction begins to relieve contact interference, and hence decreases contact pressure, and proceeds further into the expansion in proportion to the load. The relief of contact pressure due to local yielding permits a higher applied load to reach further into the expansion because the benefit of the friction reaction is reduced at the beginning of the expansion as higher loads are applied. Local yielding occurs further into the expansion so that contact pressure is relieved there as well, and so on, so that ultimately there is very little additional capacity achieved for the 6" and 8" expansion with regards to the 4" expansion. This trend of results was reported during the original 1980's joint qualification program, and is also present in the Reference 24 analysis model. In short, there is decreasing utility in increasing the length of the joint above 4". The analysis model also showed a change in the performance of the joint from friction limited, when the intact expansion is at a minimum, to yield strength limited when the intact length is longer and the applied axial load is higher. This was an expected result, since the joints must yield as applied load is increased.

2.3 Flaw Dispositioning Criteria Development A flaw dispositioning criteria was analytically built, in part, on these performance features of the kinetically expanded joint. The analysis model was able to conservatively evaluate the performance of the intact and flawed kinetically expanded joints. For example, Reference 24, Section 3, Figure 3-12 shows the expected distribution of contact pressure in a 6" expansion [i.e.,

5.5" of expanded tube and a 0.5" transition] of a peripheral tube after a 2" 100% through-wall axial defect is introduced midway through the expansion length. The axial defect completely relieves contact pressure along its length and, in fact, influences the contact pressure for a length greater than 2" because of the "edge" effect as previously described. The expected pull out load for this configuration is 2509 lbs., which compares well with the capacity of the 4" expansion in a peripheral tube from Figure 3-11 (2516 lbs.) of Reference 24. Thus, a 2" axial defect in a nominal 6" expansion, without including tube internal pressure, forms an equivalent 4" expansion that also satisfies the qualification program criterion for resisting slip. The general conclusion from this and other similar calculations is that the kinetic expansions are flaw tolerant of axial defects (and for circumferential defects of limited extent also, as will be shown below) with respect to pull-out load. The required intact expansion for slip/pull-out load may be continuous or distributed in segments anywhere within the expansion length, provided the tube condition prevents tube parting.

The prescriptive conditions that were used to develop the design basis axial load for the MSLB include primary pressure equal to 2500 psi (Reference 17). Tube internal pressure should be included in the tube-to-tubesheet analysis model in order to identify the increase in contact pressure, in addition to residual contact pressure from formation, due to "pressure tightening".

As the internal pressure within the tube increases, the tube is tightened within the tubesheet.

When this pressure tightening was included in the analysis model, the analysis model results (Reference 24, Section 3, Figure 3-1 1) indicated that, for a lower bound yield strength tube having the design wall thickness, slightly less than a 2" expansion depth is required to resist pullout in a peripheral tube.

9 ECR #02-0112 1,Rev. 2

The finite element model conservatively assumed that contact pressure was completely released by the presence of a hypothetical flaw as a ring 360 degrees around the circumference of the tubing, and not only locally. This is a conservative treatment since compressive stresses are present in the expanded tubing, and the distribution of contact pressure around a flaw would actually follow the pattern expected for stress distribution of tension around a flaw in a plate.

(The far field conditions maintain a uniform tension while a stress concentration develops locally around the flaw.) This conservative treatment, of relaxed contact pressure around the full circumference of the tube, is implemented irrespective of the estimated depth of the flaw. So flaws estimated as, for example, 10% throughwall result in an assumption of a complete release of contact pressure over a full 360 degree "ring" of tubing. [As an example of the conservatism of this assumption, suppose that a kinetic expansion flaw is an ID-initiated volumetric flaw with eddy current estimated dimensions of 0.2" axial extent by 0.3" circumferential extent. An estimate of the area of this flaw is (0.2" times 0.3", or) 0.06 square inches. The area of expansion that is assumed to be released of all contact pressure is (pi times the tubing external diameter of 0.625" times the axial extent of the flaw, or) 0.39 square inches. So, for this example, the affected area is more than a factor of 6 times larger than the estimated area of the flaw.]

2.3.1 Required Length of Expansion The Reference 24 analysis model defined the maximum axial flaw length that could be present within a kinetic expansion and still meet the requirement to resist pullout (as a function of the radial location of the tubes.) Since the analysis model assumed that the flawed lengths of kinetic expansion do not contribute to the pullout capacity, subtracting the length of the maximum allowable flaw from the expansion length provided the minimum necessary length of defect-free expansion to resist pullout. Table 3-5 of Reference 24 provides results of analyses that were based on finite element modeling of a 6" expansion (5.5" kinetic expansion length plus a 0.5" expansion transition.) [Note 4 of that table states, "These criteria are only applicable for the fully-expanded region from 0.5" to 6" above the bottom of the kinetic-expansion joint." The length of the kinetic expansion transitions at the bottom of the kinetic expansions is approximately 0.5".] Table 3-5 provides "allowable defect lengths" within the 5.5" fully expanded length. For example, for a given tube location Table 3-5 may report that the allowable defect length is 4.4". Another way to state this is that a minimum of (5.5" minus 4.4", or) 1.1" of the kinetic expansion must be "defect free". In summary, the "required defect-free" lengths of the kinetic expansions, based on the finite element analysis, is the 5.5" modeled length of the kinetic expansions minus the calculated "allowable defect length".

For the 17" expansions, Table 3-5 of Reference 24 may be summarized as follows:

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Table A: Summary of Finite Element Modeling Results for 17" Expansions Column A Column B Column C Tube Bundle Location Allowable Axial Defect Minimum Required Length Kinetic Expansion Length

_ ( = 5.5" minus Col. B)

Periphery (Radial Location = 4.4" 1.1" 59.344") _

Mid-Radius (Radial Location = 3.2" 2.3" 42") _

Center .

(Radial Location = 2.8"- 2.7" 0.000") _

Note that Column B values in Table A above were plotted in Figure 3-20 of Reference 24. If we plot the Column C values, the minimum "required defect free lengths" of the kinetic expansions are depicted over the radius of the tube bundle:

Figure 19 Summary of Finite Element Results for 17" Expansions 3.5 4- 3 0

-J on0 2.5

._ 'r- 2 0 M 0 E

0 .: 1.5 E cc 1 0.5 0

0 10 20 30 40 50 Tubesheet radial Location (inches from center)

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The minimum required length of kinetic expansion, as described in the above paragraph, is based on finite element analysis only; the required expansion lengths were increased to conservatively account for field examination uncertainties. (Reference 24 determined structural requirements for the kinetic expansions based on structural analysis only and did not consider examination uncertainties.) For the inspection acceptance criteria additional length was added to the dimensions calculated in Reference 24 to conservatively account for the expected uncertainty in locating eddy current indications along the axial length of the kinetic expansion with respect to the expansion transition, and any uncertainty in locating the transition reference point itself.

When applied in the field the minimum "defect free" length is 2.1" for a peripheral tube. Table I provides the resulting list of minimum required lengths of defect-free expansion, AKELM,1N, for the various kinetic expansion lengths and their radial locations within the OTSG tube bundles.

Table I provides AKELmins that include the results of the finite element analysis plus additional length for conservatism and to account for possible examination errors.

The following plot (Figure 20) illustrates the conservatism of Table I with respect to the minimum defect free lengths for 17" expansions that were calculated by the model. Note that a fixed margin between the Table I value and the finite element modeling result was not used. A minimum of 0.5" margin was utilized. The minimum 0.5" margin was added to account for examination uncertainties.

FIGURE 20 Comparison of Table 1 Values to Model Results (17" Expansions) 3.5

- 3 2 .5--\ _,-Minimum Required Defect Free Length

0) (inches)

C 1.5 l Table I AKELMin Value (inches)

.;D 1 -.

0.5 0 10 20 30 40 50 Tubesheet Radial Location (inches from center) 12 ECR#02-01121,Rev. 2

The derivation of Table I values for the 22" expansions was performed in a similar manner to the above. However, the 22" long expansions are discussed in Section 2.6 of this document.

Figure I provides an illustration of a typical 17" deep kinetic expansion within the TMI-I upper tubesheet. As described above, TMI-I uses required kinetic expansion lengths that are conservative and are longer than those defined by the analysis model. TMI inspects and dispositions only these required expansion lengths. (Refer to Table 1.) A TMI-1 eddy current analyst reviews the tube's MRPC signal to locate the top of the kinetic expansion transition (i.e.,

that point where the tube is fully kinetically expanded against the tubesheet bore). This point is designated by the eddy current analyst as location ETL+0.00". (ETL = Expansion Transition Location) The analyst reviews the eddy current signals from the fully-expanded section; if no flaws are detected over the minimum required defect free length then the tube is dispositioned as "NDD" (i.e., No Detectable Degradation). If a flaw is detected, it is characterized, located with respect to the ETL+0.00" reference point, and additional kinetic expansion length is reviewed by the analyst to detect/characterize any other flaws that might be present. If the additional analyzed length contains flaws such that sufficient defect free tubing is not identified, the tube is repaired. If the additional kinetic expansion length is analyzed and sufficient defect free tubing length is identified, the expansion then may be left in service (provided it meets all other criteria to remain in service).

FIGURE 1 17" KINETIC EXPANSION (TYPICAL)

(NOT TO SCALE)

TUBE I z TOP OF UPPER TUBESHEET a z- w 2 -

. T BOTTOM OF UPPER TUBESHEET 13 ECR#02-01121,Rev.2 l

The kinetic expansion acceptance criteria apply only to tubing that has been fully kinetically expanded. As described above, the plant's analysis guidelines require that that point at which the tubing is fully expanded against the tubesheet bore is identified and is given the ETL + 0.00" reference point. This provides a reference point to locate any indications that may be present.

(See Figure I above.) All kinetic expansion examination results are referenced to the ETL+0.00" reference point. All minimum axial kinetic expansion lengths (AKELNUNS) are measured from the ETL+0.00" reference point.

2.3.2 Evaluation of Circumferential and Axial Indications (Note that, beginning in the IR16 refueling outage scheduled for the fall of 2005, AmerGen will plug all tubes with circumferential flaw indications in their kinetic expansion's required length upon detection-including circumferential flaw indications that were detected during prior 1997 through 2003 outage examinations that remain in service in the kinetic expansions' required lengths. As a result only ID Volumetric IGA indications will remain in service in the kinetic expansions' required lengths. This Section 2.3.2 remains in this document, however, to describe the treatment of the circumferential extents of volumetric indications. Volumetric indications have both axial and circumferential extents. That is, the axial extent of a volumetric indication is evaluated with respect to axial criteria, and the circumferential extent of a volumetric indication is evaluated with respect to the circumferential criteria.)

Evaluation of circumferential defects in the kinetic expansions was performed based on tube parting considerations. A tube may have a through-wall circumferential defect of 1300 (0.64", as measured on the ID) in extent and still have a sufficient ligament to resist the design axial load (36 percent of the tube circumference is permitted to be flawed). This evaluation assumed that the defect is located at the bottom of the expansion region where the axial force is at its maximum. (At higher elevations within the expansion region, part of the axial force would be transmitted to the tubesheet by the friction restraining force, thereby reducing the axial force in the tube wall. As a result, the allowable circumferential defect in higher areas of the expansion region would be greater than 0.64".)

For multiple circumferential defects in the expansion region, the allowable combined length of the defects would be 0.64" if the elevation difference is less than a separation criterion. These separation criteria were conservatively evaluated as part of the analytical work. The resulting flaw combination criteria are based on providing the required shear path between defect elevations in order to transfer the total load. It is conservative to include total load for shear transfer since membrane transfer also occurs. A reasonable separation distance was judged to be I" considering that 1.13" of intact tube length is required at the plane of the defect for membrane stress. A I" separation provides 2" of shear transfer path (I" at each side of a defect) at an allowable stress of 60% of that for membrane stress. For example, if two circumferential defects are separated by an axial distance greater than I", each one may not exceed 0.64" in length.

These criteria will ensure that the tube within the expanded region will not part.

The I-inch separation distance represents the required shear path to transfer the axial load applied to the joint from the elevation of a circumferential flaw to the next elevation of another circumferential flaw. Flaws with separation distance greater than I inch do not interact. This separation criterion for combining the effective length of nearby circumferential flaws is based on the ASME Section III method for determining the maximum allowable average shear stress.

The externally applied axial load is assumed to be reacted in shear by the ligament separating 14 ECR #02-0112 1,Rev. 2 I

flaws at different elevations. If the required distance is not satisfied, then it is concluded that the flaws interact.

The I-inch separation distance between two flaws assumes that the tube is freespan. No credit was taken for the expansion. In summary, the I" separation distance over which flaws are combined was determined based on the freespan condition; no credit was taken for the presence of the tubesheet, or compressive stresses that are present from the tube-to-tubesheet expansion joint. This is a conservative practice because all of the axial load is assumed to be reacted in shear disregarding the portion of the load that is actually reacted by the friction force developed by contact pressure associated with the ligament between the flaws. The actual load reaching the flaw elevation is also assumed to be the maximum disregarding the reduction in load as a function of depth into the kinetic expansion joint.

The "edge" effect, described in Section 2.2, is an additional factor that must be included when evaluating the impact of circumferential defects. The edge effect of a circumferential defect degrades the pullout capacity of the tube much like an axial defect, as discussed above. For purposes of developing a flaw dispositioning criteria, a 0.25" axial influence will be added to each circumferential defect. In this way, the results for the contact pressure redistribution in the presence of only an axial defect form the basis for the comprehensive dispositioning criteria with respect to pullout resistance.

The resulting inspection acceptance criteria for the OTSG kinetic expansion region are given in Table I and Table 2. Note that criteria differ for periphery, mid-bundle, and center tubes due to the effect of tubesheet bow, to be described below. As a result, the Table I and Table 2 values for a given tube are a function of the radial location of that tube within its OTSG tube bundle.

Table 2 provides an example to clarify how the edge effect is applied.

Note also that Table 2 requires that the 0.25" axial influence "edge" effect be added to the axial length of each axial defect, except the first defect. This exception is present because the finite element model's calculation of the minimum required expansion lengths assumed one defect was present, including that defect's edge effect. For multiple defects (i.e., the second, third, and so on), the axial length of each additional defect is considered and the additional edge effect is added. A 0.25" axial edge effect is assumed for all circumferential defects.

Note that the "edge" effect and the I-inch separation distance are two different, and independent, parameters: The edge effect is a dimension (i.e. 0.25") that reflects the length of kinetic expansion tube-to-tubesheet joint that might be adversely affected due to the presence of an individual flaw. The 1-inch separation distance is the length of tube required between two adjacent flaws beyond which the two flaws will not interact. Edge effects are considered for all flaws; the separation distance is only applicable when considering the proximity of two (or more) flaws. Table 2 describes how the 1-inch separation criterion and edge effects are implemented.

(Since the separation criterion was calculated assuming the tube was in a freespan condition, the edge effect 0.25" value and the separation criterion I" value are not added to create a new 1.25" separation criterion. The I" separation criterion is independent of the tube-to-tubesheet contact pressure and is a function only of the tubing material shear strength. If credit were taken for any additional benefit of tube-to-tubesheet contact pressure between the elevations of adjacent flaws, the required separation distance would be less than I inch.)

The loads, methods, and assumptions that were used in the analysis are conservative. The 3140 lb. axial load that was used to develop the inspection criteria is from a conservative analysis based on conservative assumptions with respect to TMI-1 regarding main steam line size, and 15 ECR # 02-01121, Rev. 2

maximum emergency feedwater flow and duration. For example, more recent analyses addressing expected MSLB thermal/hydraulic conditions and tube loads (described in Section 5.0) indicate that the maximum axial tube load is about 1300 lbs. as opposed to 3140 lbs. Thus, the use of the axial tube load from the analysis in the development of the inspection criteria incorporates a conservative factor of at least 1.8 with respect to the maximum axial tube load for the lower bound yield strength material, i.e., 2400 lbs.

Each kinetic expansion defect was assumed to locally relieve the tube-to-tubesheet contact pressure to the same extent as a 3600 cut regardless of its circumferential extent. Therefore, the relief of contact pressure due to any acceptable circumferential defect is overestimated and actual pull-out capacity is higher than that calculated. In addition, with regard to acceptable circumferential defect location, no credit is taken for the reduction in applied axial tube load within the expansion due to friction. The assumption provides more conservative results for defects that are further within the expanded zone, (i.e., the full axial load is assumed to be imparted on a circumferential defect, regardless of its location within the expansion). These structural analyses of joint integrity assumed that all defects are 100% through-wall. Any difference between actual depth and the assumed 100% through-wall depth of the analysis model represents an additional conservatism.

2.4 Fatigue Analysis The analysis of the joints also evaluated the possibility that defects that are acceptable for the faulted condition could propagate by fatigue during normal operation. The important contribution to propagation by fatigue is the axial tube load due to the cooldown, because bending stress, such as that due to flow induced vibration or due to local bending at the elevation of a defect, does not occur in the expanded tube above the transition. Crack propagation by fatigue was conservatively evaluated previously during the repair of the OTSGs considering a defect located in the free span. The previous calculation was useful for guidance because, while it did not identically match the kinetic expansion condition, it was representative. The previous calculation considered a smaller through-wall, circumferential defect (0.36" circumferential extent), but also included local bending stress. The sum of these is practically the same as the membrane stress for the kinetic expansion analysis (i.e., the kinetic expansion analysis had a longer defect and no bending stress). The results indicated that, on a per cooldown cycle basis, the expected crack propagation is about 104 inches in circumferential extent per cycle. For example, assuming six cooldown cycles per year for two years of operation, propagation by fatigue results in practically no increase in circumferential extent. It is, therefore, not necessary to reduce the extent of the acceptable critical defect size in the expanded tube because of expected propagation due to fatigue during the forthcoming operation cycle. In addition, re-inspection of representative ID volumetric indications left in service in the kinetic expansions will take place during subsequent refueling outages in order to verify that flaw extent is not increasing to unacceptable size. (Additional discussion regarding the possibility of growth of existing flaws in the kinetic expansions is provided below.)

2.5 Tubesheet "Bow" Analysis The analysis model (and the resulting inspection criteria) for the OTSG tubes includes an additional feature of the performance of the joint: tubesheet bow (due to tube axial load and due to primary-to-secondary pressure differences during an MSLB) is assumed to open the tubesheet bores below the tubesheet center plane and close them above. The tubesheet bore dimension was adjusted in the analysis model to reflect the expected bending strain distribution at the elevations of the expansion due to tubesheet bowing. The effect is greatest for a center tube where bowing 16 ECR # 02-01121, Rev. 2

is maximum. There is no effect for a peripheral tube. As a result of the upper tubesheet bowing inward, the applied axial tube load on the affected tubing is reduced, with the minimum occurring at the center. However, as another result of tubesheet bow, the contact pressure of the tube-to-tubesheet joint is reduced due to enlargement of the tubesheet hole in the area of the joint. This effect is greatest at the secondary face of the tubesheet.

The greatest impact of tubesheet bowing is for the 22" deep expansions where the original 6" qualification length was further below the tubesheet center plane, and closer to the secondary face, than for the 17" expansions. In fact, for a 22" expansion at the center, tubesheet bow eliminates most of the residual contact pressure even when considering tube internal pressure.

(The effects of tubesheet bow were not evaluated during the original kinetic expansion qualification program of the early 1980's.)

The kinetic expansion inspection criteria identify the minimum required defect-free kinetically expanded tube length that must be present within the inspected distance (Table 1)as well as the flaw, or combination of flaws, allowable within the inspected distance (Table 2). The inspection may continue beyond the nominal qualification length, if necessary, in order to demonstrate the presence of a satisfactory joint since the tubes were kinetically expanded over the entire length of the tubesheet above their original 6" qualification length. The absence of consideration of the effects of tubesheet bow as part of the original qualification program will not impact nuclear safety as long as the 22" expansions within the center and mid-radius locations of the tubesheet are inspected to the same elevation as the 17" expansions and evaluated to similar criteria. (Note that the lower 5" of the center and mid-radius 22" kinetic expansions [from ETL + 0.00" to ETL+5.00"] are also evaluated as freespan tubing, as is discussed below.)

2.6 Implementation of the Inspection and Repair Criteria The inspection of a kinetic expansion always includes a concurrent inspection of its transition.

(This is required by the plant's eddy current guidelines and is also necessary to determine the location of the ETL+0.00" reference point as described above.) All kinetic expansion examination results are referenced to the ETL+0.00" reference point at the top of the expansion transition. All AKEL,,,N minimum axial kinetic expansion lengths (for both 17" and 22" expansions) are measured from the ETL+0.00" reference point. Section 4.0, which follows, provides details regarding the eddy current inspection of the kinetic expansions.

Volumetric indications are dispositioned by combining the results that were derived separately for axial and circumferential defects. That is, the criteria for axial defects shall be used for the axial extent of the volumetric indication and the criteria for circumferential defects shall be used for the measured circumferential extent of the volumetric indication. (The majority of TMI-1 OTSG kinetic expansion flaws are volumetric ID IGA indications, similar to those found in the freespan tubing of the TMI-I generators.)

As is apparent in Table I and Table 2, field implementation of these inspection criteria is specific with respect to both tube location and expansion length. The analysis model determined allowable defect sizes (plus influences) as a function of relative radius of the tube bundle for 17" and 22" expansions. The analysis model calculated values at specific radial locations; it is conservative to apply results specifically for tubes that are located at a smaller radius as governing for tubes located at a larger radius. This logic represents an additional factor that contributes to the conservatism of the inspection criteria.

17 ECR # 02-01121, Rev. 2

As is also apparent in Tables I and 2, disposition of defects in the 22" expansions is notably different than for 17" expansions. (The majority of the TMI- I kinetic expansions are 17" in length. Of 31,062 tubes in the TMI-1 steam generators only 431 tubes have 22" kinetic expansions that remain in service during the plant's current operating Cycle 15.) The lower 5" length of the 22" expansion at center and mid-radius locations does not contribute to slip resistance under postulated MSLB conditions due to the tubesheet bowing. For this reason the required defect-free expansion lengths (AKEL1,NI) for the 22" expansions located near the center of the tube bundle are 5" longer than that for 17" expansions located at the same tube bundle radial position. Indications in the lower 5" length of the 22" expansions located near the center of the tube bundle are dispositioned using more stringent free span criteria, since this length of expanded tubing loses contact with the tubesheet as a result of postulated tubesheet bow.

Amendment #237 to the TMI-1 Technical Specifications incorporated a requirement to implement the freespan tubing acceptance criteria for volumetric ID IGA indications within the lower 5" of the 22" long expansions at the center of the tube bundles. Amendment 237 also implemented a requirement that 100% of the 22" long expansions at the center of the tube bundles be examined during each tubing inspection. In summary, the lower 5" of the 22" long expansions at center and mid-radius locations are a special subset in which both the freespan inspection acceptance criteria and the kinetic expansion acceptance criteria are applicable. (The freespan acceptance criteria are more stringent than the kinetic expansion criteria.)

To derive the Table I values for the 22" long kinetic expansions in the periphery of the tube bundles, Figure 3-20 of Reference 24 was used in a manner similar to that described for the 17" long expansions (-as described in Section 2.3.1, above). The following plot illustrates the conservatism of Table I with respect to the minimum defect free lengths for 22" expansions that were calculated by the model. Note that a fixed margin between the Table I value and the finite element modeling result was not used. A minimum of 0.5" margin was utilized.

18 ECR # 02-01121, Rev.2

FIGURE 21 Comparison of Table 1 Values to Model Results (22" Expansions) 6 (A

en W

0

-Minimum Required Defect am Free Length (inches) r

-* -Table 1 AKELMin Value (inches)

C:

0)

S CD

-j qa, A,

U a,

At

"-A-a, W

42 44 46 48 50 52 54 56 58 Tubesheet Radial Location (inches from center) 2.7 Sleeved Tubes There are 502 tubes in the TMI-1 steam generators with kinetic expansions that have sleeves installed. These tubes were sleeved during the plant's 1991 and 1993 outages as a preventive measure to mitigate primary-to-secondary leakage resulting from high cycle fatigue cracks. (All of the in-service OTSG plants installed sleeves to prevent these fatigue cracks.) Each of the TMI- I sleeves was manufactured from Inconel 690 material and completely spans its tube's kinetic expansion to form a new pressure boundary. The B&W rolled sleeves installed in the TMI-l steam generators extend from the primary face of the upper tubesheet to a point more than 80 inches down into the steam generator tube (i.e., deeper into the tube than the 17" or 22" deep kinetic expansions). Figure 22 depicts a typical TMI-I steam generator tube sleeve.

19 ECR#02-01121, Rev. 2

Any tubes with flaws detected in the sleeves, or in the parent tube adjacent to the sleeve between the lower sleeve end and the parent tube kinetic expansion transition, will be "plugged-on-detection".

Given that the kinetic expansions in TMI-1 sleeved tubes have been removed from service, kinetic expansion examinations are not conducted in these sleeved tubes and the subject inspection and dispositioning criteria are not implemented in those tubes.

20 ECR#02-01121,Rev.2 l

FIGURE 22 B&W ROLLED SLEEVE Not to scale. Dimensions are nominal and provided in inches.

21 ECR #02-0112 1,Rev. 2 1

3.0 MATERIAL CONDITION ASSESSMENT 3.1 Stress Corrosion Cracking Mitigation Resulting From Kinetic Expansion The impact of kinetic expansion on the TMI-I OTSG tube material condition can be considered in two separate parts. First, there is the effect on pre-expansion defects. Secondly, there is the formation of, and benefits from, post-expansion residual stresses. Kinetic expansion is not a corrosive process; rather it is a mechanical, cold work process that produces plastic strain. It is reasonable to assume that defects that may not have been initially detectable may have been enlarged by the kinetic expansion and, thereby, made more detectable. It is possible, particularly for the axial component of defects because of the induced permanent circumferential strains, that defect dimensions increased due to the expansion, that the distance between defect planes (i.e.,

crack opening displacement) increased, that grain drop-out increased the defect volume, or that a combination of these changes occurred. As a result of the effects described above, kinetic expansion probably enhanced flaw detection. In addition, the eddy current techniques used to examine the kinetic expansions during recent refueling outages are more sensitive than the techniques used during the 1980's, when the expansions were created.

No defect growth has been observed over the course of recent operating cycles for kinetic expansion defects that have been reviewed with the same ECT technology.

3.2 Growth Monitoring and Examination Scope TMI-I has monitored the growth of eddy current indications within the kinetic expansions for the past several outages (since MRPC inspections were started) and has reported these results to the NRC (References 32, 35, 36, 37). Since the original 1997 submittals regarding the kinetic expansions (i.e., References 25 and 26) TMI-l has provided additional details regarding growth of indications in the TMI- I steam generators. Reference 30 provided information regarding the methods with which TMI-I has monitored the growth of the ID degradation found in the kinetic expansions, and as well as growth within the unexpanded tubing. Indications have been evaluated for changes in axial extent and circumferential extent over successive outages, and over multiple outages. Analysis of indication growth, and an assessment of that indication growth relative to the repair criteria, is required by the plant as part of operational assessments each outage.

Reference 30 provided information regarding the reliability of ECT techniques used for indication detection and sizing. TMI-1 has examined all of the population of inservice kinetic expansions, by examining approximately one third of the tube population during each of the last four plant refueling outages (Outages 12R, 13R, IR14, and IR15).

TMI-1 will continue to monitor for growth of flaws in its steam generators, including flaws in the unexpanded tubing within the tubesheets and kinetic expansions. The following parameters are compared for kinetic expansion flaw indications:

-change in axial extent of ID volumetric indications

-change in circumferential extent of ID volumetric indications

-whether or not new indications are detected 22 ECR # 02-01121, Rev. 2

The results of these evaluations are evaluated in several ways: "scatter plots" of the data are created to visualize the trend of the data. Average changes, standard deviations of the changes, and maximum changes are calculated and reviewed. During the last TMI-I outage, sign and paired-t statistical tests were performed on the ID volumetric IGA indications in the kinetic expansion region, as are performed for the ID volumetric IGA indications in the freespan tubing.

(Refer to Sections 3.2.1.5 and 3.2.1.6 of this report.)

Statistical analysis of the growth results is necessary. The kinetic expansion indications are relatively small in size, and the variability of the eddy current examination process must be considered to evaluate the population for growth. For example, the following (Figure 23) is a "scatter plot" that depicts the change in axial and circumferential extents (in inches) of the ID volumetric indications in the "A" steam generator kinetic expansions over the last operating cycle. This plot illustrates some of the variability of the growth data.

FIGURE 23 Growth of ID Kinetic Expansion VOLs TMI-1 SG A 0.6-0.5 x 1R14to1R15DeftaExtent

- Extreme Value Box 0.4-0.3

-J U

U 02 l U

0.1 I x xI l 4'0 xx,

.0.1

-0.2

-0.3

-0.3 40.2 -0.1 0 0.1 02 0.3 0.4 0.5 0.6 Measured Delta Axial Length In addition to monitoring for the growth of existing flaw indications in the kinetic expansions, TMI-I also evaluates the population of new flaw indications identified during the examinations.

Section 3.2.1, which follows, provides a technique with which new flaw indications will be evaluated during each examination.

The following tables (Tables D and E) provide the results of evaluations of changes in kinetic expansion indications from the plant's most recent Outage IRI5.

23 ECR # 02-01121, Rev. 2 l

Table D Historical Growth Summary for Kinetic Expansion Volumetric IDIGA Indications Average Change in Average Change in Operating Period SGA [

Circ. Extent SGB SGA Axial Extent SGB IR14 - IR15 0.003" -0.005" 0.005" -0.008" 13R - IR15 -0.010" -0.024" -0.008" -0.034" 12R- IR15 -0.013" -0.049" -0.001", -0.034" Table E Growth of Kinetic Expansion Circumferential Indications Average Change in Operating PeidSGA J Circ. Extent SGB 1R14 - 1RI5 +0.010" -0.005" 13R - iRIS -0.009" -0.046" From 12R to IRI 5, the average change in circumferential extent of kinetic expansion circumferential indications (within the tubes' minimum required expansion length, AKELmin) in both steam generators was -0.02".

In order to monitor the kinetic expansions for the possible onset of new degradation, and monitor the existing degradation for possible growth, examination of the tubing is required each outage.

Specifically, beginning with the IR16 Refueling Outage planned for the fall of 2005, all of the non-plugged, non-sleeved tubes' kinetic expansions and their kinetic expansion transitions will be scheduled for inspection with rotating coil eddy current probes during each refueling outage.

The TMI Unit I plant operating cycle length is presently 24 months.)

Approximately one third of the plant's kinetic expansions have been examined during each of the plant's last four refueling outages (i.e., 1997 through 2003 refueling outages). These samples were sufficient to detect whether significant growth of existing flaws in the kinetic expansions was occurring, or if any new degradation began to appear within the kinetic expansions. The plant's steam generator program requires that condition monitoring assessments and operational assessments be performed based on the results of the outage examinations. The operational assessments must contain an evaluation of the potential for growth during the following operating cycle.

To conservatively address the appearance of new kinetic expansion flaw indications, TMI-I will plug/repair any tubes having kinetic expansions with new flaw indications in their required expansion length that were not detected during the 1997 through 2001 refueling outage examinations. "Lookbacks" will be used to evaluate whether or not an indication may have been present in this previous outage data. (Each of the in-service kinetic expansions was first examined with an MRPC probe during the 1997, 1999, or 2001 outage.)

24 ECR # 02-01121, Rev. 2 l

3.2.1 Procedure for Monitoring Growth of Kinetic Expansion Indications 3.2.1.1 Introduction This section provides the procedure for continued growth monitoring of the indications within the kinetic expansions of the TMI-1 steam generators. [Note that much of this procedure is nearly identical to the growth monitoring procedures already incorporated into the TMI-I Technical Specifications for monitoring of ID Volumetric IGA indications in the unexpanded tubing per ECR TM 01-00328. Approximately 80% of the indications in the kinetic expansions are ID Volumetric IGA, so the techniques used to monitor growth in the unexpanded tubing are also applicable to the majority of the kinetic expansion indications. (At the close of the 2005 refueling outage examinations, the only flaw indications remaining in service in the kinetic expansions' required lengths will be ID Volumetric IGA.) However, since the bobbin coil probe is not used in the kinetic expansion area, it was necessary to revise the growth monitoring procedures outlined in ECR TM 01-00328.] The procedure is a multi-step process including statistical tests to detect changes in the apparent growth distributions.

Eddy current indications found within the TMI-I kinetic expansions during recent refueling outages have been of two types: ID-initiated volumetric IGA and circumferential indications.

Thus, the appearance of new OD-initiated indications, axial indications, or other type of degradation differing from the aforementioned two types would be evidence of a new form of degradation. Beginning in the IRI6 refueling outage scheduled for the fall of 2005, AmerGen will plug all tubes with circumferential flaw indications in their kinetic expansion's required length upon detection-including circumferential flaw indications that were detected during prior 1997 through 2003 outage examinations that remain in service in the kinetic expansions' required lengths.

The procedure for growth evaluation of the kinetic expansion ID Volumetric IGA indications consists of screening the data for extreme values, followed by two statistical tests that will be applied to axial and circumferential length measurements from the kinetic expansion ID IGA inspection data. The two tests will be the application of a sign test and a paired t-test. These two tests will be applied to each of the two variables. If all tests are passed (that is, if all four tests demonstrate that the ID IGA growth rate is less than a small positive value), it will be concluded that the kinetic expansion ID IGA population is not growing. If these tests are unsuccessful in demonstrating that growth is less than a small positive value, a cycle-specific growth model and NRC notification are required.

3.2.1.2 Capability of Statistical Tests to Detect a Change in Mean Growth of ID Volumetric IGA Indications in the Kinetic Expansions Increases in measured eddy current parameters do not necessarily indicate actual growth of flaws as they also reflect the NDE uncertainties associated with sizing relatively small flaws (as discussed earlier in this report). The validity of classical statistical tests for no growth depends strongly on the assumption that the data are normally distributed. Departures from normality such as excessive peakedness or skewness affect the results of the tests and may lead to incorrect conclusions (for example, concluding that flaw dimensions have changed when, in fact, they have not). The methods described in this report may be applied even if the data does not have a 25 ECR # 02-01121, Rev. 2

normal distribution. Generally, large datasets tend to be normally distributed; so the risk of error is not substantial.

3.2.1.3 Procedure for Assessing ID IGA Growth in the Kinetic Expansions As described above, a statistical procedure will be used to assess ID IGA growth. The procedure consists of initial screening of the data for extreme values, followed by two statistical tests that will be applied to axial and circumferential length measurements from the kinetic expansion ID IGA inspection data:

(l) Sign test (2) Paired t-Test These two tests will be applied to each of the two variables (i.e., axial and circumferential extent) for a total of 4 tests. If all tests are passed (that is, if all test statistics calculated from the ID IGA growth data are statistically insignificant), it will be concluded that the kinetic expansion ID IGA population is not growing.

If the test results are unsuccessful, then some evidence exists in the apparent growth data that the population of kinetic expansion ID IGA indications may have changed. At this point it is necessary to develop a cycle-specific growth model that should be applied in the operational assessment.

An outline of the procedure follows:

Perform Extreme Value Screening and Perform Statistical Tests for Change in the Kinetic Expansion ID IGA Flaw Population:

1.Extreme Value Screening

2. Sign Test

3. Paired t-Test Because of the limited data population in the "B" OTSG, data from the two steam generators will be combined for these tests. (The majority of the kinetic expansion indications are in the "A" OTSG.)

Data from individual indications will be compared back to the first outage with acceptable MRPC data for these statistical tests.

3.2.1.4 Step Ia. Extreme Value Tests for Largest Growth Rates An extreme value analysis will be used as an initial screening for kinetic expansion volumetric ID IGA indications that may be outliers in the datasets. For example, if an indication is mis-analyzed or mis-characterized as volumetric ID IGA (in either the current outage or a previous outage), the extreme value screening will help identify the indication. Similarly, if an indication were to grow or "shrink" by a large amount, this test will help to identify it. The extreme value screening serves to identify (mathematically) those indications that might also be found by visual inspection of a scatter diagram of the data for outliers.

26 ECR# 02-01121, Rev. 2

Samples from normal distributions yield extreme (in this case maximum apparent growth) values that are described (for large sample sizes) by the so-called Type I Extreme Value distribution.

Since the number of volumetric ID IGA flaws in the kinetic expansions of the TMI-I steam generators is relatively large, the Type I distribution is expected to provide a good representation of the expected frequency of extreme growth values. This screening is performed by comparing the largest observed growth value with the 5% critical value. If the largest growth value is less than the critical value, it will be concluded that the IGA growth data extreme value is not statistically significant.

If the extreme value screening identifies indications with erroneous data, the erroneous data will be corrected prior to using that data in the subsequent screenings, or subsequent Sign and Paired t statistical tests. If the extreme value screening identifies indications with large apparent growth rates, and are not due to erroneous results, these indications will be used in the subsequent statistical tests.

In summary, the extreme value analysis will be performed to identify possible outliers or erroneous data.

3.2.1.5 Step lb. Perform Sign Tests for Change in Kinetic Expansion ID IGA Population The Sign test is a statistical test for detecting differences in the median of a binomial distribution from a reference value. This test will be used to identify the presence of statistically significant (i.e., positive) change in the kinetic expansion ID IGA flaws based on two eddy current measurements: measurements of axial and circumferential lengths. This approach will not require that the data be normally distributed.

The Sign tests will determine if the growth of the kinetic expansion ID IGA indications is bounded by the following small, positive reference values between examinations: 0.01" axial extent increase and 0.01" circumferential extent increase. (The use of small positive values will reduce the possibility that random process error alone could result in mistakenly concluding that actual physical growth has occurred. These small extent values are very small in comparison to the repair criteria.) The maximum Type I error (i.e., the probability of erroneously concluding that there is growth when there is actually no growth) is 5%.

The variables for the Sign tests are:

a = the significance level of the test = 0.05 for a one sided test O = the standard = inches (for axial or circumferential length)

Xi = each observation (change in inspection parameter for each indication) for a given parameter, from 1 to nocal n,,t,, = the total number of indications for which there is data or observations for a given parameter X = average of Xi 27 ECR # 02-01121, Rev. 2 l

r = the number of observations less than the standard rcrj, = critical value of "r" for the sign test which is taken from Table A-33 in Reference 39 Note that the significance level of the test has been chosen to be equal to 0.05 which is a generally accepted value within industry. The significance level of the test, as well as the number of observations, affects the probability of making a correct determination.

If r is greater than rat, it is concluded that there is no reason to believe the measured parameter change is different from zero and therefore, there is no reason to believe the defects were growing in the given outage interval.

3.2.1.6 Step Ic. Perform Paired t-Tests for Change in Kinetic Expansion ID IGA Population The Paired t-test is a standard statistical test for hypothesis testing as regards the significance of differences in sample means. The standard paired t-test will be used to further evaluate whether growth is indicated by this parametric test. For this application, again the null hypothesis is that the mean change (growth) in the kinetic expansion ID IGA flaws is bounded by the following small, positive reference values between examinations: 0.01" axial extent and 0.01" circumferential extent. (As in Step Ib, the use of small positive values will reduce the possibility that random process error alone could result in mistakenly concluding that actual physical growth has occurred.)

a = the significance level of the test = 0.05 for a one sided test mn0 = thestandard = inches(foraxialorcircumferentiallength)

X; = each observation (change in inspection parameter for each indications) for a given parameter, X = average of Xi n = the total number of defects for which there is data or observations for a given parameter u = difference between the observed average and the standard = X - mn 0 ti ,= percentile of the t distribution, taken from Table A-4 of Reference 39, as a function of level of significance, a and degrees of freedom, df S

U1crit = t] a r s = standard deviation df = degrees of freedom = n- l 28 ECR# 02-01121, Rev. 2

If u is less than ucn it is concluded that there is no reason to believe the measured parameter change is different from zero and there is no reason to believe the defects were growing in the given outage interval. If u is greater than un,,, it is concluded that the defects were growing in the given outage interval.

Sign and Paired t-testing were performed in accordance with the above procedure on both the axial and circumferential extent changes of 434 volumetric IDIGA indications found in the TMI-I kinetic expansions during the plant's IR15 and IR14 Outages. The results of the tests supported the "no growth" assumption (i.e., no reason to believe that growth had occurred).

3.2.1.7 (This section, and Step II, were deleted.)

3.2.1.8 (Deleted) 3.2.1.9 Step III. Evaluate the Number of "new" Kinetic Expansion Indications Identification of new indications is expected as analyst sensitivity and technique sensitivity (i.e.,

data quality) change. To conservatively address the appearance of new kinetic expansion flaw indications, TMI-I will plug/repair tubes having kinetic expansions with new flaw indications in their required expansion length that were not detected during the 1997 through 2001 refueling outage examinations. "Lookbacks" will be used to evaluate whether or not an indication may have been present in this previous outage data.

As described above, detection of OD indications or axially-oriented indications will also be indicative of a new form of degradation, since these types of degradation are not normally found in the kinetic expansions. The results of this analysis will be provided to the NRC in the "90-day report" currently required by the plant's Technical Specifications 4.19.

3.2.1.10 Step IV. Develop Cycle Specific Growth Model If Steps I through III, above, are successful in demonstrating the lack of statistically significant growth in the kinetic expansion eddy current indication population, Step IV is not necessary.

However, in the event that future TMI-I kinetic expansion field data indicates that growth is greater than a small positive value change from the historical population, or apparent growth as evidenced by the inability to demonstrate statistically insignificant growth via the procedures in Steps I through III, it will be necessary to develop a cycle-specific model of growth. This growth model will characterize changes in the mean, variability and extremes of apparent growth and will be important as a basis for a cycle-specific growth allowance to be used in operational assessments for forthcoming cycles.

After using the procedures in Steps I through III, TMI-l will notify the NRC during any outage in which growth is greater than a small positive value change from the historical population, or apparent growth as evidenced by the inability to demonstrate statistically insignificant growth.

(Refer to Section 5.9 regarding reporting methods.)

It may be necessary to re-verify the analyst-to-analyst variability that is applicable to the field data at hand and to evaluate the components of variability so that an accurate model of actual growth can be obtained. Any growth analysis performed using the cycle specific growth model 29 ECR#02-01121,Rev.2 l

described here will require a revision to this report to include information substantiating the growth conclusions reached and the basis for the conclusions. The revised report will be submitted to the NRC well ahead of the subsequent refueling outage with any actions to address potential growth.

3.3 Residual Stresses Kinetic expansion produces residual compression in both the circumferential and axial directions. This can be understood by considering the mechanics of the process. The residual contact pressure from formation is an external pressure on the tube OD due to the interference between the tube and tubesheet. The resulting residual hoop stress in the tube is compressive at a level approaching the yield strength of the tube material. In addition, during the expansion, as contact between the tube and tubesheet increases, the tube is extruded against the friction that is also developing in the contact zone. The friction reaction due to contact pressure causes residual axial compression by resisting extrusion.

Service conditions will not completely remove residual compression of the kinetic expansions in either the circumferential or axial directions. At operating conditions, increases in both internal pressure and temperature cause an increase in contact interference, resulting in higher compressive circumferential stresses. Axial tube loads applied during normal operation will not remove residual axial compression completely because contact pressure due to radial interference is not lost. Axial load on the joint is at a maximum during the normal cooldown transient but will not exceed about one-third of the applied axial load during the faulted condition. The normal cooldown transient will not remove contact interference even for the limiting 22" expansion at the tubesheet center. Any reduction in axial compression is temporary with full elastic restoration following any (and all) cooldown transient(s).

The kinetic expansion joints, under normal operation, have compressive residual stresses in both the axial and circumferential directions. Mitigation of stress corrosion cracking, both for new damage and propagation of existing damage, is accomplished by maintaining these compressive residual stresses within the kinetically expanded regions. Since the analytical model and structural repair criteria assume that all defects are 100% through-wall, and that circumferential defects result in a full relaxation of the tube-to-tubesheet contact pressure over 3600 of the tube circumference, there exists substantial allowance for flaw growth. In addition, the MRPC eddy current techniques provide conservative measurements of flaw extents within the kinetic expansions. (See Section 4.0, which follows). With these conservatisms and the other conservatisms of the finite element model, the as-called eddy current indication length and widths are evaluated with respect to the repair criteria. Additional factors or increments to account for flaw growth are not used and are not necessary.

4.0 BASIS FOR DISPOSITION OF INDICATIONS AND NDE PROCESS VARIABILITY The basis for dispositioning indications in the kinetic expansion has been, and continues to be, that even full through-wall damage can be acceptable with respect to both structural integrity and primary-to-secondary leakage, depending on indication location and extent. Post-expansion ECT inspections of the kinetic expansion performed in the 1980's identified previously undetected indications. Depth sizing was not possible with the inspection technology that was used at the time (i.e., 8XI probe). It was concluded (NUREG 1019, Table 3.3-1) that small indications possibly having through-wall extent would not impact the reliability of the joints. More recent analyses, described herein, have also reached this conclusion.

30 ECR# 02-01121, Rev.22

4.1 Examination Techniques and Variability 4.1.1 Examination Techniques Kinetic expansion examinations are currently performed with MRPC probes (i.e., Motorized Rotating Pancake Coil). These probes contain a mid-frequency Plus-Point coil and a 0.080" diameter high frequency shielded pancake coil that are used to detect and/or evaluate indications in the kinetic expansions. The 300 kHz Plus-Point coil data is used for detection and depth sizing of detected flaws in the kinetic expansion region. The 300 kHz Plus Point coil data is used for length sizing of kinetic expansion circumferential or axial "crack-like indications". The 600 kHz 0.080" pancake coil data is used for measuring the axial and circumferential extents of ID volumetric flaws. These examination techniques are able to characterize the flaws in terms of morphology, surface extent, depth of the flaws, and axial location of the flaws within the expansions.

TMI-I will not change eddy current techniques used for examining the kinetic expansions unless prior NRC approval has been obtained.

4.1.2 Noise Levels Prior to the 2001 Outage 1R14, Outage 13R (October 1999) eddy current noise levels in the steam generator were compared to the applicable qualification data for the Plus Point coil by measuring the volts peak-to-peak and vertical volts ("volts vert-max"). This comparison was based on 300 actual in steam generator noise measurements and 168 qualification data set noise measurements. The comparison revealed that the general population of tubes was expected to have noise levels equivalent or less than the qualification data set noise levels. In fact only one of the 300 in-steam generator noise measurements exceeded the measured qualification noise measurement and the vertical volts noise measurement difference at this location was less than 0.05 volts. (The volts peak-to-peak measurements were all less than the maximum measured volts peak-to-peak measurement for the qualification data.)

Prior to 2003 Outage IR15, Outage IR14 eddy current data noise levels were compared to prior examination data for tubes that were in situ pressure tested at TMI. The in situ pressure tested flaw population at TMI-I is large in comparison to the flaw population in the industry qualification data and structural and leakage performance has been acceptable for all of the flaws in situ pressure tested to date. The in situ pressure tested flaw population also provides a more diverse population of flaws. To date, more than 69 ID IGA flaws having eddy current measured circumferential and axial extents up to 0.37" and 0.40" extents, respectively, have been in situ pressure tested. To date, 10 circumferential indications have been in situ pressure tested with the maximum measured eddy current extent being 0.51". (Refer to Table 5 for a summary of TMI-I steam generator tube in situ pressure tests performed to date.) All of the in situ pressure tested flaws were located below the kinetic expansion and the test results are conservative compared to expected performance for a similar flaw inside the expansion because the tested flaws would not have had the tubesheet ligament providing structural support. Thus, if the measured noise in the kinetic expansion region is equivalent or less than the measured noise in the in situ pressure tested tubes, similar or better examination results will be expected. These measurements were made using RMS noise vertical and horizontal measurements as described in the EPRI Steam Generator Examination Guidelines (Reference 38). The in situ pressure tested tube noise values were based on 32 tubes with 107 flaw locations. The kinetic expansion tube noise measurement values were based on 70 tubes evenly distributed in both steam generators in order to obtain a 31 ECR # 02-01121, Rev. 2

representative sample of the tube bundles. Tube locations in the generator were chosen to assure a sampling across the tubesheet array. The noise values were measured for both the 300 kHz Plus Point coil channels and the 600 kHz 0.080" shielded pancake coil channel (the channels used for kinetic expansion flaw detection and sizing). The study illustrated that the noise within the kinetic expansion data is comparable to (i.e., is not more noisy than) data from TMI- I's in-situ pressure tested tubes.

Eddy current noise levels are monitored during TMI-I steam generator tubing examinations as part of the data quality verification process. There are currently no formally accepted procedures for quantifying the effects of measured noise levels on the probability of detection (POD) or sizing of flaws. The tube noise studies described above determined that the eddy current noise in the kinetic expansion region was similar to the noise in other regions where MRPC probes are utilized in the TMI-I generators. These studies also confirmed that the eddy current noise levels in the kinetic expansion region are comparable to the noise levels that were present in the data used to qualify the MRPC probes. This confirmation supports a conclusion that in generator POD and sizing errors are similar to, or better than, those supported by the qualification data and in situ pressure tested flaw population. Noise monitoring will continue to be performed during future TMI-1 examinations.

4.1.3 Examination Technique Qualification 1997 Analyses PWSCC and ID IGA sizing performance of rotating coil examinations in OTSG tubes were evaluated prior toTMI-l's 1997 Outage 12R. Machined flaws were introduced into OTSG tubes in order to represent circumferential, axial and volumetric damage. Table C below provides a summary of the machined OTSG tubing flaws used in this study. The study concluded that the 300 kHz mid-frequency Plus Point coil examination technique provided the best depth sizing performance and the best flaw extent measurement performance for axially- and circumferentially-oriented flaws. The 600 kHz 0.080" high frequency shielded pancake coil examination technique provided the best extent measurement performance for ID volumetric (ID IGA) indications.

Table C OTSG Tubing Machined Flaws Used in 1997 Study Flaw Type Flaw Quantity Nominal Depth Nominal Axial Nominal Range in Length Range Circumferential Percent in Inches Length Range in Throughwall Inches Axial Notch 10 20 to 80 0.06 to 0.25 0.004 Circumferential 20 20 to 100 0.004 0.06 to 0.50 Notch Volumetric Pit- 23 20 to 100 0.02 to 0.16 0.02 to 0. 16 L ik e_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

32 ECR # 02-01121, Rev. 2 l

Data from EPRI Appendix H qualifications for axial and circumferential PWSCC (ETSS's 96703, 96702, and 96701 based on 0.750" and 0.875" diameter tubing) was evaluated with the same analysis techniques in order to confirm the validity of the measured performance from the OTSG machined flaw tubing examinations. A comparison of the Appendix H qualification results against the OTSG machined flaw results confirmed the validity of the defined examination performance in the study and that use of the examination technique performance in evaluating kinetic expansion data would result in a conservative dispositioning of identified in generator degradation.

Prior to examining a large number of kinetic expansions in the 1997 12R Outage, the contributing sources of expected error during the MRPC examinations were segregated and evaluated separately. The primary source of error was technique error involving differences between the "as-called" values compared with metallurgical "truth". The other contributing factors were analysis variability due to differences between the results of eddy current analysts, and equipment/technique variability due to differences among multiple trials for the same analyst.

In order to establish examination extent and acceptance criteria it was necessary to establish the magnitude of each of these contributing sources of examination error. Using length sizing performance as an example, the relative sizing error was greater than the sum of analysis variability and equipment/technique variability. This result has significance because the average error for both circumferential and axial length sizing is an overcall. This means that the sum of all of the error contributing factors remains an overcall for axial and circumferential extent.

Since the overall performance was shown to be consistent overcall of flaw lengths, this helps ensure that tubes with unacceptable flaw lengths will be removed from service. Since the examination techniques overcall these extents, the "as-called" circumferential and axial dimensions, without any statistical correction, are used for length sizing.

Only those defects estimated to be greater than 67% through-wall are included in the kinetic expansion accident-induced primary-to-secondary leakage evaluation. (Leakage is highly improbable from shallow defects.) The logic for addressing the expected errors when depth sizing was similar to that for length sizing. In this case, however, an additive correction is used because the typical Plus-Point depth sizing error is an undercall. (ECT estimated the throughwall extent to be less than the actual throughwall depth.) The additive correction to the "as-called" depth is large enough to ensure the sum of all factors that contribute to error will result in an overestimate of throughwall depth.

Specifically, the additive correction factor for the mid-frequency Plus-Point probe depth estimate is 32.6% through-wall. Thus, for field implementation, any indication having an "as-called" depth greater than 67% through-wall is considered as potentially contributing to primary-to-secondary leakage, and is included in the leakage assessment calculations.

1999 Analyses Subsequent to the 1997 outage, additional analyses were performed to evaluate eddy current analysis errors for TMI-1 steam generator tube flaws. This study included the addition of 9 TMI-I pulled tube ID IGA flaws and 6 OTSG tube laboratory induced PWSCC flaws. The majority of TMI- I flaws are volumetric ID IGA indications. Axial and circumferential extents of the volumetric ID IGA indications in the freespan are measured using the 0.080" shielded high frequency pancake coil operated at 600 kHz. AmerGen's Reference 30 (RAI Question I) 33 ECR# 02-01121, Rev.2

response provided to the NRC the following information concerning length and width sizing of volumetric ID IGA indications:

"...TMI-I has evaluated eddy current techniques and expected analyst uncertainties so as to assure that the dispositioning of the ID IGA indications using MRPC probes is conservative. Before 1997's Outage 12R, a study was performed to evaluate the acquisition, analysis, and technique errors expected during the MRPC examinations of the ID IGA indications. Volumetric flaws manufactured by EDM were used in the 1997 study. This study was updated before 1999's Outage 13R so as to incorporate the data from the ID IGA flaws in the tube samples pulled during the 1997 outage. A team of 5 production analysts and I senior (resolution) analyst was used in the study.

"Acquisition variabilities were obtained by running three separate MRPC exams of the ID volumetric flaws. Comparison of the three separate exams by a single analyst enabled the acquisition errors to be evaluated. Since each flaw was a separate test, a pooled variance was used to combine the results. For the 0.080" HF pancake coil (the coil utilized by TMI-I to measure the extents of the ID IGA indications), the acquisition pooled standard deviations were 0.0114" for axial length and 0.0084" for circumferential length.

"Analysis variabilities were obtained by comparing the different analysis results of the six different eddy current analysts. For the 1999 study, this dataset included 23 EDM flaws and 9 flaws from the 1997 TMI-1 pulled tube, for a total of 32 volumetric flaws. For the 0.080" HF pancake coil (the coil utilized by TMI-1 to measure the extents of the ID IGA indications), the analysis pooled standard deviations were 0.022" for axial length and 0.03 1" for circumferential length.

"Technique variabilities were obtained by comparing the results of the eddy current analyses to the actual metallurgy of the flaws. Again, for the 1999 study, this dataset included the 23 EDM flaws and 9 flaws from the 1997 pulled tube, for a total of 32 volumetric flaws. For the 0.080" HF pancake coil (the coil utilized by TMI-1 to measure the extents of the ID IGA indications), the technique standard deviations were 0.039" for axial length and 0.033" for circumferential length. For the 0.080" HF pancake coil, the technique average errors were a 0.124" overestimate of axial extent and 0.127" overestimate of circumferential extent.

"The conclusion of the 1999 error analysis and performance evaluation is that ". ..the rotating coil techniques have demonstrated that axial and circumferential extents are consistently overestimated. Even when analysis and technique / equipment variability are applied at a 95% confidence level, the extents measured by eddy current are larger than the actual extents." The overestimation of axial and circumferential extents is of sufficient magnitude that no correction to the repair limits is necessary to account for eddy current acquisition, analysis, or technique uncertainty. Since the eddy current coils interrogate a volume of metal larger than the volume of the flaws themselves (i.e., "look ahead" and "look behind") the result is a consistent overestimate of flaw extents.

"Note that tube pull results from the 1997's Outage 12R demonstrated that the MRPC probe typically overestimates the axial extents of the ID IGA flaws by a factor of approximately three. This occurs due to the "look ahead" and "look behind" phenomena of eddy current coils used in steam generator tube examinations. Additional information on analyst uncertainty is provided in the response to RAI Question No. 4."

34 ECR# 02-01121, Rev. 2

Similar length sizing studies were performed for axially- and circumferentially-oriented indications prior to the 1997 and 1999 outages using the 30 machined notches from the 1997 study and 6 laboratory-induced, axially-oriented PWSCC cracks (added during the 1999 study).

These measurements were made using the mid-frequency Plus Point coil similar to measurements made in the field. The results of these studies indicated that the Plus Point coil, like the pancake coils, overestimates crack length.

In addition to the "look ahead" and "look behind" effects described above, another reason that the eddy current probes tend to overestimate the extents of the kinetic expansion flaws is that the flaws are typically small in comparision to the eddy current probe coil field sizes. No studies were performed to investigate whether the eddy current probes would overestimate the extents of flaws larger than those used in the sizing study. Refer to Section 4.1.4 for flaw sizes recorded during the 2003 examinations.

In the kinetic expansion region flaw depth measurements are made using the mid-frequency Plus Point coil. Prior to the 1997 and 1999 outages Plus Point coil depth sizing performance studies were performed in a manner similar to that described above for the length sizing studies. The 1999 study was performed using 68 total flaws that were comprised of 10 machined axial notches, 20 machined circumferential notches, 23 machined ID volumetric IGA like indications, 6 laboratory grown PWSCC indications in OTSG tubing, and 9 TMI pulled tube ID IGA indications. The 6 PWSCC samples were axially oriented ranging from 0.08" in length to 0.32" in length and 35% to 99% through wall. The 9 ID IGA pulled tube flaws ranged from 0.016" to 0.032" in circumferential length, 0.020" to 0.066" in axial length, and 19% to 49% through wall.

The studies indicated that the measured 95% lower confidence level (LCL) through wall measurement error is expected to be -28.1% through wall. [Note that the additive correction factor for the mid-frequency Plus-Point probe depth estimate was not changed from 32.6% to 28.1% after the 1999 study. Thus, for field implementation, any indication having an "as-called" depth greater than 67% through-wall is considered as potentially contributing to primary-to-secondary leakage, and is included in the leakage assessment calculations.]

It should be noted that the measured eddy current through wall estimate is used for estimation of accident-induced leakage only; the eddy current measured axial and/or circumferential extent is assumed to be 100% through wall for evaluation of structural integrity (resistance to pull-out) as described in previous sections of this report. Based on the eddy current examination results, and in situ pressure tests of freespan indications performed at TMI to date, accident-induced leakage from kinetic expansion indications remaining in service is expected to be very small.

In summary, the eddy current techniques used at TMI-I are based on qualification datasets that included pulled tube samples from TMI-l and other samples representative of TMI-1 's ID degradation. Performance studies have demonstrated that eddy current sizing is conservative, and both pulled tubes and in situ pressure testing to date have demonstrated that the techniques used at TMI-I are able to reliably disposition steam generator tube flaws.

4.1.4 Examination Technique Qualification Performance Applicability for TMI-1 OTSG's Sections 4.1.1 through 4.1.3 provide information on the extensive examination performance studies that were performed using machined flaws, laboratory induced cracking flaws, in situ pressure tested tubes and TMI-1 pulled tube flaws. The techniques and examination errors provided in these studies can be applied to TMI-1. The eddy current qualification described in 35 ECR # 02-01121, Rev. 2

this report provides a strong case that the errors identified are applicable to TMI-I, however, other applicable factors have validated the examination techniques. The factors are listed below and will be further described in this section:

• The qualification data represents TMI-l flaws

• In situ pressure testing of similar freespan degradation has supported structural and leakage integrity conclusions

• There is a large population of known flaws examined each outage

• Flaws caused by this damage mechanism have successfully been in service since the mid-1980's

• Primary-to-secondary leakage is not present

• Statistical evaluations conclude that this damage mechanism is non-active The qualification data represents TMI-1 flaws - The TMI-l OTSG tubing qualification data represents the in steam generator degradation in terms of morphology, size and through wall dimension. Both the qualification data and indications remaining in service are ID initiated. The 1997 destructive examination of a TMI-1 tube with known ID volumetric degradation did not identify additional degradation beyond that identified with eddy current. Figure 24 below provides the measured axial and circumferential extent of "VOL" (ID IGA indications) and circumferential extent of "Circ" (circumferential) indications detected and sized during Outage IR15. Figure 24 below is based on 995 ID volumetric indications and 110 circumferential indications detected during Outage IR15. Only 3 of the ID volumetric indications measured

>0.30" axial extent and 8 measured >0.30" circumferential extent. In fact, 86% of the ID volumetric indications measured <0.20" circumferential extent and 94% of the ID volumetric indications measured <0.20" axial extent. Of the circumferential indications only 3 exceeded 0.50" in circumferential extent. The qualification flaw dimensions are similar to these dimensions (see Table C). The qualification flaw data set included flaws from 20% through wall to 100% through wall (essentially the full spectrum of flaw depth).

Figure 24 Distribution of 1R15 Kinetic Indications 900 800 700-a 600 - laVOL Axial Length 500-400 ._ a VOL Circ Length 0 300 - 'Circ' Circ Length 200 .

100N 0

0.00 to 0.11 to 0.21 to 0.31 to 0.41 to 0.51 to 0.10 0.20 0.30 0.40 0.50 0.60 Length In Inches 36 ECR# 02-01121, Rev. 2 l

In situ pressure testing of similar freespan degradation has supported structural and leakage integrity conclusions - TMI-I has in situ pressure tested a large population of flaws.

To date this includes more than 69 ID volumetric indications and 10 circumferential indications.

The maximum tested ID volumetric indication axial and circumferential extents in situ pressure tested to date are 0.40" and 0.37" respectively. The maximum circumferential extent of a tested circumferential indication is 0.51". All of these ID indications were tested at locations below the kinetic expansion (more conservative test because the tubesheet ligament does not provide additional structural support). All of these indications demonstrated acceptable structural and leakage integrity. Based on results from Outage 1R15; the maximum ID volumetric indication axial and circumferential extents of indications remaining in service is 0.35" and 0.41" respectively and the maximum measured circumferential extent of circumferential indications remaining in service is 0.60". Based on the Outage IRI5 examination results, the prior in situ pressure tested flaw population strongly represents the remaining inservice population of kinetic expansion indications.

Table 5 provides summary data of the in situ pressure tests performed on TMI-I flaws during Outages 12R (1997), 13R, (1999), and 14R (2001). This data was excerpted from the outage reports previously forwarded to the NRC (References 32, 35, and 36). [Note that TMI-I changed its voltage normalization criteria between Outages 12R and 13R; therefore voltages between Outage 12R and later outages must be adjusted to make voltage comparisons.]

There is a large population of known flaws examined each outage - The planned initial sample of tubes to be examined in the kinetic expansion region is large (approximately 30,000 tubes each outage). This scope is 100% percent of the in-service tube population. This large population assures that, even if there were a lower than expected probability of detection, new degradation would be evident in the examination results. This large population of known flaws assures that changes to flaw dimensions will be successfully identified.

Flaws caused by this damage mechanism have been in service since the middle 1980's - The kinetic expansion flaws are due to the sodium thiosulfate intrusion that occurred in the early 1980's. Tubes damaged by this mechanism have been in service since that time and no active growth has been shown based on statistical studies. Tubes damaged by this mechanism have demonstrated structural and leakage integrity since restart following repairs for this damage mechanism (about 18 calendar years of service) and the sodium thiosulfate inventory has been eliminated from the plant's design.

Primary-to-secondary leakage is not present - The TMI-I steam generators have demonstrated acceptable primary-to-secondary leakage during recent operating cycles. This indicates that tubes damaged by the sodium thiosulfate intrusion continue to perform acceptably.

Statistical evaluations conclude that this damage mechanism is non-active - Detailed statistical evaluations referenced in this report have concluded that this damage mechanism is non-active. The statistical evaluations also provide evidence that the applied examination techniques are repeatable. This document requires that growth studies be continued in future examinations, with the results reported to the NRC.

In summary, all information provided in Section 4.1 of this report, when considered as a whole, provides strong evidence that the applied examination techniques and their related uncertainties have been demonstrated to be applicable and conservative for TMI- 1.

37 ECR# 02-01121, Rev. 2 l

4.2 Conservatism of Measured Depth Criterion The 67% throughwall threshold for the leakage estimate is a very conservative criterion considering:

- the 33% TW eddy current accuracy (i.e., 100% minus 67%) was based on the results of the 1997 eddy current analysis with a 95% single tailed lower confidence level. A team of analysts was used for the study to evaluate error. In addition, a 1999 evaluation determined that 28% accuracy could have been used.

- a number of additional conservatisms are incorporated into the leakage assessment methodology. For example, volumetric indications are hypothesized to form both a circumferential crack and an axial crack, with the entire measured eddy current extents used to calculate expected accident leakage.

- the majority of the indications within the TMI kinetic expansions are ID volumetric IGA indications. In-situ pressure testing of ID volumetric IGA indications at TMI to date has not identified any indications that have demonstrated measurable leakage (i.e., leakage above detectable levels) at simulated normal operating or accident conditions. For example, 69 ID volumetric indications were in situ pressure tested, without leakage, during the plant's IR14 refueling outage in 2001 (Reference 32).

The results of in situ pressure tests performed during recent refueling outages also provide some additional evidence that the depth estimates of TMI-1 steam generator tube flaws are conservative. For example, during the IR14 Outage, seven TMI-1 tube indications whose estimated depth by Plus-Point was greater than 80% throughwall were insitu pressure tested.

(Reference 32) None of these seven indications leaked at a delta pressure equivalent to three times the delta pressure during normal plant operation (i.e., 3NODP). One of these seven indications, with an estimated depth of 97% throughwall, leaked at a rate of 0.014 gpm, a small leakrate, at a delta pressure of 6450 psi, approximately five times the delta pressure during normal plant operation. All seven of these indications had estimated depths greater than 67%

throughwall and would have been assumed to leak at MSLB delta pressure, which is less than 3NODP delta pressure, under the kinetic expansion leakage criteria.

4.3 Evaluation of Kinetic Expansions with Indications If any flaws are detected within a kinetic expansion, the eddy current analysts document the locations, measurements, and types of flaws within the expansions. Evaluation of the flaws with respect to the repair criteria, and leakage estimates, are performed by the plant's engineers.

Note that the expansion transition (i.e., below the ETL+0.00" reference point) is considered freespan for indication disposition purposes. The kinetic expansion transitions are treated as freespan tubing since they are not expanded against the tubesheet bore and do not benefit from any compressive residual stresses such as those present in the expansions.

4.4 Repair Criteria Application As described above, kinetic expansion evaluations are performed beginning at the ETL + 0.00" location to verify that sufficient defect-free lengths are present. Structural evaluations of the kinetic expansions require that a kinetic expansion be removed from service if insufficient defect-free length is identified over its examined length. That is, if a defect (or a combination of 38 ECR # 02-01121, Rev.2

defects) is detected that exceeds the allowable circumferential extent acceptance criterion, or an insufficient axial length of defect-free expansion is present, the expansion is removed from service. The inspection of a kinetic expansion may proceed farther (i.e., higher) in the tubesheet if flaws detected during the course of the examination within that expansion are within the conservative structural acceptance criteria. Figure 2, below, provides a visual presentation of the "defect-free" concept for a kinetic expansion with two indications.

If a volumetric ID IGA flaw is detected in a kinetic expansion, the TMI-l dispositioning criteria conservatively assume that the joint is not usable for structural purposes over the entire axial length of that flaw. For example, if a small volumetric flaw is detected with an eddy current-measured axial extent of 0.15", the entire 0.15" length of the expansion (360 degrees around the surface of the tube) is not credited in the evaluation of the joint structural integrity. In addition, no credit is taken for defect-free tubing along additional axial lengths of the joints adjacent to flaws (known as flaw "influence zones"). In summary, sufficient defect-free tubing must be detected to verify the integrity of an expansion during an inspection; no credit is taken for the length of the kinetic expansion where any defect is present, or where any defect might influence joint integrity.

(Note that, as described above, beginning in the IR16 refueling outage scheduled for the fall of 2005, AmerGen will plug all tubes with circumferential flaw indications in their kinetic expansion's required length upon detection-including circumferential flaw indications that were detected during prior 1997 through 2003 outage examinations that remain in service in the kinetic expansions' required lengths.

While the kinetic expansion structural dispositioning criteria are very conservative, there is no requirement that the defect-free joint length be "continuous". The kinetic expansions are flaw tolerant. (Burst is precluded due to the presence of the tubesheet; residual compressive stresses are present; bending stresses and vibration are limited; secondary side loose parts are prevented from impacting the tubing.) Small defects do not influence the reliability of the kinetically expanded joints. For example, a small volumetric ID IGA pit on the surface of a kinetic expansion will not impact the ability of defect-free tubing, located above or below that pit, to maintain the structural requirements of the joint (e.g., no tube parting, no joint pullout). Outside of the flaw influence zones a small ID-initiated axial crack present along the length of a kinetic expansion would not adversely affect the structural integrity of defect-free tubing located above or below that crack. From a structural standpoint, so long as no flaw or combination of flaws is present with a circumferential extent greater than 0.64", the defect-free tubing located above or below the flaw is an integral part of the kinetic expansion joint. (If the 0.64" circumferential extent value is exceeded prior to the required defect-free length being observed, the kinetic expansion is repaired, since the tube, conservatively assuming 100% throughwall degradation, could theoretically be parted under calculated accident-induced loads.) The expansion evaluations only "move higher into the tubesheet" if the examination data is available, and the repair criteria are not exceeded. The technical basis for this continued inspection (i.e., higher in the tubesheet) is provided in the finite element analyses of Reference 24.

39 ECR#02-01121,Rev.2 l

FIGURE 2 "Defect Free" Concept (Inside Surface of a Hypothetical Kinetic Expansion "Flattened" for this Sketch)

---Not to Scale---

EDI Indicadon and Indication Influence" Zone M I -Dcfect Fred' Zonc 5.0 LEAKAGE ASSESSMENT METHODOLOGY 5.1 Introduction and Background Primary-to-secondary leakage during an accident must not degrade the ability to provide adequate core cooling capacity nor cause unacceptable or unanalyzed radiological consequences.

The kinetic expansion inspection criteria provide assurance of joint structural integrity to the ends that joint failure will not occur either by slipping or by tube parting. Each of these failure modes has as a theoretical consequence the introduction of primary-to-secondary leakage.

40 ECR # 02-01121, Rev. 2 l

Theoretically,' through-wall defects that may be present in the kinetic expansion region may leak when subjected to MSLB conditions, even if these defects are not large enough to create a tube slipping or parting concern. The hypothetical MSLB axial loads and differential pressures could cause defects to open and provide a less restrictive leakpath than that provided by the tube-to-tubesheet joint during normal operation.

Primary-to-secondary leakage from the expansions is expected to increase during a postulated MSLB. The joint was originally qualified as leak-limiting and not leak-tight. However, in order to address even the possibility of increased primary-to-secondary leakage due to defects in the joint, a number of very conservative assumptions have been made in the leakage assessment methodology.

Defects that are judged to be through-wall, or near through-wall, by the inspection techniques are included in the primary-to-secondary leakage evaluation. While the analysis model for kinetic expansion structural evaluation assumed 100% through-wall, the analysis of accident-induced leakage utilizes through-wall depth information provided by the ECT.

In addition, some potential defects could be located at elevations where contact pressure between the expanded tube and the tubesheet bore remains, albeit reduced, during the accident. The presence of contact pressure considerably reduces leakage. The analysis model results showed that, for tubes that are not affected by tubesheet bowing (i.e., peripheral tubes), no part of the minimum required intact expansion loses residual contact pressure during the accident. Tubes that are affected by tubesheet bowing (i.e. tubes near the center of the bundle and mid-radius tubes) will locally lose contact pressure during the MSLB event. As a result, .the radial location of a tube within the bundle affects the estimation of leakage from flaws found in its kinetic expansion.

"As found" and "as left" leakage estimates for the kinetic expansions are calculated after each inspection. Because no flaw growth has previously been detected, and no growth is expected, it is necessary only to consider defects found in the joint that are dispositioned as acceptable and left in service as potential sources of future primary-to-secondary leakage. Defects that are unacceptable are repaired by plugging.

The purpose of this section is to describe the methodology that is used to evaluate the total primary-to-secondary leakage that may occur during a guillotine rupture of a main steamline as a result of assumed through-wall (>67% throughwall as measured by eddy current) cracks in the kinetic expansion region of the OTSG tubes. In Reference 17 it was demonstrated that the limiting accident scenario which results in the largest tube loads is that which results in a large SG tube-to-shell temperature differential (AT). The most restrictive limits were determined to be when the tubes are colder than the steam generator shell.

In order to establish the total primary-to-secondary leakage that would be acceptable during the MSLB event from assumed through-wall cracks in the kinetic expansion region, a calculation determined the maximum leakage that would meet the offsite dose criteria of 10% of 10CFRI00 limits for the 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> Exclusion Area Boundary (EAB) and 30 day Low Population Zone (LPZ)

(Reference 2). The revised dose consequences for the FSAR MSLB analysis were submitted to the NRC for approval (Reference 3). The results were as follows:

1. Integrated Primary Coolant Leakage @ 2 hrs (gallons @ 579 F) = 3228.

2. Total Integrated Primary Coolant Leakage (gallons @ 579 F) = 9960 41 ECR # 02-01121, Rev. 2

The methodology used to estimate leakage from the kinetic expansion indications, and to determine if these leakage limits are met, is discussed in the following sections. Section 5.2 provides an overview of the methodology and the subsequent sections provide additional detail.

5.2 Overview of Methodology As described in Section 2.2, the structural criteria utilized to disposition the kinetic expansion indications were based on Reference 17, a Main Steam Line Break analysis performed in 1980.

The resulting peak axial, tensile load on the steam generator tubes from this analysis was 3140 lbs.

In order to evaluate theoretical accident-induced leakage from the kinetic expansions new TMI-1 plant-specific MSLB analyses were completed in 1997. The Reference 17 1980 MSLB analysis was updated for the following reasons:

- The 1980 analysis was a 'generic' analysis for the B&W Owners Group (BWOG) plants (i.e., the analysis was not TMI-1-specific).

- The 1980 analysis assumed an EFW flow of 1650 gpm with operator action to isolate EFW to the affected OTSG after 20 minutes. The TMI EFW design includes cavitating venturies and a safety grade level control system. The response of the TMI- I EFW system to a MSLB would be to limit break flow to a maximum of about 570 gpm to the affected (depressurized) OTSG and to control level at 25 inches in the unaffected OTSG.

The difference in EFW flow to the affected OTSG of 1650 gpm vs. 570 gpm has a very significant effect on the cooldown of the steam generator tubing.

- The 1997 analysis assumed operator action to terminate EFW after 10 minutes. This is consistent with the plant's licensing basis FSAR MSLB analyses and emergency procedures. Since the volumetric flowrate of EFW used in the 1997 analysis (590 gpm was conservatively used) is considerably less than that of the 1980 analysis (1650 gpm),

there would only be a small effect on the 1997 results if the EFW isolation time was changed from 10 minutes to 20 minutes. (The difference in termination times would be very significant for the 1980 analyses since a very large EFW flowrate was assumed.)

- The 1997 analysis was more conservative than the 1980 analysis regarding reactor vessel mixing.

- The 1980 analysis assumed a 36-inch break at the OTSG nozzle to bound all of the BWOG plants. For TMI-I, this assumption was conservative because the plant has four 24-inch steam lines that only connect downstream of the Main Steam Isolation Valves, at the turbine chest. This difference has the most pronounced effect during the initial blowdown of the OTSG. (After the initial blowdown, the cooldown is dominated by the amount of EFW that is boiled out of the OTSG and would not be any different for a 24-inch or 36-inch break.) The 1997 analysis assumed rupture of a 24-inch TMI-I steam line.

The computer codes used for the 1997 analysis differed from those used for the 1980 analyses.

The 1997 analyses used RETRAN for the short term analysis and GOTHIC for the long-term analysis. The 1980 analyses used TRAP2.

The following sections describe the 1997 MSLB analyses performed to evaluate kinetic expansion tube leakage in more detail.

The methodology used for the MSLB-induced leakage analysis performed in 1997 involved the following activities that are depicted in Figure 3:

42 ECR # 02-01121, Rev.2 l

A. Develop the time varying thermal hydraulic (T-H) information from the design basis Main SteamLine Break (MSLB) event analysis.

B. Determine the OTSG tube tensile and differential pressure loads from the T-H data.

The loads vary as a function of time throughout the transient and as a function of radial distance from the center of the steam generator to the peripheral tube.

C. Calculate the theoretical crack opening area (COA) separately for postulated circumferential and axial cracks. The COA varies with the applied load, crack orientation, and crack length.

D. Determine the theoretical leakage flow as a function of the crack area. The total mass released from the crack is obtained by integrating the leakage flow over the first 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and over the entire transient interval.

E. The integrated leakage flow for each of the identified cracks (based on crack size and radial position within the tube bundle) is summed and the total is compared against the leakage limits specified in the offsite dose calculation (Reference 2) based upon 10% of the 10CFRlOO limits.

If the calculated leakage exceeds the limits established in Reference 3, then a decision will be made as to which tube(s) will be repaired (i.e., the leakage contribution from the repaired tube(s) can be eliminated from the total to meet the allowed as-left leakage limits.)

Additional details and references regarding each of the activities discussed above are provided in the sections which follow.

43 ECR # 02-01121, Rev. 2 l

FIGURE 3 LEAKAGE EVALUATION METHODOLOGY OVERVIEW 44 ECR # 02-01121, Rev. 2 l

5.3 Main Steam Line Break Analysis 5.3.1 Overview A conservative plant MSLB analysis was used to generate the transient thermal hydraulic parameters that were needed as input to define the OTSG tube loads and to calculate the leakage from each kinetic expansion flaw.

The transient analysis was accomplished in two phases: a short term phase and a long term phase. The short term phase duration was 10 minutes (600 sec) and utilized the transient systems analysis code RETRAN-02, Mod 5 (Reference 5). The long term phase thermal hydraulic conditions were developed by applying assumed operator actions, based upon TMI-I Anticipated Transient Procedures (ATPs), to recover from the event and to calculate the OTSG shell metal cooldown rate in order to develop a technical basis for cooling down to DHR conditions without violating tube-to-shell differential temperature limits. The long-term analysis began at 10 minutes and extended to the end of the transient (approximately 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />). Details of these evaluations are provided below.

5.3.2 Short Term Analysis 5.3.2.1 Basis of Duration As indicated above, this portion of the MSLB thermal hydraulic analysis included the first 10 minutes (600 sec) of the event. There were multiple reasons for choosing this duration. First, this portion of the transient is characterized by the most complicated and dynamically changing thermal hydraulic attributes. The affected OTSG is blowing down, the Heat Sink Protection System initiates a closure of the Main Feed Water (MFW) control valve and the MFW block valve and also initiates Emergency Feed Water (EFW) on low OTSG level. The RCS is depressurizing and cooling down, the pressurizer is emptying and refilling, an RPS trip occurs, ESAS is initiated, etc. Because of the complexity of this portion of the transient, a relatively sophisticated systems analysis code (RETRAN 02, Mod 5) was used to establish the thermal hydraulic parameters during this period (Reference 5).

Another reason for this duration is that no operator recovery actions were assumed to take place until after 10 minutes had passed. This is a licensing basis for TMI- l.

Following the first 10 minutes, credit for operator actions is permitted.

The peak axial, tensile tube loads for this event also occur within the first 10 minutes and the thermal hydraulic conditions at the end of this duration are important since they represent the end of the peak load period and the transition to reduced OTSG tube loads.

In this manner, the first 10 minutes of the MSLB analysis set the stage for the entire leakage determination effort. At the end of this period, the system is not characterized by rapid changes in thermal hydraulic conditions and is in transition to the recovery from the event.

45 ECR# 02-01121, Rev. 2

5.3.2.2 Methodology The RETRAN-02 MOD005 computer code and a TMI plant model were used to perform this analysis (Reference 4). The TMI RETRAN model has been extensively benchmarked against plant data and previously approved licensing codes. The benchmarks demonstrate the adequacy of the TMI RETRAN model for performing safety analysis. The TMI RETRAN model has also been approved by the NRC for referencing in licensing applications (Reference 5). The TMI Base deck (Reference 6) as shown in Figure 4 was used for this analysis.

FIGURE 4 - Three Mile Island Unit 1 RETRAN Two Loop Model Nodalization Diagram 5.3.2.3 Assumptions The analysis assumptions and initial conditions as discussed below were chosen to provide a conservative RCS overcooling and pressure history for the MSLB event and the resulting tube loads.

46 ECR#02-01121,Rev. 2 l

5.3.2.3.1 Initial Conditions The reactor was assumed to be operating at rated power prior to the hypothetical MSLB accident (2568 MWt). The initial pressurizer liquid level was set at 220 temperature-compensated inches, which is the typical hot full power (HFP) pressurizer level. The initial RCS pressure was 2170 psia in the hot leg, which is the normal operating value.

The TMI design basis MSLB assumes that offsite power is available and that was the assumption in this analysis. The effect of high RCS loop flow is to minimize the OTSG tube average temperature during the initial phase of the event. Thus, OTSG tube axial loads are maximized.

5.3.2.3.2 Break Modeling The initiating event was assumed to be a double-ended rupture of a 24-inch steam line on one steam generator. This is the largest possible break which results in the maximum cooldown rate. The faulted steam generator steam line was nodalized as shown in Figure 5, so as to model each steamline individually. The flow area of the two break junctions were consistent with the 24-inch steam line piping.

A Moody choking model was used for these break junctions with a contraction coefficient of 1.0 to maximize break flow rate.

The break was assumed to occur in the plant's Intermediate Building upstream of the Main Steam Isolation Valve (MSIV). This is an appropriate break location because it results in a ground level release of coolant activity.

47 ECR # 02-01121, Rev. 2

FIGURE 5-Break Nodalization 3 1 5.3.2.3.3 Reactor Vessel Mixing The amount of mixing that was assumed to occur within the reactor vessel was a ratio of the difference in hot leg temperatures to the difference in cold leg temperatures:

THOT (unfaulted) - THOT (faulted)

RATIO =

TcOLD (unfaulted) - TcOLD (faulted)

A value of RATIO = 0.0 implies perfect mixing while RATIO = 1.0 implies no mixing.

For the purposes of this analysis, a target value of RATIO = 0.5 was chosen to conservatively bound the analyses at an upper value.

To simulate this mixing in RETRAN, the reactor vessel was modified to include two equal parallel flow paths by splitting the downcomer, the lower plenum, the core, and the upper plenum as shown in Figure 6. For the most part, these parallel flow paths behave independently, with the exception of common connections with the bypass and upper head volumes. These common flow paths keep the loop pressures in balance but contribute little to mixing of loop flows.

48 ECR#02-01121,Rev.22

5.3.2.3.4 Reactor Kinetics Parameters To minimize the power increase response to the core temperature decrease, the moderator temperature coefficient (MTC) was set to a value of zero. This was conservative since it will not increase the power prior to trip and results in lower RCS temperatures. Post trip, the MTC determines the extent to which the core energy generation is increased by sub-critical multiplication. An MTC of zero will assure that the post trip reduction in temperature will not lead to increases in power generation above the normal decay heat power. The absence of a return to power after the trip results in a greater cooldown, and therefore a larger axial load on the steam generator tubes.

Decay heat was based on the ANS5.1 1979 decay heat standard. In order to maximize RCS cooldown following reactor trip, a 0.95 multiplier on decay heat was used. The 5% reduction was chosen since it is greater than a 2(sigma) uncertainty for thermal fission of U235 under equilibrium operating conditions.

FIGURE 6 RPV Nodalization 18 1

_ -~

UNFAULTED FAULTED LOOP - - LOOP 134 233 IR E

3 13 133 0t,2T BP 12 I 1 I8I 49 ECR # 02-01121, Rev. 2 l

5.3.2.3.5 Reactor Trip With an MTC of zero, the reactor power will not increase with the decrease in moderator temperature, so the reactor will trip on low RC pressure. Since this analysis was primarily interested in steam generator tube temperature, a trip setpoint of 1900 psig plus a 30 psi error was used. This limits the amount of energy the core model generates, resulting in a lower primary system temperature during the event. It should be noted that this setpoint results in an earlier trip, which is conservative for tube temperature calculations. For the steam line break event, the trip setpoint will be reached rapidly due to the dramatic overcooling which would occur.

5.3.2.3.6 Initial Steam Generator Mass The initial steam generator inventory provides a measure of the heat removal capability of the secondary system. For a steamline break, a larger initial secondary system inventory in the steam generator associated with the break will lead to a higher integrated heat removal. The larger the heat removal, the lower the resultant reactor coolant temperature. The OTSG design has the maximum inventory at full power conditions.

Thus the event should start from full power to maximize the heat removal capability of the steam generator. The steam generator inventory can increase if fouling of the SG tube bundle region occurs. The inventory predicted for full power and fouled conditions has been conservatively determined to be approximately 55,000 pounds per SG, and this value was used in the model. In addition, the mass of feedwater between the isolation valves and the affected steam generator, which was calculated to be 35,500 Ibm, was also modeled and available to cool the affected steam generator.

5.3.2.3.7 Main Feedwater and Emergency Feedwater Flow The MSLB accident in this calculation assumed the worst single failure, which is the failure of the feedwater regulating valve to close on the affected generator. This maximizes the overcooling of the event by maximizing the main feedwater 50 ECR# 02-01121, Rev. 2

FIGURE 7 Main FeedWater Flow Rates C

Main Feedwater , I Flow (Ibs./sec)

z Elapsed Time (sac) 51 ECR # 02-01121, Rev. 2 l

(MFW) flow to the affected generator as a result of the preferential feeding to the broken, depressurized, side. Feedwater flow to the affected steam generator is shown on Figure 7 above. MFW flow was terminated to the affected steam generator after the MFW block valve closes in about 30 seconds after a low SG pressure of 600 psig is reached.

For this transient, the Emergency Feed Water (EFW) system would be initiated by a low OTSG level signal. The OTSG low level indication signal of 10 inches is measured by the startup range instruments. The setpoint is calculated in the RETRAN model as the collapsed liquid level in the tube region. (Zero inches indicated level is 6 inches above the upper face of lower tube sheet.) EFW controls level at 25 inches indicated. Due to the continued MFW flow to the broken SG until the MFW block valve closes, the OTSG level does not drop below the low level initiation signal until about 67 seconds after the start of the transient.

FIGURE 8 Emergency Feedwater Flow Rates To Byken SG Emergency Feedwater Flow (Ibs/sec)

Intact SG o __________ +/-

.t.

100 210 000 400 soo Elapsed Tim (sec) 52 ECR # 02-01121, Rev. 2 l

The start of the motor driven EFW pumps (MDP) is delayed by 5 seconds after the initiation signal and a coastup time of 10 seconds. Subsequent to the EFW initiation signal, the steam admission valve to the turbine driven pump (TDP), MS-V 13A, receives an immediate open signal and is fully open in 24 seconds. Turbine testing shows the TDPs are at full speed in 11 seconds after the steam admission valves are full open. An additional 8 seconds for flow coastup is typically modeled resulting in TDP flow delivery at 43 seconds.

For this analysis, 2 MDPs and TDP were conservatively assumed to deliver flow instantaneously to the steam generator following an EFW initiation signal (See Figure 8 above).

5.3.2.3.8 High Pressure Injection The plant's high pressure injection (HPI) system is actuated during the cooldown period following a large area steam line break. The system supplies borated water to the RCS to recover the RCS shrink and to provide core cooling if necessary, and to increase the core shutdown margin. Boron addition to the reactor coolant, during the controlled cooling to atmospheric pressure, will prevent criticality at lower temperatures. For this analysis, no credit was taken for boron addition resulting from HPI actuation, since the BOL kinetics and best-estimate rod worth will result in keeping the core shutdown. To minimize the primary system temperature, and thus tube temperatures, full HPI was initiated in the model on a signal of 1600 psig plus a 30 psi error at the pressure measurement tap location. This is conservative, since a rapid actuation of HPI will maximize the overcooling.

5.3.2.3.9 Steam Generator Downcomer Modeling The RCS cooldown was maximized by minimizing the amount of liquid carried over from the steam generator out of the break. To minimize the liquid carryover, the downcomer was modeled with a single bubble rise volume and a large bubble velocity (1E6 ft/sec) which produced less liquid carryover.

5.3.2.4 Summary of Results 5.3.2.4.1 Power Results The results of the MSLB analysis for the first 10 minutes (600 sec) are provided in this section. The reactor scram occurs on low reactor pressure in about 10 seconds as shown in Figure 9. This reflects a trip setpoint of 1900 psig plus a 30 psi error.

The reactor power in Figure 9 also indicates that there is no return to power as a result of the absence of a negative moderator temperature feedback. This is a conservative result with respect to the cooldown.

53 ECR # 02-01121, Rev. 2

FIGURE 9 Reactor Power Normalized Power

'i -

0 100 200 o00 400 500 400 Elapsed Time (8eC) 5.3.2.4.2 Loop Temperature Results The hot and cold leg temperature responses to the MSLB are shown in Figure 10. A rapid overcooling results from the event with the cold leg temperature reaching about 435 degrees F about 70 seconds after the break. After the OTSG blowdown is completed, the primary to secondary heat transfer is reduced and the cold leg and hot leg temperatures are essentially the same. The temperature is about 450 degrees F at this point and is maintained for the duration of this portion of the event. The final temperature for this phase of the event reflects the fact that the intact OTSG acts as a heat source as discussed below.

54 ECR#02-01121,Rev.2

FIGURE 10 RCS Faulted Loop Temperatures r.

. I Temp (deg F) To Io To 100 . 200 300 4*0 500 400 elapsed Time (9cc) 5.3.2.4.3 OTSG Pressure Results The pressure response results for both the faulted and unfaulted OTSG are shown in Figure 11. The faulted OTSG is fully depressurized in about 100 seconds.

The unfaulted OTSG responds initially in a normal post trip manner, increasing to the MSSV setpoint, but is slowly reduced in pressure as a result of reverse heat transfer to the RCS.

5.3.2.4.4 RCS Pressure Results The RCS pressure results are depicted in Figure 12 and reflect a rapid drop in pressure due to the initial cooldown. The decrease in pressure results in a reactor trip, ESAS actuation, and a small influx of Core Flood Tank flow. After the cooldown has stabilized, the RCS repressurizes in response to HPI injection flow refilling the pressurizer. At the end of 10 minutes, the RCS subcooling margin is less than 100 degrees F.

55 ECR # 02-01121, Rev. 2

FIGURE 11 Steam Generator Pressure Response Inw tDSG SG Pressure (psia)

\ . DEmkn SG

_I . l 0

• 0 200 200 420 Ekapsed Time (sc) 56 ECR # 02-01121, Rev. 2 l

FIGURE 12 Pressurizer Pressure To To g

So Pressure (psia) g n

9 Io

-t

• t 200 300 400 500 400 Elapsed Time (Secl 5.3.3 Long Term Analysis 5.3.3.1 Approach Following the first ten minutes, it was assumed that operator action would be taken to terminate EFW to the affected OTSG and to begin a controlled cooldown and depressurization to DHR conditions using the unaffected OTSG. The limitations imposed by the various cooldown P-T limits and tube-to-shell differential temperature limits would be observed. The following assumptions reflected this approach.

57 ECR # 02-0112 1,Rev. 2

5.3.3.2 Assumptions

1. The operator will control the NSSS such that the tube-to-shell differential temperature tensile limit of -70'F (tube temp minus shell temp) is observed (Reference 9).

2. RCS temperature will not be allowed to increase to reduce the tube-to-shell differential temperature (Reference 10). Procedure guidance has the operator minimize the RCS reheat following an overcooling event. Increasing RCS temperature for this analysis would reduce (i.e., make less negative) the tube-to-shell differential temperature and reduce the tube load. Reduced tube load would lead to reduced tube leakage.

3. RCS pressure will be maintained at a subcooled margin of 750 F. Reference 10 directs the operator to minimize the RCS pressure increase following an overcooling event. The minimum SCM limit is 250 F (Reference 9). An RCS pressure control value of 751F SCM is reasonable. Higher RCS pressure leads to greater tube leakage.

4. As RCS temperature and pressure decrease, additional pressure limitations are established. The operator will maintain RCS pressure in excess of the emergency RCP NPSH limit (Reference 9). A margin of 50 psi is considered to be adequate. A high margin maintains RCS pressure high, increasing tube leakage. However, a large margin to the NPSH curve could prevent initiation of DHR. Therefore, a margin of 50 psi is reasonable. Additionally, the operators will maintain RCS pressure such that the minimum RCP seal differential pressure (275 psid) is maintained (Reference 12). .Seal return can be dumped to the sump instead of being sent to the Makeup Tank. A margin of 25 psig is maintained to the limit of 275 psid. Therefore, a minimum RCS pressure of 300 psig is established.

5. The transient after 600 seconds is quasi-steady-state. Therefore, large time steps could be used in the model. A time step size of 600 seconds was chosen as reasonable.

6. Operator action is assumed to take place at 10 minutes. The following actions would be taken by the operator for a MSLB event (Reference 9 and 10):

a. Terminate EFW to the broken OTSG (MFW is already isolated).

b. Control/terminate HPI to the RCS to control RCS pressure.

c. Adjust the TBV on the Unbroken OTSG to prevent RCS temperature from increasing.

5.3.3.3 OTSG Cooldown Analysis As indicated above, the operator will control the NSSS such that the tube-to-shell differential temperature tensile limit of -70'F is observed. The maximum 58 ECR#02-01121,Rev.2

possible cooldown rate that meets this criterion is established by the rate at which the affected OTSG shell cools down.

To determine the shell cooldown rate, the GOTHIC computer code, version 5.0e, was used with a six (6) volume model as shown in Figure 13 (Reference 11). Two volumes (volumes I and 2) represented the primary (tube) side of the OTSG, two volumes (volumes 3 and 4) represented the secondary (steam side) side of the OTSG shell inside the shroud, and two volumes (volumes 5 and 6) represented the secondary side of the OTSG outside the shroud (i.e., between the shroud and the shell metal). The volumes were divided to correlate with the division of the downcomer region into upper downcomer and lower downcomer regions.

The analysis began at 10 minutes and allowed the RCS to cool down as the shell cooled down to preserve the -70 deg limit and thus account for the impact of the cooler RCS tube temperature on the cooldown rate of the shell.

The shell cooldown rate results from this analysis are shown in Figure 14 below.

FIGURE 13 GOTHIC Model For Shell Cooldown Analysis Post-YSLB OTSG Cooldown - Multi-Node SG Wed Sep 17 09:53:36 1997 GOTHIC Version 5.0(QA)-e - October 1996

'I-Tf- b -t I - I I I I I INI I I I.

I I

=-

-. L J I I

I I I I I I I I I I4 l I L I L _ _ _

59 ECR # 02-01121, Rev. 2 l

5.3.3.4 Results Figures 14 and 15 below provide the results of the long term analysis. The figures also include data from the first 600 seconds of the analyzed event as well. The results reflect the application of the criteria described above. The average shell temperature is a weighted average of the upper and lower shell temperatures at the outside metal surface of the OTSG. The RCS temperature is the average of the hot and cold leg temperatures for the affected OTSG.

FIGURE 14 MSLB Temperature Response 650

. . . RCS9T---Mm

-AveMe Sh~eN 600

_ W-

.1 * .

200

• 10 100 1000 10000 100000

• Tkne (s"c}

60 ECR # 02-01121, Rev. 2 l

FIGURE 15 MSLB Pressure Response i

100 Tim. (soc) 5.4 OTSG Tube Loads 5.4.1 Introduction The resulting steam generator tube loads were determined from the T-H parameters provided from the analysis presented in Section 5.3 above. The method of calculating the tube loads evaluated the theoretical tubesheet deflection under a differential pressure and tube axial load and as a function of the different OTSG tube, tubesheet and shell metal temperatures. The resulting pressure and tensile loads were used to determine the leakage area that would develop for a given crack length and orientation as described in Section 5.5 below. Since the thermal hydraulic conditions changed with time, the resulting tube loads also change accordingly. As a consequence of the tubesheet deflection from the center to the periphery, the tube loads varied as a function of the radial distance from the center of the OTSG. In this way, a plot of the tube loads as a function of radial distance from the OTSG center to the OTSG periphery would be different for each set of consistent T-H conditions. The discussion below provides an overview of the methodology used by both GPUN and FII to independently determine the OTSG tube loads using the T-H data in Section 5.3, and a presentation of the results.

61 ECR # 02-01121, Rev. 2

5.4.2 Methodology 5.4.2.1 GPUN Methodology The methodology that was employed by GPUN for the determination of the tube loads is described in Reference 16 and comprised the following steps:

• Establish the tubesheet behavior as a function of applied load and material properties as a function of temperature.

• Establish the tube loading (pre-load) in the OTSG as a function of the measured gap between the separated sections of a failed tube at the temperature at the time of measurement. The calculation will be based on the assumption that very few tubes have parted so that the loading on the balance of the intact tubes is unchanged.

• Separate the three major OTSG components (tubes, shell, and tubesheet) to free components (bodies), remove all loads acting on them and find their unloaded geometry.

• Establish the physical variables that will result in deformation of the free bodies and calculate these deformations, including an accounting for the Poisson effect on the tubes and on the shell.

• Re-combine the deformed free components by pulling the tubes until they meet the final tubesheet location. The final tubesheet location must simultaneously satisfy both of the following conditions:

• The tubesheet periphery must be at the same location as the shell.

• The tensile load from all of the tubes must be equal to the shell compressive load.

5.4.2.2 Framatome Technologies, Inc. (FTI) Methodology An ANSYS finite element model of the OTSG was used to determine the tube load contribution for various system operating parameters. The ANSYS model was basically identical to the NASTRAN model used in the 'OTSG Tube Topical Report' (Reference 17). The NASTRAN model was converted to ANSYS due to some extra features ANSYS possessed at the time.

The model was an axisymmetric thermal and structural model of the OTSG. The model included the steam generator shell sections, upper and lower heads, upper and lower tubesheets, support skirt, and twelve beams representing twelve effective tube regions.

The tubesheet model accounted for the material properties which were adjusted to account for the tubesheet temperature and the effects of the perforated plate.

Several different load cases (parameter study) were executed to establish the variation in tube loads due to change in primary pressure, secondary pressure, tube-to-shell delta T (both tubes hotter and cooler than the shell), and average tube temperature. The end 62 ECR # 02-01121, Rev. 2

result was a series of equations as a function of average temperature and tubesheet radius, that provided the load in the tubes for each of the pertinent system parameters.

Using the postulated MSLB system transient parameters discussed in Section 5.3 above, the total tube loads for the transient, as a function of transient time and tubesheet radius, were determined.

5.4.3 Results 5.4.3.1 GPUN OTSG Tube Loads The GPUN analysis results are provided in Figure 16. This figure shows the OTSG tube loads for three radial positions in the OTSG (Center, Average, and Periphery) as a function of time from the start of the MSLB transient. The peak axial tube load of 1310 lbs. occurs 60 seconds into the transient at the periphery of the OTSG. The smallest loads occur at the center of the OTSG tube bundle as was discussed earlier.

5.4.3.2 FTI OTSG Tube Loads The FTI results (Reference 22) are provided in Figure 17. As can be seen, they were very similar to the GPUN load results. The peak axial tube load was 1135 lbs. at 60 seconds and also occurs at the OTSG periphery, with the smallest loads at the center as well.

A comparison of the GPUN and FTI results is provided in Section 5.4.4 below with an explanation for the loads that were used to perform the subsequent tube-to-tubesheet interface pressure and the leakrate analyses (which are described in Sections 5.5 and 5.6).

63 ECR#02-01121,Rev.2

FIGURE 16 GPUN Tube Loads 1000 gmoof

-3

-,0

- a 400 id ico 1.00 100000 Time Isec)

FIGURE 17 FTI Tube Loads 1400

. _Rt~ -50.4" 1200' 100 40

- 10 1000 . 10000 100000 Time psec) 64 ECR # 02-01121, Rev. 2 l

5.4.4 Analysis of Loads Figure 18 provides a comparison of the FfI and GPUN OTSG tube load results. Results are presented for three points in time as a function of radial distance from the OTSG center to the periphery. While the results were very close, it can be seen that the GPUN results tended to be more conservative than the FTI results as radial location (R) increases. Similarly, for smaller R, the FTI results were slightly more conservative. The plot of area ratio vs. radial position (right side ordinate axis is the area ratio) shows that there are substantially more tubes at the higher R values than at the lower R values. It was judged that the GPUN results would be more conservative since they would result in higher loads on a greater number of tubes. As a result, for this study, the GPUN-calculated loads were used to perform the subsequent crack area and crack leakage analyses described below.

FIGURE 18 Tube Load Comparison Peak Load: Bosac Local Maxdmum: 00 sec Long-Tenm Minimwn: 24800 sac 1400 1.1 1300 1200 1100

• 1X o 900

-I

?- S0o

.9700 0 5 10 15 20 25 30 35 40 45 50 55 10 Tubeshs Radius (In)

The two sets of independent analyses were confirmatory and demonstrated that the calculated OTSG tube loads are reasonable.

65 ECR # 02-01121, Rev. 2 l

5.5 Crack Area Determination 5.5.1 Introduction The crack opening area (COA) determination was based upon the methodology provided in Reference 13 and established a method for calculating the crack opening area for through-wall cracks in tubes. Primary-to-secondary leakage was calculated using two potential crack orientations in combination with a specific applied load (Reference 14). These were:

I. Circumferential Through-Wall Crack in Tension (Note: The contribution of primary pressure is included in the applied tension load.)

2. Axial Through-Wall Crack Subjected to Internal Pressure Using these methods, the user could calculate the crack opening area (COA) for a crack given the specified conditions and use that area to determine the tube leakage (See Section 5.6).

There are conditions particular to the capture of the tube within the kinetic expansion region that separates the COA within the kinetic expansion from the COA for a defect in the free span.

(Therefore, the subject leakrates calculated for flaws in the kinetic expansion are not usable for flaws in the free span.)

It is arguable whether any COA occurs at all within a kinetic expansion because the tube will not slip or rotate within the expansion. Within any expansion region, the tubesheet, due to its proximity alone, guides the tube and prevents rotation at the elevation of a defect that could result in increasing COA. In addition, remaining contact pressure on the tube OD surface further provides a friction reaction that prevents bending of the tube that could result in increasing COA.

Therefore, for the purpose of leakage assessment from flaws in the kinetic expansions, COA depends on applied axial tension only because there is no rotation at the elevation of a defect due to remotely applied tension. COA is assumed to develop because of asymmetry local to the section as the symmetrically distributed load comes into equilibrium with the asymmetrical section containing the defect.

NUREG/CR-3464 (Reference 13) provides the solution for COA for circumferential defects in OTSG tubes under applied axial tension. The COA for axial defects is also provided. This reference has been widely used in the nuclear industry and, in particular, was the source for COA evaluation for the leak-before-break analysis of RCS piping in B&W plants (Reference 18).

5.5.2 Methodology (Kinetic expansion region)

Reference 13 provided the equations necessary to calculate the crack opening area for circumferential through-wall cracks in tension and axial through-wall cracks subjected to internal pressure. The methodology was implemented in Reference 14 and is summarized herein.

66 ECR # 02-01121, Rev. 2

5.5.2.1 Circumferential Through-Wall Crack in Tension The crack opening area as a function of the axial, tensile, tube load was calculated based on the applied axial stress (a,), Young's Modulus (E) for the tube material, and a non-dimensional function (It (0)) formulated from the stress intensity factors:

It=(,rR2 )I,()

E The applied stress was calculated given the axial tensile load (P) and the mean tube radius (R) with the tube wall thickness (t), or the inner and outer tube radius (Ri and R.,

respectively):

P P 2 z :2_ 2 5.5.2.2 Axial Through-Wall Crack Subjected to Internal Pressure The crack opening area for an axial through-wall crack with internal pressure was calculated based on the membrane stress (a), Young's Modulus (E) for the tube material, mean tube radius (R), tube wall thickness (t), and a non-dimensional function (G(X))

formulated from the stress intensity factors:

A = a(2;rRt)G(A)

E The applied stress was calculated given the differential pressure (p), mean tube radius (R), and tube wall thickness (t):

af=- pR t

This methodology was used to calculate the crack opening area for through-wall cracks of tubes with an outer radius to wall thickness ratio (R/t) of less than or equal to 10.0 with no bending moment applied. The crack opening areas for R/t ratios of less than 10.0 are conservatively large.

67 ECR # 02-01121, Rev. 2

5.6 Crack Area Leakage Analysis 5.6.1 Overview The leakage flow for a given crack area (from Section 5.5) was determined by the PICEP (Pipe Crack Evaluation Program) computer code developed by EPRI (Reference 15). A brief description of the code is provided in this section.

The crack area as a function of time for a given crack length and crack orientation was provided from the analysis described in Section 5.5 above. The T-H parameters were provided in Section 5.3 above. The PICEP code utilizes a crack area, the RCS pressure, RCS temperature, and OTSG pressure at a single point in time and calculates a leak rate through the crack for that specific time. In order to develop a leak rate as a function of time, the code has to be run numerous times throughout the MSLB transient duration. The PICEP analysis was run at the MSLB transient model data intervals. The result was a leak rate as a function of time, which was then integrated to provide a total leakage volume for a given crack. This process was repeated for each type of crack indication at different radial locations within the tube bundle. (See Section 5.7.)

The contact pressure between the expanded tube and the tubesheet causes a significant reduction in leakage. However, the calculations took no credit for the leakpath between the tube and the tubesheet.

5.6.2 Code Description The PICEP program (Reference 15) was used to calculate the crack opening area, the critical crack length and the flow rate through various sizes and types of cracks in kinetic expansions.

Options are available to calculate the leakage with a crack area that is supplied by the user. For subcooled or saturated liquid discharge, the critical flow equations are based on the Henry/Fauske homogeneous non-equilibrium critical flow model with modifications to account for fluid friction due to surface roughness, crack turns, and non-equilibrium 'flashing' mass transfer between liquid and vapor phases. The flow was assumed to be isenthalpic and homogeneous with non-equilibrium effects introduced through a parameter, N, which is a function of equilibrium quality and flow path length-to-diameter ratio, I/D.

The PICEP program was used to estimate calculate the theoretical leakage from the axial, circumferential, and volumetric indications in the TMI-1 kinetic expansions. (As described above, volumetric indications are conservatively assumed to result in both a circumferential crack and an axial crack.) The PICEP program predicts the theoretical flow through straight cracks. The volumetric morphology of the ID IGA flaws, the predominant flaws within the kinetic expansions, is dissimilar to the morphology of straight cracks. However, given the constraint of the tubesheet, it is very conservative to predict leakage based on the assumption that each volumetric flaw will result in one circumferential, throughwall, straight crack and one axial, throughwall, straight crack.

Numerous inputs were required for the PICEP calculations to estimate the leakage from the kinetic expansion flaws:

- Tensile loads on the tube were set to zero for the axial cracks (since tensile loads tend to tighten these cracks and reduce leakage).

68 ECR # 02-01121, Rev. 2

Surface roughness was set to 0.0002 inches, a value of roughness typical for corrosion-induced cracks.

No credit was taken for any tortuosity of the crack channel. (The number of 45 degree turns was set to zero for the computer code runs.)

Minimum tube wall thickness of 0.034" was assumed.

Validation/benchmarking of the PICEP program was based on a large number of flaws and is described in Appendix C of EPRI NP-3596-SR (Reference 15). PICEP crack flow results were assessed using several sets of leak data including data from EPRI (Battelle Columbus and Wyle Laboratory), NRC (UC Berkeley), Canada (AECL), Italy, and Japan. The types of cracks used for this validation work were varied. For example, PICEP results were compared with flow data from cracks formed by parallel plates, pipes with circumferential cracks, and rectangular slits. Among the test results with which PICEP was compared were those results described in NUREG/CR-3475, "Critical Discharge of Initially Subcooled Water Thru Slits". (The PICEP results showed good agreement with the NUREG's results.) Additional work to benchmark the PICEP code is described in EPRI NP-6897-L, "Steam Generator Tube Leakage Experiments and PICEP Correlations" (Reference 33). In that study the PICEP results were benchmarked against numerous steam generator tube laboratory leak tests. (48 leak tests were conducted on I-600 steam generator tube specimens with laboratory-generated flaws.)

5.6.3 (This Section was deleted.)

5.6.4 (This Section was deleted.)

5.6.5 Leakage from Defects Above the Required Kinetic Expansion Length Estimated leakage from flaws that are located above the AKELI.N expansion lengths will be very small in comparison with flaws that are located nearer to the expansion transitions. In classical equations for laminar flow through a small annular orifice formed by concentric members with circular cross sections - a highly idealized representation of the kinetic expansions in which the tubing was expanded, twice, against a drilled tubesheet bore with explosive force - flow is linearly inversely proportional to length of the orifice (Reference 34). Thus, if it was conservatively assumed that a kinetic expansion flaw's leakpath were a concentric annulus, expected leakage from a hypothetical flaw 3.0" into the expansion would be 10% of the expected leakage from an identical flaw located 0.3" into the expansion.

To conservatively account for flaws that may be present above the kinetic expansions' required lengths, where the tubing is not examined, an MSLB-induced leak rate will be assumed. As previously described, the kinetic expansion joints were designed to be "essentially leaktight". The results of the original leak rate testing (Reference 29) for the 69 ECR # 02-01121, Rev.2

kinetic expansions indicated a 99% confidence that 99% of the expanded tubes will have leak rates less than 460 x 0I lb/hr-tube during normal plant operations. A 6" minimum defect-free length was used for the original design leak testing.

For the purposes of outage kinetic expansion evaluations, TMI-I will assume 15 gallons of accident-induced leakage over the first 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> of the MSLB transient, and 170 gallons of accident-induced leakage over the duration of the MSLB transient, from the uninspected lengths of each steam generators' kinetic expansion joints (i.e., deeper into the kinetic expansion joints). These are conservative values that were derived as follows:

- The shortest required kinetic expansion length is 2.1 " (Table 1). This is approximately 1/3 of the original leak tested joint, so the expected leakage from a hypothetical leak into a "concentric annulus", as described above, would be 3 times greater than a leakrate at 6" deep into the expansion under the same conditions.

- The peak MSLB break differential pressure at TMI-I is less than twice the normal plant operations' differential pressure. If a factor of 2 increase in differential pressure is assumed, the expected primary-to-secondary MSLB leakage, based on Bernoulli's theorem, would be increased over the normal plant operations leakage by a factor of the square root of 2. A factor of 2 will be used, which is conservative. (Figure 15 of this report shows the expected primary-to-secondary differential pressure for a MSLB event.)

- Each TMI-I steam generator has less than 15,000 tubes in service.

- Combining these factors yields:

(460E-6 lbs/hr-tube) ( 3 ) (2) (15000 tubes/generator) = 42 lbs/hr-generator A reference density of 0.7094 grams/cc was used for the kinetic expansion leakage evaluations, which is equal to 5.92 lbs/gallon. (Primary side temperature was 579F.

Refer to Section 5.1.) Therefore, to convert this mass flow rate to a volumetric flow rate:

(42 lbs/hr-generator) / (5.92 lbs/gallon) = 7.1 gals /hr- generator Accident-induced leakrates are tabulated in this document on a volume basis over a 2-hour period and over the duration of a hypothetical MSLB, as described in Section 5.1.

(The duration was calculated to be 23.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br />.) Converting this leakrate to provide consistent units with these other calculated leakrates:

2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> leakage = (7.1 gals/hr-generator) (2 hrs.) = 15 gallons/generator Duration Leakage = (7.1 gals/hr-generator) (23.5 hrs.) = 170 gallons/generator Note that primary-to-secondary leakage from the TMI-I kinetic expansions has been less-than-detectable over the past several operating cycles, so there is no evidence that the kinetic expansion joints leak during normal plant operations.

70 ECR#02-01121,Rev.2 I

5.7 Total Leakage Evaluation 5.7.1 Overview This section describes the approach taken to determine the total leakage for the purposes of comparison against the leakage limits. A calculation methodology was developed that integrates the OTSG tube loads with the thermal hydraulic data and analysis needed for leakage through the cracks and combines the results into leakage assessment tables.

These calculated leakages are based on implementing the methodology discussed in Sections 5.3 through 5.6 above. Also discussed in this section are the ways in which the unaffected OTSG will be treated since the tube loads are quite different (i.e., smaller) and the steamline is intact.

5.7.2 Leakage Results Calculations were created to apply the methodology discussed in earlier sections of this report to calculate the leakrates from postulated tube cracks in the kinetic expansions of the OTSGs(Reference 21). The crack opening area was calculated based on the tube tensile load or the differential pressure depending on the orientation of the crack. The mass flux was calculated using the PICEP computer program given the crack geometry and the fluid properties as discussed in Section 5.6. The mass flux was converted to a volumetric leakrate based on a reference density (579 degrees F and 2200 psi) and the crack opening area. (This reference density corresponds to the same value as was used in determining the FSAR leakage limits.) The calculated leakage from cracks of various sizes was integrated over a period of 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and for the duration of the MSLB transient.

The results of this calculation can be provided by 'binning' of integrated leakage from cracks in the range of sizes for circumferential and axial leakage. The circumferential crack size bins for a given radial position in the OTSG are the same, but the integrated leakage for a given crack size is different as a function of radial position. This is necessary for circumferential crack leakage-- but is not necessary for axial crack leakage which is not sensitive to radial position, only differential pressure.

The circumferential crack integrated leakage results, presented as leakage tables according to crack size for 5 concentric, radial "zones" (from the center of the tube bundle to the periphery), are provided in Table 4. For axial cracks, the leakage is provided as crack size bins in Table 4. The bins for all of the circumferential crack tables range from 0.05 inch crack size (.05 inch leakage is used for all cracks from 0.02 to 0.05 inches) through 0.65 inches. Table 4 also provides the leakage calculation results for axial indications up to I inch in length. In the field all circumferential and axial extents are 'rounded up' to the next 0.05 inch increment. [Note that the circumferential crack integrated leakage 5 bins are slightly different than the 11 bins of the original version of this document. Reference 26 originally placed the results into 11 bins. One of those I I bins was eliminated since it was for the very center of the steam generator (radius = 0") and there are no tubes at the center of the generators. The remaining 10 bins were combined into 5 bins.]

As previously described, if an indication is determined to be volumetric, it is treated as two cracks. Each volumetric indication is treated independently as if there were one axial and one circumferential crack of lengths equal to the volumetric flaw's measured axial and circumferential extent, respectively. It is very conservative to estimate the 71 ECR # 02-01121, Rev. 2

theoretical leakage from volumetric flaws in the kinetically expanded tubing by considering them as a combination of a 100% throughwalI circumferential crack of length equal to the as-called circumferential extent of the volumetric flaw and a 100%

throughwall axial crack of length equal to the as-called axial extent of the volumetric flaw. This treatment of the volumetric flaws is conservative for a number of reasons including:

- the fact that the tubing is expanded into the tubesheet and is unlikely to crack axially. (Expansion and deformation of the tube in the hoop direction are prevented by the constraint of the tubesheet.)

- pulled tube examination results from TMI-I have demonstrated that the MRPC examinations tend to overestimate the extents of the ID volumetric IGA flaws (as a result of the "look-ahead/look behind" effect and the proximity of the ID flaws to the surface-riding coils). The majority of flaws within the kinetic expansions are ID volumetric IGA flaws, as is also the case for the freespan tubing in the TMI-I steam generators.

- bending of the tubing is prevented by the presence of the tubesheet. (Crack formation is less likely since movement/displacement of the tubing is severely restricted.).

- the presence of the tubesheet prevents formation of a volumetric "hole"; thus only a tortuous flow path through an intergranular flaw surface (similar to a crack) would be expected.

5.7.3 Affected OTSG Versus Unaffected OTSG Since both the affected OTSG and the unaffected OTSG will experience tube loads, leakage is possible from both generators. Since either of the two OTSGs might be the affected one, it is necessary to assume that the OTSG with the greatest volume of estimated leakage is the affected generator.

The leakage from each of the indications has to be summed, and the total leakage for the OTSG can then be compared against the total leakage limits of 3228 and 9960 gallons (at 579 degrees F, 2200 psia) for the 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> EAB and 30 day LPZ, respectively, discussed in Section 5.1. Since OTSG tube loads were not specifically determined for the unaffected OTSG, it is necessary (and conservative) to treat the unaffected generator as if it had the same loads as the affected generator. Thus, the same process used for the affected OTSG will be used for the unaffected OTSG. The leakage calculations assume that either steam generator could leak (as if it were the affected generator during an MSLB) and determine the leakage based on the sum of the cracks in that generator without taking credit for the intact steamline of an unaffected generator.

The estimated leakage from kinetic expansions is calculated for each of the steam generators based on outage inspection results. Since either of the TMI-I steam generators could have been the affected OTSG during a hypothetical MSLB that occurred in the operating cycle prior to the inspection, it is necessary that each of the OTSGs has an "as-found" estimated leakage less than the above leakage limits. Since either of the TMI-I steam generators could be the affected OTSG during a hypothetical MSLB that occurs during the operating cycle following the inspection and required tube 72 ECR#02-01l121,Rev. 2

repairs, it is necessary that each of the OTSGs has an "as-left" estimated leakage less than the above leakage limits. (Note that estimated leakage from flaws in the steam generator tubing located in areas other than the kinetic expansions, possible leakage from other tubing repairs, and possible primary-to-secondary leakage during the operating cycle must also be considered in this evaluation of possible leakage versus the steam generator performance criteria limits.)

5.8 Leakage Assessment Methodology Summary The leakage assessment methodology allows for a determination of the leakage that may occur during a Main Steam Line Break (MSLB) event from conservatively assumed through-wall cracks in the kinetic expansions in the upper tubesheets. Eddy current indications with throughwall estimates greater than 67% are assumed to be 100%

through-wall cracks that will leak during the MSLB.

The amount of leakage is determined by calculating the leakage area resulting from the MSLB-induced tube loads (differential pressure only for axial cracks), and then calculating the subsequent leakage flow rate and total event integrated leakage for each applicable indication based upon the thermal hydraulic conditions associated with the MSLB event. The estimated leakage for all cracks is compared against 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> and event duration leakage limits. These leakage limits for the TMI-I steam generators ensure that exclusion area boundary and 30 day low population zone doses do not exceed a small fraction of 10 CFR 100 requirements if the MSLB event were presumed to occur.

The implementation of this leakage assessment methodology using OTSG eddy current data provides reasonable assurance that the leakage that could occur during a design basis MSLB from indicated cracks in the kinetic expansion region may be conservatively determined.

5.9 Reporting Requirements Kinetic expansion inspection results will be reported to the NRC. These results will include the number of tubes plugged, the types of degradation detected, the radial location and required expansion lengths of tubes with degradation, results of growth assessments, and the calculated theoretical MSLB-induced leakage from kinetic expansion indications. The following is a list of the information to be reported to the NRC, including the method by which it will be reported, and the time period in which it will be reported.

73 ECR # 02-0112 1, Rev. 2

ITEM TO BE REPORTED TYPE OF REPORT DATE DUE TO NRC Number of kinetic expansions Written report. This information Shall be reported to the NRC within inspected. will be reported with the 90-day 90 days following completion of the report currently required by TMI- I inspection and repairs (main Tech. Spec. 4.19.5 (b). generator breaker closure).

[Same as the 90-day report currently required by TMI- I Tech. Spec.

4.19.5 (b).]

Location, percent of wall-thickness Written report. This information Shall be reported to the NRC within penetration, voltage, and axial will be reported with the 90-day 90 days following completion of the and/or circumferential extent for report currently required by TMI- I inspection and repairs (main each kinetic expansion indication. Tech. Spec. 4.19.5 (b). generator breaker closure).

[Same as the 90-day report currently required by TMI- I Tech. Spec.

4.19.5 (b).J Tubesheet radius location and Written report. This information Shall be reported to the NRC within minimum defect-free kinetic will be reported with the 90-day 90 days following completion of the expansion length required report currently required by TMI- I inspection and repairs (main (AKELI,,N) associated with each tube Tech. Spec. 4.19.5 (b). generator breaker closure).

with degradation detected in its [Same as the 90-day report currently required kinetic expansion region. required by TMI- I Tech. Spec.

4.19.5 (b).]

Number of tubes plugged due to Written report. This information Shall be reported to the NRC within kinetic expansion indications. will be reported with the 90-day 90 days following completion of the report currently required by TMI- I inspection and repairs (main Tech. Spec. 4.19.5 (b). generator breaker closure).

[Same as the 90-day report currently required by TMI- I Tech. Spec.

4.19.5 (b).]

An assessment of the growth of If no growth is detected: This Shall be reported to the NRC within indications within the kinetic information will be reported with the 90 days following completion of the expansions in accordance with Sect. 90-day report currently required by inspection and repairs (main 3.2 of this report, including the TMI- I Tech. Spec. 4.19.5 (b). generator breaker closure).

number of tubes with new [Same as the 90-day report currently indications located in the required required by TMI- I Tech. Spec.

kinetic expansion region. 4.19.5 (b).]

If growth is detected: NRC shall be Telephone call to NRC shall be notified by telephone during the during the outage in which growth is outage in which growth is detected. detected. Report(s) required by Additional notifications/reports, 10CFR50.72 and IOCFR50.73, if shall be made in accordance with the applicable, shall be made in requirements of IOCFR50.72 and accordance with schedule prescribed IOCFR50.73. if applicable. in those documents.

An assessment of the theoretical If as-found leakage is projected to be Shall be reported to the NRC within MSLB-induced leakage from less than 3228 gals for 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> 90 days following completion of the indications within the kinetic duration leakage and less than 9960 inspection and repairs (main expansions in accordance with Sects. gals. over the MSLB duration: generator breaker closure).

5.7 and 5.8 of this report Written report. This information will [Same as the 90-day report currently be reported with the 90-day report required by TMI- I Tech. Spec.

currently required by TMI- I Tech. 4.19.5 (b).]

Spec. 4.19.5 (b).

74 ECR # 02-01121, Rev. 2 l

ITEM TO BE REPORTED TYPE OF REPORT DATE DUE TO NRC If as-found leakage is projected to be Telephone call to NRC shall be greater than 3228 gals for 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> during the outage in which said duration leakage or greater than leakage is determined. Report(s) 9960 gals. over the MSLB duration: required by IOCFR50.72 and NRC shall be notified by telephone IOCFR50.73, if applicable, shall be during the outage in which said made in accordance with schedule leakage is determined. Additional prescribed in those documents.

notifications/reports, shall be made in accordance with the requirements of IOCFR50.72 and IOCFR50.73, if applicable.

6.0 INSPECTION CRITERIA AND LEAKAGE ASSESSMENT

SUMMARY

Kinetic expansions were installed in the upper tubesheet region of more than 30,000 TMI-1 steam generator tubes in the early 1980's. Finite element analysis modeling has demonstrated that the kinetic expansions are relatively flaw tolerant. These expansions are protected from a number of types of stresses, vibrations, bending, and secondary-side loose parts by the presence of 24" thick tubesheets.

Eddy current inspections of the TMI-l kinetic expansions are required by the plant's steam generator program. This document provides the required inspection scope, reporting requirements, leakage assessment methodology, and acceptance criteria that conservatively disposition kinetic expansion inspection results. Kinetic expansions that contain flaws that might be adversely influenced by MSLB-induced stresses are removed from service under the subject conservative criteria. This document also requires a conservative evaluation of the estimated leakage that might occur from flaws detected within the kinetic expansions.

The criteria require, beginning with the Fall 2005 refueling outage, that only ID volumetric indications may remain in service in the kinetic expansions' required lengths, that new indications be removed from service, and that 100% of the in-service kinetic expansions be examined each refueling outage. These and other conservatisms implement conservative criteria with which to disposition the kinetic expansions during each examination.

75 ECR # 02-01121, Rev. 2 I

7.0 REFERENCES

1. Letter from James W. Langenbach to USNRC, "Once-Through Steam Generator Kinetic Expansion Inspection Criteria", August 08, 1997 (6710-97-2348).

2. GPUN Calculation No. C-1 101-224-6612-057, "Offsite Doses from OTSG Tube Leakage Due to Main Steam Line Break", August, 1997.

3. Letter from James W. Langenbach to U.S.N.R.C., "TMI-1 License Amendment Request No. 269 'Revised Steam Line Break Accident Analysis Dose Consequence"', August 14, 1997 (6710-97-2345).

4. GPUN Calculation No. C-I 101 -224-E510-061, "MSLB Analysis for OTSG Tube Integrity", August 21, 1997.

5. Letter, USNRC to Mr. C. R. Lehman, "Acceptance for Referencing of the RETRAN-02 MOD005.1 Code", April 12, 1994.

6. GPUN Calculation, C-I 101-202-5412-114, Rev. 2, "TMI-1 RETRAN Base Deck".

7. Weimer, J. A., "Thermal Mixing in the Lower Plenum and Core of a PWR", EPRI NP-3545, May, 1984.

8. GPUN Calculation C-I101-224-E610-060, "Long Term MSLB Transient Data for Tube Leakage Calculations", September, 1997.

9. GPUN Procedure ATP 1210-10, "Abnormal Transients Rules, Guides and Graphs," Rev.

32.

10. GPUN Procedure ATP 1210-3, "Excessive Primary to Secondary Heat Transfer," Rev.

20.

11. GPUN Calculation C- 1101-224-E610-058, "Post MSLB Cooldown of a Voided OTSG",

Rev. 0, September 1997.

12. GPUN Procedure OP 1104-6, "Reactor Coolant Pump Operation", Rev. 62.

13. NUREG/CR3464, "The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulating Circumferential Through-all Cracks",

September, 1983.

14. GPUN Calculation C- 1l01-224-E610-054, Rev. 1, "Tube Crack Opening Area Calculation Methodology", August 14, 1997.

15. EPRI NP-3596-SR, Revision 1, Special Report, "PICEP: Pipe Crack Evaluation Program (Revision 1)", December, 1987.

16. GPUN Calculation C-1101-224-E520-061, Rev. 0, "Methodology for Calculating OTSG Tube Axial Loads", September 30, 1997.

76 ECR#02-01121,Rev.2

17. Babcock and Wilcox Report No. BAW 10146, "Determination of Minimum Required Tube Wall Thickness for 177-FA Once-Through Steam Generators", October, 1980.

18. Babcock & Wilcox Report No. BAW-1847, "Leak-Before-Break Evaluation of Margins Against Full Break for RCS Primary Piping of B&W Design NSS", Rev. 1, 77-1153295-01, September, 1985.

19. "Analysis of Remaining Contact Pressure for TMI OTSG Kinetically Expanded Tubes for Leakage Assessment Purposes", enclosure to letter from H. W. McCurdy (MPR) to S.

D. Leshnoff (GPUN), September 29, 1997, "Three Mile Island Generating Station Development of OTSG Kinetic Expansion Inspection Acceptance Criteria."

20. Framatome Technologies, Inc., Report #51 -1264463-00,"TMI- I EDM Notch Sample Hot/Cold Leak Tests Results", August 15, 1997.

21. GPUN Calculation C- 1101-224-E610-059, Rev. 0, "OTSG Tube Leakage Methodology for Tubesheet Region", September, 1997.

22. Framatome, Technologies, Inc., Report #51-5000542-00, "GPU MSLB Tube Load Summary", September 19, 1997.

23. GPUN Calculation C-i 10I-224-E520-063, Rev. 0, `TMI OTSG Analysis of Tube Axial Loads During Plant Specific MSLB".

24. MPR Report #MPR-1 820, Revision 1,"Three Mile Island Nuclear Generating Station OTSG Kinetic Expansion Inspection Criteria Analysis", MPR Associates, Inc.,

September, 1997.

25. GPU Nuclear Letter #6710-97-2348, J.W. Langenbach to U.S. Nuclear Regulatory Commission, "Once-Through Steam Generator Kinetic Expansion Inspection Acceptance Criteria", August 8, 1997.

26. GPU Nuclear Letter #6710-97-2441, J.W. Langenbach to U.S. Nuclear Regulatory Commission, "Leakage Assessment Methodology for TMI- I Once Through Steam Generator (OTSG) Kinetic Expansions", November 26, 1997.

27. TMI-l UFSAR, Section 14.1.2.9.

28. NUREG 1019, "USNRC SER Related to Steam Generator Tube Repair and Return to Operation", TMI-1, November 1983.

29. Babcock & Wilcox Report #BAW-1760, Rev. 1, March 1983.

30. AmerGen Letter #5928-01-20169, M. E. Warner to U.S.N.R.C, "Additional Information

- License Change Application No. 291 - Once Through Steam Generator (OTSG)

Surveillance Specifications Applicability Following Cycle 13", July 13, 2001.

31. U.S. Nuclear Regulatory Commission Letter, T. G. Colburn to M. E. Warner, "Three Mile Island Nuclear Station, Unit I (TMI-1) Surveillance Following Cycle 13 (TAC No.

MB0664)", October 5, 2001.

77 ECR#02-01121,Rev.2

32. AmerGen Letter #5928-02-20036, "Cycle 14 Refueling (T IR14) Inservice Inspection (ISI) Summary Report", M. P. Gallagher to U. S. Nuclear Regulatory Commission, March 5, 2002.

33. EPRI NP-6897-L, "Steam Generator Tube Leakage Experiments and PICEP Correlations".

34. Keller, George, "Hydraulic Systems Analysis", published by the Editors of Hydraulics and Pneumatics Magazine, Penton/IPC, Cleveland, 1978.

35. AmerGen Letter #1920-99-20679, J. B. Cotton to US NRC "Cycle 13 Refueling (13R)

Inservice Inspection (ISI) - ASME NIS I &2 Owner's Data Report Forms with Reports of the Once Through Steam Generator (OTSG) Tube Inspections, Pressure Tests, and ASME Section XI, Subsection IWE & IWL containment Inspections", January 14, 2000.

36. GPU Nuclear Letter #6L20-98-20004, J. W. Langenbach to US NRC, "Cycle 12 Refueling (12R) Outage Once Through Steam Generator (OTSG) Tube Inspection Report with ASME NIS Data Reports for Inservice Inspections (ISI)", January 12, 1998.

37. AmerGen Letter #5928-04-20063, M. P. Gallagher to USNRC, "Cycle 15 Refueling (TlRI5) Inservice Inspection (ISI) Summary Report", February 24, 2004.

38. EPRI Pressurized Water Reactor Steam Generator Examination Guidelines, EPRI TR

1. 1003138, (Revision 6, including interim guidance).

39. Natrella, Mary G., "Experimental Statistics, National Bureau of Standards Handbook 91 "; Issued August 1, 1963, reprinted October 1966 with corrrections.

78 ECR#02-01121,Rev.2

TABLE 1 INSPECTION ACCEPTANCE CRITERIA FOR OTSG KINETIC EXPANSION REGION (REQUIRED EXPANSION LENGTH)

Minimum Defect-Free Kinetic Kinetic Expansion Length Bundle Expansion Length Required BundleAKEL,,,,

0.00" - 20.00" 3.4" 20.01" -42.00" 3.2" 17 42.01"- 46.00" 3.0" 46.01" - 50.00" 2.7" 50.01" - 55.00" 2.4"

> 55.00" 2.1" 0.00" - 20.00" 8.4" 20.01" - 42.00" 8.2" 22" 42.01" - 47.00" 8.0" 47.01" - 50.70" 5.2" 50.71"9- 54.30" 4.2"

>54.30" 3.2" 79 ECR # 02-01121, Rev.2 l

TABLE 2 INSPECTION ACCEPTANCE CRITERIA FOR OTSG KINETIC-EXPANSION REGION (FLAW DISPOSITIONING CRITERIA)

Defect Type (Note Iland Note 2) Requirement(s)

The AKELNIN length (Table 1) of defect-free tubing must be present.

For multiple defects, I/4-inch shall be added to the length Axial of each defect, except the first defect. Also, for each circumferential defect, a defect length of 'A-inch shall be added. Example: Three axial defects are found, with one defect I-inch long and two defects each '%-inch long.

In addition, two circumferential defects are found. The effective length of the %-inch defects is: % inch + 1/4 inch = 3/4 inch. The combined length of the three axial defects is: 1-inch + 3/4-inch + 3/4-inch = 2 1/2-inch. The effective axial influence of the two circumferential defects is: 2/4-inch + 1/4-inch = 1/2-inch.The total length of axial influence is 2 1/2-inches+ 1/2-inch= 3inches.

The AKELNI,.4 length (Table 1) of defect-free tubing must be present.

For single defects, no defect may be longer than 130 Circumferential degrees or 0.64 inches. For multiple defects:

• If separated axially by less than I-inch, their length shall be combined, and the total shall be less than 0.64-inch.

• If separated axially be more than 1-inch, the individual defects shall each be less than 0.64-inch in extent.

NOTES:

1. For volumetric defects, the criteria for axial defects shall be used for the axial length of any volumetric defect, and the criteria for circumferential defects used for the circumferential length of any volumetric defect.

2. Note that flaws other than ID volumetric IGA in the kinetic expansions' required lengths are removed from service under this criteria. (Only ID volumetric flaws may remain in this area, provided they meet the requirements of this table and steam generator projected total leakage required by this report is not exceeded.) This table is used to disposition the axial extents and circumferential extents of the ID volumetric IGA defects. This table is also used for condition monitoring of axial or circumferential defects in the kinetic expansions' required lengths, prior to their removal from service.

80 ECR # 02-0112 1,Rev. 2 l

TABLE 3

[This table was deleted.]

81 ECR # 02-01121, Rev. 2 l

Table 4 Leakage Assessment Evaluation Data CIRCUMFERENTIAL INDICATIONS:

Theoretical MSLB Leakage Based on Circumferential Extent Tubesheet Circ. Extent 2 Hour Leakage Duration Radius Location (Inches) (gal) Leakage of Tube (gal)

(inches) 0 - 0.01 0 0 0.02 - 0.05 0 0.05 0.06 - 0.10 0.03 0.28 0.11 - 0.15 0.08 0.85 0.16 - 0.20 0.18 1.93 0.21 - 0.25 0.35 3.77 0.0 - 11.525 0.26 - 0.30 0.61 6.66 0.31 - 0.35 1.01 11 0.36 - 0.40 1.61 18.21 0.41 - 0.45 2.63 29.93 0.46 - 0.50 4.1 47.04 0.51 - 0.55 6.21 71.34 0.56 - 0.60 9.14 105.1 0.61 - 0.65 13.17 151.16 0- 0.01 0 0 0.02 - 0.05 0.01 0.06 0.06 - 0.10 0.04 0.35 0.11 - 0.15 0.11 1.07 0.16 - 0.20 0.24 2.42 0.21 - 0.25 0.46 4.7 11.526 - 23.05 0.26 - 0.30 0.81 8.3 0.31 - 0.35 1.34 13.98 0.36 - 0.40 2.28 24.02 0.41 - 0.45 3.66 38.99 0.46 - 0.50 5.64 60.54 0.51 - 0.55 8.42 90.73 0.56 - 0.60 12.25 132.12 0.61 - 0.65 17.43 187.93 82 ECR # 02-01121, Rev. 2 l

Table 4 (Cont'd)

CIRCUMFERENTIAL INDICATIONS:

Tubesheet Circ. Extent 2 Hour Leakage Duration Radius Location (Inches) (gal) Leakage of Tube (gal)

(inches) 0 - 0.01 0 0 0.02 - 0.05 0.01 0.08 0.06 - 0.10 0.05 0.47 0.11 - 0.15 0.15 1.42 0.16 - 0.20 0.34 3.22 0.21 - 0.25 0.65 6.25 23.051 - 34.575 0.26 - 0.30 1.15 11.08 0.31 - 0.35 2.05 20.02 0.36 - 0.40 3.41 33.81 0.41 - 0.45 5.4 53.99 0.46 - 0.50 8.19 82.54 0.51 - 0.55 12.04 121.84 0.56 - 0.60 17.25 174.88 0.61 - 0.65 24.18 245.4 0 - 0.01 0 0 0.02 - 0.05 0.01 0.11 0.06 - 0.10 0.07 0.63 0.11 - 0.15 0.21 1.9 0.16 - 0.20 0.47 4.3 0.21 - 0.25 0.92 8.35 34.576 - 46.1 0.26 - 0.30 1.74 15.93 0.31 - 0.35 3.05 28.43 0.36 - 0.40 4.99 47.19 0.41 - 0.45 7.76 74.12 0.46 - 0.50 11.6 111.51 0.51 - 0.55 16.79 162.14 0.56 - 0.60 23.69 229.5 0.61 - 0.65 32.78 318.02 0 - 0.01 0 0 0.02 - 0.05 0.02 0.14 0.06 - 0.10 0.1 0.84 0.11 - 0.15 0.29 2.5 0.16 - 0.20 0.64 5.65 0.21 - 0.25 1.32 11.51 46.101 - 57.625 0.26 - 0.30 2.5 22.24 0.31 - 0.35 4.31 38.98 0.36 - 0.40 6.94 63.64 0.41 - 0.45 10.64 98.41 0.46 - 0.50 15.67 145.89 0.51 - 0.55 22.37 209.32 0.56 - 0.60 31.18 292.75 0.61 - 0.65 42.65 401.91 83 ECR # 02-01121, Rev. 2 l

Table 4 (Cont'd)

AXIAL INDICATIONS:

Theoretical MSLB Leakage Based On Axial Extent Axial Extent 2 Hour Leakage Duration Leakage (Inches) (gal) (gal) 0 -0.01 0 0 0.02 - 0.05 0.01 0.02 0.06 - 0.10 0.04 0.13 0.11 - 0.15 0.12 0.45 0.16 - 0.20 0.31 1.19 0.21 - 0.25 0.7 2.73 0.26 - 0.30 1.53 5.72 0.31 - 0.35 3.14 11.21 0.36 - 0.40 5.81 20.49 0.41 - 0.45 9.87 36.51 0.46 - 0.50 15.64 61.2 0.51 - 0.55 23.45 96.42 0.56 - 0.60 33.61 144.31 0.61 - 0.65 46.45 206.92 0.66 - 0.70 62.33 286.28 0.71 - 0.75 81.64 384.43 0.76 - 0.80 104.81 503.54 0.81 - 0.85 132.33 646.46 0.86 - 0.90 164.68 815.41 0.91 - 1.00 245.97 1238.97 I

84 ECR#02-01121,Rev.2 I

TM!-1 IN-SITU PRESSURE TEST LIST OTSG-A TrUBE AND EDDY CURRtENT IN FOROATION IN-SITU TF.ST RESUL7S Tubelnfori 1atin i

_lln riot__ l Dobbin Dali Dnut oPM GrM 0 GPON Max Rtgion Row Tube 1Acat Ltngth (in.) Volts Est% Orlenutlon Volts Esta. C""meits NOrD MSLW RG. 1.121 hressurv Tue Region~N LofonT Un, 93 119 urs + 1.50 0.25C 17.135 s7 IDSCI 2.39 93%s Uper Tubeshect Circ 0 0 0 4400 Tubshet 93 119 UTS+i.93 C.2lAxrO.1 C 9.2 N/A ID Vol. 2.39 93% UTS Volumetric 0 0 0 4400

. 93 11 _ UfS +0.55 0.24AX o.27C 3169 __N/A ID VOL 0.95 4t0tS urs Vohmdric 0 4400 107 120 E J-0.49 0.27A x 0.25C 2.36 IVA ID VOL NDD NIA Kb-r Volumetrie I O _O0 440 0 107 320 E`TL-25 2 Ql2Axo.9c 3.21 N/A ID VOL 1.47 27% ursYohmtetre 0 o0 0 4400 307 20 Fn-W3.9S Q1 6Ax0Q16C 2.42 N/A ID .444OL .33% UTS Vohrmetrie I 0 0 0 4400

(:0 (A

00 Is I-A TMI-t IN.SITU PRESSURE TEST LIST' M 1_

OTSC-B (-9 t

__*_TUBE AND EDDY CURIEN'r INFORMATION l IN-SITU TEST RESULTS Tube Information _ ln Pointn Bobn Data GPM 1 GPMO ( CPM @ Max Regelm Tube LocAtion Lh (in) Vohs Esn.% Orietation Volts ESmlI Comnients NOPD MSLO 3-O. 1.121 Presure

_ _ _ _ ~~ . _ _ _ _ I_ Iw I_ I 1W I__ _ _ _ _ _ _

4400

.. 5 _. UTS-0.7 0.29C 6.99 NA l IDSCI 4.28 l 7% I rslcbe 0 0 0 Uer jj134 1 IS -020 l 014Ax 0.30C 6.04 N/A ID VOL 2.8l 60'h UTSF Vohnuetric 0 0 0 4400 Tubetheet II$1 UTS+174 0.51C 5.11 WIA IDSCI 9.S2 43% l l rst C 0 0 0 4400 FReespan 79 5 5+ 40.30 0 86Ax0.50C 3.39 N/A ODVOL l.3 I3' Fre1sp3%NQ 0 0 0 4400 Freespan 79 60 15+44.57 0169Ax 0.39C .3.5 INA I O1D OVOL 4.77 2 Fre2pnNQ 0 0 0 4400<W

~0

";a To 0

IN SITU PRESSURE AND LEAK TEST RESULTS ~1 TMI-1 SG B 09/99 13R 0 It Is 02 M

9 Co (A IT' TUBE AND EDDY CURRENT INFORMATION l IN-SITU TEST RESULTS 00 C71 so 1

RCto Region bL Row Tube Information

-IUpperTS 6 19 l cLcatlon T 0 .3 1 Langthl Axial ttio.fh 0.31 Plus Point Data 4.517 94 ID SCI IComments NDD Bobbin Data Clrc Volts I Est. Sl Orientation Ind Volts Est. %

UTS Cirt PM NOmet GP6GP 0

IJISLB 0

GxNPM Maximum M@ftGPM~ Pressiure 350 40 CD It 0-3 (A SGi Freespan

[ 80 50 2

15S 29 051o 433.7 I1 046 OSAI NOI 0.3 47 FreespanAx 0 0 4350 \0 M

rA 1 ___ 1131 2 14S 27.94 to 29. 1 1 0 431 OD SAI NOl 06(7 67 Freespan Ax D0 I4350 to

• Note: An additional axial load was applied duringt Ihs test to Impart 1402 tbs. axial tensile load on this indication. This indication was also tested at S0Wpsi with a 2350 lbs. axial load us applied (to simnulale SB LOCA) with no leakage.

n 4t 0

t,3 iz CD

Table 5 (Continued)

In Situ Pressure Test Data Summaries Outage 14R (2001) 188 Si~tu -Test -i-s-t an-dR-e-s-u-lts: OTSG ,-A CUEANO EDDY CURRENT INFORJATION _______IN4I171 TEST R~slr REGION3A¶25 _____PtUS PON. DATA 30wsw AT OF a 050 070GPM a51 M'J y

___RWiCI W-A11CN AS111~INVOLTS MT71 OR1ONTATION VMI7MTSIST% 0.04140117 NHOP IIL IP 64*Osd PRESSUREO (SW0.

fteM56 72 072.-13.07 0.10 0.0 0.12 56 00 mN14-t "CA 022 55 jcC81k=C 0 0 L,0 - 412 17 Lo 52 de IITL .143 [017 2.14 97 1CM ah m1 - rnc

- 0 I GM 1 m-1 0%08 6 10C Pc_mrs11t ID10.4A5C 0 0 0 001 GM5 17-U-e? 2 5 413. 4112 .82 64 I0CI NR 07I44l aC 0 0 640 7 M 0.18s04 c2 C 0 0 a

• S 0.1127 Ot 0.14 0.30 NM 0 a a * £154 us 13.12II9CM0.elm Oilu 3t40 0.4 o0 0 30 402F-&016 A 035 __ o 0 0-W 100 L9: 0.M 030 __04 p4 0 0 0

• 45 135 1.462 GU3 ,022 015 30 0 04?

a- 435 13tu 1 0150702 016 034 p0 NOD0a 4354 US .1405 .390220.15 21 1 2 a0? a 4374 1 1u 148.0 8204 -102250.35 N04O 2 0 0 0 435*

11 11.1 0.2* 0.2 0.10 Svc au 0 440 U3S 02 C400 0.16 0.15 *4' 025 1 a 0 0

• 4351 135 $1240 024- 0.12 0.31 Z; 0 0 a

• 135 1AS .245 0.18 0.18 0.10 lo1 0.38 0230 40.

143-.2359 0.3 .1 0.31 XX 0A6 17 0 0 0 44 145 M024 0.13 0.12 0.10 MOO 0-4 0 0 . 40 1158 $17.62 0.13 ml7 0.31 14 0.40 0 0 .0 - £5 ftewi 13 2 :95.1ISM 0.13 0.17 O.2 __ 1Vdmoot1 I40 0OM 82DIWO0C 0 0 0 - 40. 17 152 .2134 0.13 023 0.34 14R 0.51 0 0 0 - 4400 Ass .24.70 CM0. 0.T2 0214 PWM a a a 4500 PVM .17.24 000 all 0am11. 0 a a 4400 UTZ .13 35 006 0.12 0.22 - 14 0.20 0 0 0 . 4400 UTII 11.75 OM1 0.12 0.M NOD a 0 0

• 4450 U~TZ -lId8 0.18 0.17T 0.12 _ _0 8 00 0 0 - 43

- UTS .9669 00 0.12 0.L2' 0 0 0 . .....

07440.40 0.17 0.14 03 170 0 0 - 40 Fr.oupa 112 as 79 2 .12 Q7 0.13 __ 9Vd~.mlfo NW10 IC."CC 0 ~ 0 I. 44MI I7 110 *1918 0.1l 0.17 01cl* 02 0 O0 4M 125 .7.70 0.1IS 0.15 0.22 'lD0 a 0 . 4*50 129 .0.52 QIII 0.15 0.25 Pdo a 0 a £450 1295411!8 CM~ 0.1501 NWO 0 0 0 . 4400 FftfM15 127 125- .t1lo 032 0.2 0S.3C5ld0 135 -. I1.31 0.11 015 018i im a.5 0 a 4 TM5 .1t41 O.1l 0.11 0.37 NWD 0 0 0 . 4.40 15S .435 0.1S 0.1? 0.40 SVC9* 0.28 0 0 0 "£40 lO 1167 0.20 0.17 1.? NM ~ 0

a. 4ADO Ul.T 2 I124 071 .2.16 0.32 0.37 CV' ID4 me1401 i 82055C 452 1

_mI __ O" 82104*5CC I 0_ 0 0 640 7 87 ECR #02-0112 1, Rev. 2 I

Table 5 (Continued)

In Situ Pressure Test Data Summaries

.Outage 14R (2001)

In Situ Test Tist and Results: OTSG-B UBE AND EDDY CURRENT INIORMAT O1~N

'______ IN-SITU TEST RESULTS RE-0lc.4 TIMEOFOR11ATIi PLUS P011T DATA______ AFTR~ACAMI G1P14 wGOM wmat M"3MU-1 .Pj OC~CH~ t VOLrS1 E0371OR.ENTWATIO CIE4 L-I VOTI E1 % NCCPA #40Wt. 3NOOP PR#ESSURd O.'C9)

• ErI .027 06 0-2* I Os I00M 0 0 0 O *00 UppfrTS 8OSdeL-L .0.38 0.30 0.19 095 __ VVkoret ID 1400 to IMSCC a0 0 0 4400 17 E" OW 0.29 0.93 _____14C ___I_ 0 0 4400 063 *0,04 0.11 0.10 0.29 MN 1129 0 0 0 4400 1063 .18,77 0.11- -0.10 016O MR _028 a 0 0 4400 083 .21 45 0.16 010 10C2 EOD 0 0 0 4100 085 .14.7a 0.11 0.15 017 IN 029 0 0 0 14100 Rd"a 4 09S .13.23 0.16 010 013 ID WC 0.25 ID CA= 0

_40___

01S .7.25 0.18 0.18 013 MV 0.31 0 0 a 440 1 09S *1.74 0.10 0.15 0.20 NOD 0 0 0 4.400 103 .7.91 0.14 10.12 OM3 GY 090 0 a 4.400 JiS .11.92 0.10 0.10 0.28 .3 OY 0 0 0 4400 IIS .e.4? 0.16 0.16 0.23 ~ ~ 032 ___ 0 0 0 *400 ur __903 O_ elNO 0 0 0 40 U___8 ad 0.17 32 M A __1- 2 0 0 0 4W UTS-4._ 428?xi 82 I NC 017 8 ~~lfcIa 0 0 02 17M

__0 "A 3 (183,91401 41 - I063 6? nug* o 0 0 3 17 Fitia  :'11 14 570- 3 02 OOA*Is NO 10.44 OCoo~eIGA a 0 O0 44021 17 4 0 0 14.S .45 SS .O74 __ 021 N0MI0 I I __ 0 Oj 4 Frp mn80 31 UT34026 02

_ 125050o 0I 0r. Na 1.01 _ ID MAJ0C 0 0 0 4350 1 AS _12

,2 _____3 W 0.31 0 0 0 4100 ISS .232 0.16 (15 0.17 BYC 0.55 0- 0 -0 4*00 153 .0.71 0.l8 018 0.40 goo a044 20 0 0 0 4400

.153 .14 07 0.16 10.15 0.41 1400 a 0 0 *400 1*3 .66? 0.18 I015 0.32 tcl ____3 0 0 D 143 *16.12 0.10 0.14 0 30 WYC 0.-6 0 0 0 4*00 1*3 *0.43 0.1 016 024 63 0.A7 2? 0 0 0 4400 FV.span 44 73 1ds .6.8. 0.16 015 014 _ IOV~i.w MY 0.29 #oIG.AFCC 0 a 0 4400 17 14S .13,47 0.18 015 0.21 NVe 0.37 0 0 0 1400 133 .17 S4 0.16- 016 02 #M 0 a 0 4400 133 +12.3s 0.18 0.18 0.29 SYC 0.37 0 0 0 4400 133 .8.00 0.16 0.13 0.21 901 035 so 0 0 0 4400 133 .18.12 0.16 020 0.33 90 0 42 27 0 0 0 4400

.123 *12451 0.16 0.15 0.24 NM0 0 a 0 4400 Ms3 *10.591 018 015 01d I10 wool___ 0 0 0 4400 UP.wTS 19 VIL 0.00 _ 0.28 8 el8 I~Dcam ___ __ IOAICASC 0 0 a am-- 17

- - CTL -101 017 0.13 0.IS t ol0Vurviwoc NM0 _ 1_O0KW9 0 0 0 MO50 Fu..tosa II1291UT3S40.0 273 033 1.24 3800OD 1o4010.No C55 2 0 0n014380es TVP 133 1l s-0 0-4 01`3 N

___ NM______ 0 4400 I - 2.3 0.37 091 -M: *8 M&I l 01l 43 010 0 4400 U,.ewTSL 43 2 01.T -0 00 1 017 lTSj G t m o G~S 0 0 0 n 17 Tube 66-131 in OTSG-B ruptured at 4360 psig. The 3.2 gpm leak rate occurred %Nitha measured pressure at the pump of approx. 450 psig (and a calculated differential pressure at thes defect of less than 100 psig).

88 ECR #02-0112 1, Rev. 2 1

Table of Acronyms AECL Atomic Energy of Canada, Ltd.

AKEL Axial Kinetic Expansion Length ASME American Society of Mechanical Engineers ATP Abnormal Transient Procedure BOL Beginning of [Core] Life BWOG B&W Owners Group CFR Code of Federal Regulations COA Crack Opening Area DHR Decay Heat Removal EAB Exclusion Area Boundary ECT Eddy Current Test EDM Electro-Discharge Machine EFW Emergency Feed Water EPRI Electric Power Research Institute ESAS Engineered Safeguards Actuation System ETL Expansion Transition Location F Fahrenheit FSAR Final Safety Analysis Report FTI Framatome Technologies, Inc.

FWLB FeedWater Line Break GPU General Public Utilities GPUN GPU Nuclear Corp.

HF High Frequency HFP Hot Full Power HPI High Pressure Injection ID Inside Diameter IGA InterGranular Attack KET Kinetic Expansion Transition LBLOCA Large Break Loss of Coolant Accident LCL Lower Confidence Limit LPZ Low Population Zone LRF Leakage Reduction Factor MDP Motor Driven Pump MFW Main Feed Water MRPC Motorized Rotating Pancake Probe MSIV Main Steam Isolation Valve MSLB Main Steam Line Break MSSV Main Steam Safety Valve MTC Moderator Temperature Coefficient NDD No Detectable Degradation NDE Non-Destructive Examination NODP Normal Operating Delta Pressure NOPD Normal Operating Pressure Differential NPSH Net Positive Suction Head NQI Non-Quantifiable Indication NRC Nuclear Regulatory Commission NSSS Nuclear Steam Supply System OD Outside Diameter 89 ECR#02-01121,Rev. 2

OTSG Once-Through Steam Generator PICEP Pipe Crack Evaluation Program P-T Pressure-Temperature PWSCC Primary Water Stress Corrosion Cracking R Radius RAI Request for Additional Information RCP Reactor Coolant Pump RCS Reactor Coolant System RPS Reactor Protection System RPV Reactor Pressure Vessel SBLOCA Small Break Loss of Coolant Accident SCC Stress Corrosion Cracking SCM Sub-Cooling Margin SG Steam Generator TBV Turbine Bypass Valve TDP Turbine Driven Pump T-H Thermal-Hydraulic TMI Three Mile Island TMI-I Three Mile Island, Unit I TS Tubesheet UFSAR Updated Final Safety Analysis Report UTSF Upper Tubesheet Secondary Face 90 ECR #02-01121, Rev. 2

ATTACHMENT 3 Regulatory Commitments

5928-05-20102 Page 1 of 1 List of Re-gulatorv Commitments The following table identifies commitments made in this document by AmerGen. Any other actions discussed in the submittal representing intended or planned actions by AmerGen are described to the NRC for the NRC's information and are not regulatory commitments.

COMMITMENT COMMITTED DATE The proposed TMI UFSAR update wording The proposed TMI UFSAR update will be (Attachment 2) will be implemented and implemented and posted against the TMI posted against the TMI UFSAR, and then UFSAR document within 60 days after incorporated into the subsequent UFSAR issuance of the NRC SER on alternate Update. repair criteria, and incorporated into the subsequent UFSAR Update.

Tubes with circumferential indications in the kinetic expansions' required lengths will be removed from service upon detection - Beginning with the T1 RI6 Refueling including circumferential indications that Outage (Fall 2005) and all subsequent were detected during prior 1997 through refueling outage examinations on the 2003 outage examinations that remain in original steam generators.

service in the kinetic expansions' required lengths.

Implement 100% examination scope of in- Beginning with the Ti Ri6 Refueling service kinetic expansions each refueling Outage (Fall 2005) and all subsequent outage. refueling outage examinations on the original steam generators.

Any tubes with flaws detected in the sleeves, or in the parent tube adjacent to Beginning with the Ti Ri6 Refueling the sleeve between the lower sleeve end Outage (Fall 2005) and all subsequent and the parent tube kinetic expansion refueling outage examinations on the transition, will be plugged-on-detection. original steam generators.

ATTACHMENT 4 NRC Question Response

5928-05-20102 Page 1 of 17 REQUEST FOR ADDITIONAL INFORMATION Related to TMI-1 Kinetic Expansion Inspection and Acceptance Repair Criteria August 16, 2004 Letter The Nuclear Regulatory Commission (NRC) has reviewed the AmerGen October 4, 2002, and August 16, 2004 (the August 16 submittal was a complete revision of the October 4 submittal),

submittals associated with the TMI-1 KE inspection, acceptance and repair criteria. The quality and clarity of the submittal has significantly improved such that the NRC staff has been able to review the proposal and identify a complete list of issues. Based on recent industry experience, NRC review of information related to similar alternate repair criteria and issuance of Generic Letter 2004-01, "Requirements for Steam Generator Tube Inspections," dated August 30, 2004, the NRC staff has identified concerns with the TMI-1 KE inspection, acceptance and repair criteria. The NRC staff is meeting with AmerGen to discuss these issues. The below agenda outlines the areas of concern that will be discussed during this working level public meeting.

1. Suitability of leaving circumferential flaws in-service

2. Leakage assessment including thermal hydraulic analysis model

3. Basis for assumption of no growth/no initiation of new flaws

4. Structural integrity assessment

a. Basis for steam line break axial load

b. Determination of the limiting accident

5. Consistency with recent industry experience

6. Inspection practices/techniques

7. Other issues Specific issues under the above topics to be discussed include the following:

1. Suitability of Leaving Circumferential Flaws In-service

a. The large-break loss-of-coolant accident (LBLOCA) issue applies to TMI-1, (based on request for additional information (RAI) responses provided in the October 4, 2002 submittal) and circumferential flaws are of particular concern in this event due to the increase in axial loads. This issue is not currently addressed in the August 16, 2004, TMI-1 KE report, ECR #02-01121, Revision 1, Inspection Acceptance Criteria and Leakage Assessment Methodology for TMI OTSG [once-through steam generator]

Kinetic Expansion Examinations," which is inconsistent with industry practice for joint repairs, therefore this issue remains unresolved.

Response

AmerGen has drafted a revised ECR 02-01121 to require that tubes with circumferential Indications in the kinetic expansions' required lengths will be removed from service upon detection-including circumferential indications that were detected during prior 1997 through 2003 outage examinations that remain in service in the kinetic expansions' required lengths.

5928-05-20102 Page 2 of 17 AmerGen has also revised ECR 02-01121 to provide additional discussion of the LBLOCA and its axial loads with respect to the kinetic expansion inspection criteria. Refer to Section 2.0 of the revised ECR.

Note that AmerGen is still working with the other B&W plant owners to revise BAW-2374, the LBLOCA topical, to address the LBLOCA transient for all aspects of the steam generators' design and maintenance.

b. For a 0.64-inch long circumferential crack, what is the factor of safety against tube severance under a main steamline break (MSLB), based on elastic analysis? For this same crack, does the axial thermal stress behave as a primary or secondary stress?

(For design in accordance with Section III of the American Society for Mechanical Engineers (ASME), Boiler and Pressure Vessel Code (Code), axial thermal stress is always considered secondary. For circumferentially cracked tubes, if the crack is large enough such that deformation occurs largely at the crack rather than being relatively evenly distributed along the length of the tube, then the net section stress at this location is not "self limiting" and should be treated as primary. Industry representatives (e.g.,

Nuclear Energy Institute (NEI)) have stated they are developing guidance for when thermal loads should be considered primary versus secondary. If the licensee has an alternative justification for the 0.64-inch circumferential crack criterion other than elastic analysis, the licensee is requested to provide that justification including justification for the value of the safety factor assumed to the applied load.

Response

Elastic analyses were used to determine the 0.64-inch dimension. Thermally-induced stresses were secondary.

Recent scoping calculations indicate that strain concentration at circumferential cracks will not prevent full load relaxation of the secondary MSLB thermal loads.

The calculations Indicate that the strain concentration for a 0.64-inch circumferential flaw Is approximately 3%.

The safety factor of the structural calculations (i.e., tube severance calculations) is at least 1.8 (1310 lbs. axial tensile load based on thermal-hydraulic analyses of the MSLB transient versus 2400 lbs. assumed for the kinetic expansion structural analyses.)

As described above, AmerGen has drafted a revised ECR 02-01121 to require that tubes with circumferential Indications In the kinetic expansions' required lengths will be removed from service upon detection-including circumferential indications that were detected during prior 1997 through 2003 outage examinations that remain In service In the kinetic expansions' required lengths.

c. Based on recent industry operating experience, circumferential flaws are likely to initiate in the KE expansion region at TMI-1. The NRC staff's concerns related to the "no growth" and "new initiation" statistical assessments have not been resolved and are discussed in more detail in Item 3 below.

5928-05-20102 Page 3 of 17

Response

AmerGen has addressed these items in Item 3, below.

d. No other domestic plant leaves circumferential flaws in service in the steam generator tube pressure boundary. Based on this and other issues associated with circumferential cracks (identified above), the NRC staff would like the licensee to discuss the suitability of its proposal to leave circumferential flaws in service.

Response

As described above, AmerGen has drafted a revised ECR 02-01121 to require that tubes with circumferential indications in the kinetic expansions' required lengths will be removed from service upon detection-including circumferential Indications that were detected during prior 1997 through 2003 outage examinations that remain in service in the kinetic expansions' required lengths.

2. Leakage Assessment Including Thermal Hydraulic Analysis Model

a. Did the leak tests, performed in support of this inspection/repair criteria, use deoxygenated water? Recent industry experience indicates tests not performed with deoxygenated water may be non-conservative.

Response

AmerGen has drafted a revised ECR 02-01121 to eliminate the use of the leakage reduction factor, and the leak tests that were performed to determine the leakage reduction factor are no longer discussed in the draft ECR.

b. The proposed leakage model assumes zero leakage from flaws located above the region of the tubing which is inspected and evaluated in accordance with the structural integrity criteria (i.e., the SG tube pressure boundary). This is inconsistent with industry practice for similar repair criteria. Recent industry experience indicates the leakage from this region may not be minimal, when assumed for all in-service tubes. Therefore, the assumption needs to be modified to reflect this.

Response

AmerGen has drafted a revised ECR 02-01121 to incorporate the conservative assumption that flaws above the region of the tubing that is inspected may leak.

Refer to Section 5.6.5 of the revised ECR.

c. Is the leakage assessment conservative for volumetric intergranular attack (IGA), given that the axial and circumferential extents (i.e., components) are independently assessed? Is there experimental and/or analytical evidence which indicates that summing the leak rates from projected axial and circumferential crack components of volumetric IGA indication give conservative leakrate estimates?

5928-05-20102 Page 4 of 17

Response

Considerable evidence exists that the TMI-1 steam generator volumetric indications will not leak, confirming conservative leakage assessment of these indicators. A large number of TMI-1 volumetric indications have been in situ pressure tested, without leakage. AmerGen has drafted a revised ECR 02-01121 in which in situ pressure test results are provided (in an attached Table 5) and these results are discussed in Section 4.1.4.

d. Laboratory tests were performed by the licensee to develop a leakage reduction factor (LRF). Information from other sources indicates lower reductions in leakage due to pressure. 1) The licensee calculated the LRF by putting a clamp over an electro-discharge machined (EDM) notch and measuring the reduction in leakage. To account for internal pressure, the licensee used the zero-applied contact pressure results as the basis. Clarify how these adjustments (relative to the zero-applied contact pressure) were made and the basis for concluding they are conservative. 2) Discuss whether use of a notch is conservative when the results are applied to cracks/volumetric IGA given that cracks have much lower leak rates.

Response

AmerGen has drafted a revised ECR 02-01121 to eliminate the use of the leakage reduction factor.

e. Based on the information submitted, the NRC staff understands that the following: 1)

Flaws are assumed to be 100% through-wall for the entire measured (via eddy current) extent for the structural analysis. 2) Flaws less than 67% through-wall are assumed not to leak for the leakage assessment. 3) Flaws exceeding 67% through-wall are assumed to be 100% through-wall for the entire measured (via eddy current) extent for the leakage assessment. Please confirm the above understanding. [Please note: Table 111-6 of the 2001 Steam Generator Tube Inservice Inspection Report for TMI-1 (1R14 SG Outage Report) implies that the length of assumed 100% through-wall crack length is less than the measured (via eddy current) crack length. Please clarify this discrepancy.]

Response

The NRC's understanding as stated above is correct. Table 111-6 of the 2001 outage report was applicable to freespan, vice kinetic expansion, indications.

Thermal Hydraulic Model for the Leakage Assessment

a. The NRC staff has determined that the thermal hydraulic model used for the leakage assessment is different than that used for the structural assessment and has not been reviewed by the staff. Considering the resulting axial loads are significantly lower (i.e., 1310 pounds (Ibs)) than those used for the structural assessment (i.e., 2400 Ibs),

the NRC staff has concluded that the thermal hydraulic model used for the leakage assessment must be reviewed.

5928-05-20102 Page 5 of 17

Response

As discussed at the February 16 and 17th, 2005 NRC meeting, AmerGen has drafted a revised ECR 02-01121 to provide additional information regarding the thermal hydraulic modeling for NRC review. (Refer to Section 5.2, for example.)

The differences between the various calculations and models were discussed in the above referenced meeting. Additional information is also provided in the responses to item 2.d, 2.e, 2.g, and 2.h below.

b. The leakage assessment is performed for a MSLB using revised tube loading conditions (based on use of a different thermal hydraulic model as discussed above). The loads on the tube for the revised MSLB analysis are lower than for other accidents (e.g., small and large break LOCA). As a result, it is not clear whether the MSLB is still the most-limiting accident in terms of assessing the consequences of leakage from these joints given the differences in loading conditions between the accidents and the different assumptions for assessing the radiological consequences of these accidents.

Response

AmerGen has drafted a revised ECR 02-01121 to provide additional information regarding the Main Steam Line Break and other transients. (Refer to Section 2.0 of the revised ECR.)

c. Inthe thermal hydraulic analysis for the leakage assessment, the licensee appears to have tried to maximize the cooldown rate to increase the axial tube loading. However, it is not clear whether this results in an overall conservative result given it may have decreased the differential pressure across the tubes (and the driving force for the leakage). In addition, it appears that the leakage assessment was performed based on the actual loads on the tube at various time intervals and the leakage over these intervals were summed. Regarding this approach, it is not clear how the licensee accounted for all of the uncertainties in all of the models (e.g., thermal hydraulics, tube material properties, PICEP, etc.) to ensure that the leakage estimates have high confidence (e.g., a 95% prediction interval at 95% confidence). Provide the details of how the leakage calculations are performed.

Response

As discussed during the February 16 and 17th, 2005 meeting of AmerGen and NRC, leakage was calculated In 1997 over various time intervals and then summed. The most conservative values of various factors that influence leakage, including primary-to-secondary differential pressures, tensile loads on the tubing, tubesheet bore dilations, temperatures, etcetera, do not occur simultaneously during the event. For this reason a conservative, time-dependent analysis of the MSLB was created, from which flaw leakage could be calculated.

Tube material properties are discussed in the response to Question 4, (b), below.

Conservative thermal-hydraulic analysis of the plant's response to a hypothetical MSLB event and conservative evaluation of projected leakage were utilized; no

5928-05-20102 Page 6 of 17 attempt was made to determine a single measurement of the accuracy of the leakage modeling.

Note also that the proposed revision to ECR 02-01121, as described above, eliminates the use of any leakage reduction factor. This significantly increases the conservatism of the leakage calculations. With the removal of the leakage reduction factor, the leakrates utilized for kinetic expansion indications are very similar to those used in the acceptance criteria for freespan volumetric ID IGA indications at TMI-1 (i.e., kinetic expansion tube-to-tubesheet joint contact pressure is neglected).

d. Please confirm that the methodology and input assumptions used for your MSLB analysis for generating inputs to define the OTSG tube load are consistent with that used in the MSLB analysis documented in Updated Final Safety Analysis Report (UFSAR),

Section 14.1.2.9. Identify any deviations from the licensing basis methodology, analysis assumptions and initial conditions and provide proper justification for such deviations.

Response

As discussed in our meeting of February 16 and 17, 2005, the 1997 MSLB analysis for OTSG tube loads, and to estimate OTSG tube leakage from the kinetic expansions, was not the same analysis as that discussed in UFSAR Section 14.1.2.9. The analyses discussed In Section 14.1.2.9 were performed to conservatively calculate the affect of a MSLB on reactor response and containment response. The calculations are very similar in many respects. For example, core power, break size, RETRAN Codes, initial RCS pressures and temperatures, etc., were similar for both analyses. There were a number of deviations necessary to perform the OTSG tube load analyses, since conditions that are conservative for evaluation of OTSG tube load may not be conservative for evaluation of reactor and containment response.

The OTSG tube load analysis Is described in Section 5.0 of ECR 02-01121. The following Is a summary of the deviations from the UFSAR Section 14.1.2.9 methodology's analysis assumptions, initial conditions, and their justification.

Item OTSG Tube Load UFSAR Justification MSLB Analysis MSLB Analysis -: _;_-:_-_:_-_-_-

Reactor Kinetics BOL reactor kinetics EOL reactor kinetics Cooler primary gives cooler steam generator tubes and maximizes tube tensile loads. For reactor peak clad and containment response, warmer primary Is conservative.

Delayed Neutron 0.007 0.005 Smaller value Is Fraction conservative for reactor/contalnment response. Larger value Is conservative for OTSG tube response.

5928-05-20102 Page 7 of 17 Item OTSG Tube Load' UFSAR Justification MSLB Analysis

-V- MSLB Analysis - _-

Decay Heat 0.95 times ANS5.1 ANS5.1 Similar to above 2 standard Items. Less decay heat cools the primary and increases tube stresses Reactor Trip Trip modeled for low Trip modeled for high BOL kinetics will cause primary system flux or low primary reactor trip on low pressure system pressure primary pressure Emergency Affected SG was Affected SG was Greater EFW flow FeedWater Flow assumed to receive assumed to receive increases the cooling, 590 gpm, the maximum 570 gpm. and induced axial flow allowed by stresses, of steam cavitating venturi. generator tubes.

Main FeedWater Failure of feedwater No feedwater Greater MFW flow Flow regulating valve was regulating valve failure Increases the cooling, assumed (as described was assumed. and Induced axial In Section 5.3 of ECR stresses, of steam

_ 02-01121.) _ _ _ generator tubes.

e. Provide the justification for why a reactor trip setpoint of 1900 psig plus a 30 psi error will result in a conservative calculation with respect to SG tube temperature for OTSG loads.

Response

The reactor trip setpoint of 1900 psig plus a 30 psi error is consistent with the trip setpoint used in the plant's other UFSAR MSLB analyses (for determination of reactor peak clad temperature and containment building environment). Early trip of the reactor is conservative for the evaluation of MSLB-induced steam generator tube axial loads since the tubing Is cooled more quickly. (As described in Section 5.3.2 of ECR 02-01121, the analyses minimized steam generator tube temperatures in order to maximize tube axial loads.)

f. In the long-term analysis, operator actions are credited in the analysis. Please confirm that all operator actions credited in this analysis are consistent with the plant emergency operating procedures at TMI-1 and that the reactor operators are properly trained on the plant simulators for these operations. Justify that the time allowed for operator action is adequate, and has been verified on the plant simulator.

5928-05-20102 Page 8 of 17

Response

Operator actions were credited for the successful mitigation of the MSLB event.

(Other operator actions were conservatively modeled to ensure the analysis produced a conservative result.)

The operator actions used in the 1997 analysis were consistent with the plant's emergency procedures in 1997, and remain consistent with the plant's current emergency procedures.

The operator actions that were credited in the analysis to mitigate the event include terminating EFW flow and controlling tube-to-shell delta temperatures during cooldown. Terminating EFW flow within 10 minutes if MSLB occurs is recognized as a time-critical element in the plant's Emergency Procedure program and simulator training. There is no specific time requirement for controlling tube-to-shell delta temperatures.

Operators are regularly trained on the emergency procedures.

In the 1997 analysis, no operator action was credited for the first 10 minutes of the event, consistent with the plant's other UFSAR analyses of MSLB.

g. Please compare the transient curves between the new analyses for OTSG tube load and the licensing analysis in Section 14.1.2.9 of the UFSAR. Identify each deviation and provide proper justifications (the NRC staff has noted quite a few differences).

Response

AmerGen has revised the proposed UFSAR pages to clarify that the transient analyzed in UFSAR Section 14.1.2.9 (from which the transient curves in the UFSAR were generated) is a different transient than that used for the OTSG tube load analyses.

Refer to Question d., above, for some of the differences between the UFSAR and ECR 02-01121 analyses. The differences in the transient curves are due to differences in the Input assumptions. Note also that the time scales are different between the UFSAR plots and the plots provided in ECR 02-01121.

h. In Section 14.1.2.9 of the UFSAR, it is concluded that the results of the analysis confirm that the maximum temperature differential that occurs in the OTSG does not produce excessive stress, and SG integrity is maintained. Discuss why this 100-second analysis supports such a conclusion and why your new analysis requires both a 10-minute duration and a long-term analysis to assess the SG tube integrity.

Response

The original UFSAR 100-second analysis, which was performed to evaluate reactor fuel peak clad temperatures and containment environments resulting from a hypothetical MSLB, supported the statement in the UFSAR. The UFSAR statement was written at the time of those analyses.

5928-05-20102 Page 9 of 17 The 1997 long-term analyses were performed specifically to evaluate OTSG tube stresses and hypothetical tube leakage over the duration of the MSLB event.

OTSG tube leakage was assumed to calculate the environmental consequences and this was reported in Section 14.1.2.9 (C).

3. Assumption of No Growth/No Initiation of New Flaws

a. Statistical tests performed to determine whether flaws are growing compare data from the current outage to data from the prior outage. The NRC staff believes the licensee should use data from the outage during which the first rotating probe examination was performed of each KE. This would ensure that potential slow flaw growth rates would be more evident. It is requested that the licensee discuss its plans to perform the statistical tests in this manner.

Response

AmerGen has drafted a revised ECR 02-01121 to specify that data be utilized from the first examinations of kinetic expansions. Refer to Section 3.2.1.3 of the revised ECR.

b. An extreme value test is performed to identify possible outliers or erroneous data.

Erroneous data is corrected prior to using that data in the subsequent statistical tests.

Outliers (i.e., indications with large apparent growth rates) are used in the subsequent statistical tests. Industry experience indicates that when a population of flaws grow, some grow faster than others. Therefore, the NRC staff would like to discuss why the outliers are not, in and of themselves, considered evidence of flaw growth, and therefore, an invalidation of the no-growth assumption used to calculate the flaw acceptance criteria.

Response

AmerGen agrees with the NRC that outliers, so long as they are not based on erroneous data, may be indicative of growth in and of themselves. (This is also consistent with the philosophy of the EPRI Guidelines.) Outliers indicative of flaw growth were not seen in the last outage examinations.

AmerGen has drafted a revision to the ECR 02-01121 criteria to require that all

'new' indications not previously detected during the 1997 through 2001 examinations will be plugged. Refer to Section 3.2.1.9 of the draft ECR.

The 'no-growth assumption' used In the ECR was based on the conservatism of the analytical methods. (In 1997, when the ECR criteria were originally created, there was essentially no prior outage MRPC/PlusPoint data with which to assess population growth.)

The ECR requires that the NRC be notified if growth of the kinetic expansion indications is detected.

5928-05-20102 Page 10 of 17

c. The NRC staff is not confident the threshold value of 0.05 new indications per KE examined is truly indicative of an active degradation mechanism. In addition, different criteria for circumferential and volumetric degradation may be appropriate since they are potentially two different populations. Industry guidance on this subject would indicate that one new crack results in a declaration of active degradation. Discuss why the size of the indication is not a consideration or why comparisons to prior data are not sufficient (i.e., if it cannot be seen with hindsight, it is new). Lastly, based on the information provided in Table B in Section 3.2.1.9 of the August 16, 2004 submittal, it could appear that degradation is active with an initiation rate of 0.03 indications per KE.

Please provide a discussion of the above issues.

Response

AmerGen has revised ECR 02-01121 to delete the subject threshold value. (Refer to revised Section 3.2.1.9 of the ECR.)

AmerGen has also drafted a revised ECR that requires that new indications, not detected during prior examinations, be 'plugged on detection'. (Also in revised Section 3.2.1.9 of the ECR.) This is a very conservative treatment of new indications.

The draft ECR requires that circumferential indications in the kinetic expansions' required lengths will be removed from service, as described in the responses above.

In addition, AmerGen has drafted the revised ECR to implement a 100%

examination scope of in-service kinetic expansions each refueling outage. (Refer to responses to questions above.) An examination with scope of 100% would identify active degradation should it occur.

d. The licensee indicated that some KE indications "drop out", or disappear, each outage.

These should be discussed in more detail including examples of several indications (e.g., largest, smallest, theory on reason for disappearance, etc.). If threshold-of-detection is ascribed to be the cause of disappearance, be prepared to discuss the criteria for the threshold-of-detection.

Response

These indications were discussed during the February 16 and 17 th, 2005 meeting of AmerGen and NRC. The large voltage "drop in" and "drop out" indications discussed were of an administrative nature such as differences in axial location or flaw classification (e.g., called as two separate Single Circumferential Indications (SCI) one outage and called as a single Multiple Circumferential Indication (MCI) during a different outage). The typical new indications were very small in voltage, axial extent, and circumferential extent (i.e., indications with an expected low probability of detection). These small indications are typical of small volumetric ID IGA that has been destructively evaluated from tube pulls at TMI-1.

As described in the response to Question 3(c), just above, AmerGen has drafted a revised ECR 02-01121 to eliminate the evaluation of indications that may have "dropped out" during each inspection.

5928-05-20102 Page 11 of 17

4. Structural Integrity Assessment Basis for MSLB Axial Load

a. The licensee's August 16, 2004, report states that the 3140-lb. axial load corresponds to an axial membrane stress of 49.5 ksi (thousand pounds per square inch) and a design-basis tube strain of 0.16%. The licensee further states that tubes with a lower bound yield strength and nominal geometry will experience load relaxation (from 3140 Ibs.) due to yielding, resulting in an axial load of 2400 Ibs. This load is used to determine the size of the needed "defect free zones" in the KE.

The "design basis" tube strain of 0.16% was determined assuming that all tubes were behaving elastically (Topical Report BAW-10146). The corresponding axial membrane stress of 49.5 ksi exceeds the nominal yield strength of the tubing at an MSLB temperature of 235 degrees F. Had the actual stress/strain properties of the tubes been assumed in the licensee's analysis, rather than elastic properties, the resulting tube strain could exceed 0.16% since the tube bundle and tube sheet would provide less resistance to the tendency of the SG shell (with temperature in the range of 520 to 575 degrees F) to expand axially relative to the tubes. If credit is taken for load relaxation in the tubes due to yielding (as is the case for TMI-1), why is it not necessary to also consider the corresponding increase in tube end displacements (and, thus, tube strain) when determining the axial loads in the tubes?

What are the tube end displacements and tube strains under MSLB temperatures if a realistic distribution of stress/strain properties are assumed to exist within the tube population? What would be the effect on the axial loads and minimum required defect free lengths assuming use of the realistic distribution of stress/strain properties?

Response

Calculations were performed, in 2005, to address tube end displacements for a hypothetical tube bundle containing only low yield strength tubing. The increase in displacement was approximately 10%. Strain hardening for low yield strength material Is negligible for this increase in tube end displacement.

The resulting load increase is also negligible.

What factors of safety are applied to the axial loads to ensure the joints don 't slip when determining the necessary size of the defect-free zones? What is the technical basis for the safety factors? [Note, a factor of safety of 1.0 is reasonable for thermal loads behaving as secondary, as acknowledged in the structural performance criteria in the latest Technical Specification Task Force (TSTF) submittal from NEI of the generic license change package and in the forthcoming revision to NEI 97-06, "Steam Generator Program Guidelines." But this safety factor criterion is based on the assumption of elastic analysis, recognizing that load relaxation will take place prior to failure. A safety factor of 1.0 is not appropriate if one is taking explicit credit for load relaxation, since components at the point of incipient failure under design-basis loadings would be contrary to ASME Code, Section 1II,and Section Xl philosophy.]

5928-05-20102 Page 12 of 17

Response

The factor of safety against slip is at least 1.8, based on the ratio of the assumed loads to those from the thermal/hydraulic analysis. (Refer also to the response to Question 1(b), above.)

b. The licensee states that the design-basis MSLB load for the SG tubes of 3140 lbs. was determined by assuming that all tubes remain fully elastic. It was necessary to adjust the results obtained for the high-yield strength tubes and greater wall thickness for consideration of minimum yield strength and nominal wall thickness tubes that may be present in the SGs. Additional details are required for the NRC staff to fully understand this adjustment.

• What are the dimensional tolerances for the nominal 0.625" diameter, 0.034" thick tubing?

Response

The allowable tube tolerances, based on the original tubing Purchase Order, were:

Outside diameter: 0.625" plus 0.005", minus 0.000" Wall thickness: 0.034" plus 0.005", minus 0.000" Minimum wall thickness was assumed for the ECR analyses.

Tube pulls from the TMI-1 steam generators have validated these dimensions.

For example, 1997 tube pull samples had typical O.D.'s of 0.625" to 0.629",

and typical l.D.'s were 0.554" to 0.555".

• What are the estimated nominal, upper bound, and lower bound yield strengths of the tubing at room temperature and at the temperature associated with the maximum MSLB load (i.e., 235 degrees F as reported in BAW-10146)?

Response

Section 2.2.3.1 of GPUN Topical Report 007, "Three Mile Island Unit 1 Once-Through Steam Generator Repair: Kinetic Expansion Technical Report",

(March, 1983) provides TMI-1 tubing yield strength information that was used to develop the KE acceptance criteria:

"...Tubes that are positively traceable as being in the TMI-1 steam generators have 0.2% offset yield strength values from 41.0 to 61.1 ksi. Tubes that may be In the generators, but which are not individually traceable as such, have 0.2% offset yield strength values from 41.0 to 64.9 ksi."

The minimum yield strength (i.e., 41 ksi) tubing was assumed for the derivations/calculations of the kinetic expansion acceptance criteria.

The reduction in yield strength at MSLB temperatures should not exceed approximately 5%. This Is compensated for by not including thermal tightening in the structural analysis model, and other conservatisms.

5928-05-20102 Page 13 of 17 It is not evident to the NRC staff, based upon its review of BAW-10146, that the 3140- lb. MSLB load is based on a larger than nominal tube wall thickness as is suggested in the licensee's words above. For a nominal 0.625" outside diameter (OD) tube with a nominal 0.034" thick wall, the cross-sectional area of the tube is 0.0631 square inches. BAW-10146, Table 5-6, indicates that the 3140-lb. load is based on this same nominal cross-sectional area. Provide an explanation for this apparent discrepancy. Provide a description of tube wall dimensions assumed in the calculation of the 3140-lb. MSLB load.

Response

AmerGen has proposed a revision of the subject sentence in Section 2.2 of the ECR in order to clarify it. The sentence originally read:

"...It was necessary to adjust the results obtained for the high-yield strength tubes and greater wall thickness for consideration of minimum yield strength and nominal wall thickness tubes that may be present in the SGs."

The revised sentence reads:

"...In order to create a conservative finite element analysis model it was necessary to adjust the model to reflect that many of the 1980's pull testing results were obtained using tubes of high yield strength and greater wall thickness (for consideration of the minimum yield strength and nominal wall thickness tubes that may be present in the steam generators)."

Minimum tube wall dimensions (0.034" wall thickness) were used for the derivations/calculations of the kinetic expansion acceptance criteria. There was no adjustment of wall thicknesses in either BAW-1 01 46 or the ECR; the only adjustments were those made to the ECR's finite element model to conservatively match up with the pull testing results.

The pullout resistance of the tube from the tube sheet is a function of the contact pressure caused by the expansion process, the effects of thermal tightening (differential thermal expansion between the tube and the tube sheet), tube internal pressure, and tube sheet bow. During a steamline break transient, the tube internal pressure and the tube temperature are changing. In addition, the yield strength of the tube changes with temperature. The yield strength of the tube affects the system response (e.g., the applied load due to load relaxation). It is not clear whether the analysis provided truly represents the most-limiting conditions of the transient. The licensee should confirm that the most-limiting point of the accident was evaluated for the most limiting situation (high-yield strength tubing/low-yield strength tubing) using the most-limiting input parameters (lowest contact pressure/pullout resistance of any of the test data). This approach is consistent with how we have assessed other similar amendments. The goal is that all tubes have adequate integrity so worst-case assumptions are generally made.

5928-05-20102 Page 14 of 17

Response

The MSLB transient analysis was performed over time, as described in Section 5.2 of the ECR. Leakage from theoretical flaws was calculated at various intervals over the duration of the transient. The most limiting tube axial loads, highest temperatures, highest primary-to-secondary pressure differentials, etc.

do not occur simultaneously during the event. As described above, tubing of minimum yield strength and minimum wall thickness was assumed for the analyses. Minimum yield strength material results in the lowest pullout resistance; higher yield strength material results in proportionately higher joint pullout strengths.

Determination of the Limiting Accident In determining the limiting accident, it is not clear what factors of safety were applied under all events considered (e.g., LOCA, normal operating, feedwater line break conditions, etc.). Question 4a., above, focuses on the factor of safety used in assessing the MSLB accident; however, it is not clear whether another event may be more limiting if appropriate safety factors were used (this question assumes the correct safety factors were not applied).

Response

The 'design' transient for the TMI-1 kinetic expansions in the 1980's was the Main Steam Line Break (MSLB). Other transients were considered, including LOCAs, normal operating stresses, feedwater line break conditions, etc. The MSLB transient has a combination of relatively high axial tube loads and primary-to-secondary differential pressure. In addition, the MSLB event has potentially limiting nuclear consequences as a result of theoretical leakage through kinetic expansion joints. The 1997 derivation of the kinetic expansion inspection criteria, as described In ECR 02-01121, was consistent with the original design analyses of the kinetic expansion joint; MSLB was the limiting transient.

AmerGen has drafted a revised ECR 02-01121 to provide additional discussion of the transients, and the associated factor of safety for the MSLB, in Section 2.0.

5. Consistency with Recent Industry Experience Analyses of kinetically expanded joints in other designed SGs have indicated that defect-free lengths greater than what is being proposed for TMI-1 are needed to ensure structural and leakage integrity. In addition, the contact pressures for the TMI-1 KEs appear to be significantly larger than those at other plants with KEs. Therefore, please compare and contrast the expansion process used at TMI-1 to the KE processes used at other plants to help the NRC staff understand the potential differences (e.g., joint tightness, resistance to cracking, etc.).

5928-05-20102 Page 15 of 17

Response

The TMI-1 kinetic expansions were expanded twice to maximize their contact pressure. The design objective was 99% confidence that 99% of the kinetic expansions had pullout strengths of greater than 3140 lbs. tensile load.

The TMI-1 kinetic expansions have required lengths for structural concerns (i.e.,

joint slippage, tube parting, or tube sever) that are similar to those for explosively-expanded joints at other plants.

The TMI-1 required lengths to ensure leakage integrity are shorter than those at some of the other plants since the TMI-1 17" deep kinetic expansions are located near the mid-planes of the TMI-1 tubesheets. Analyses of MSLB-induced tubesheet flexure at plants with full-depth expansions have shown that the maximum tubesheet bore dilations occur at the secondary faces of the tubesheets, reducing further Into the joint toward the mid-plane of the tubesheets.

(Note that the TMI-1 22" long kinetic expansions, which are closer to the secondary faces of the tubesheets, have longer required defect-free lengths for this same reason: they are closer to the secondary faces of the tubesheets.) In summary, the joint lengths required for leakage integrity are similar.

Note that some of the other U.S. PWRs with explosive expansions have steam generator tubing with a minimum yield strength of 35 ksi. As described in responses above, the TMI-1 tubing minimum yield strength was 41 ksi. This is another feature that influences the calculations of the required joint length.

6. Inspection Practices/Techniques

a. On page 23 of the August 16, 2004, report, the licensee indicates that if "localized" degradation occurs at a KE, then the scope of the inspection will not be expanded to 100%. The specific example given was damage from a maintenance tool. If growth or new degradation is occurring, the scope should be expanded to 100%. It is not clear what other "localized" degradation could be occurring. Discuss the intent of this statement.

Response

AmerGen has drafted a revised ECR #02-01121 that requires 100% of the in-service kinetic expansions to be examined during each refueling outage inspection. Therefore, scope expansion Is no longer applicable.

b. On page 23 of the August 16, 2004, report, the licensee states that if growth of existing degradation or initiation of new degradation in the KE region is detected, then an examination of 100% of the KEs in the affected generator(s) will be undertaken.

However, the proposed statistical tests combine data from both SGs because there is a limited data population in the "B" SG. Therefore, the NRC staff assumes the 100%

scope expansion would occur in both SGs. Please confirm this assumption.

5928-05-20102 Page 16 of 17

Response

AmerGen has drafted a revised ECR #02-01121 that requires 100% of the in-service kinetic expansions to be examined during each refueling outage inspection. Therefore, scope expansion is no longer applicable as a result of the statistical analyses.

c. The licensee states that the eddy current measurements always result in conservative overestimates of the flaw size. This is attributed to lead in and lead out affects and the flaw being small in comparison to the coil field. How did the licensee confirm that measurement uncertainty is not a function of flaw size (i.e., is there a flaw size beyond which the flaw size could be underestimated (at a 95% confidence level))? In addition, did the licensee confirm that the 95% confidence levels on uncertainty for cracks and for volumetric IGA (i.e., non-notch specimens) when analyzed separately from the notch data still result in overestimates of the flaw size?

Response

The TMI-1 kinetic expansion Indication sizes are small in comparison to the flaws used in the sizing studies. Size distributions were provided to the NRC staff at the AmerGen/NRC meeting of February 16-17, 2005, and are also provided in Section 4.1.4 of ECR #02-01121. No studies were performed to determine flaw sizes beyond which extents might be underestimated.

Note that TMI-1 has drafted a revision of ECR #02-01121 In which crack-like indications are removed from service in the required lengths of the kinetic expansions. (Refer to responses in Section 1 above.) The result will be that kinetic expansion indications remaining in service after the next refueling outage will be ID-initiated volumetric ID IGA. Tube pulls from TMI-1 have shown that eddy current conservatively sizes the TMI-1 ID volumetric IGA indications, as is discussed in Section 4.1 of ECR #02-01121.

d. Discuss the inspections performed of the parent tube/sleeve assembly in the upper tube sheet region. Describe the flaw acceptance criteria utilized for the portion of the sleeve/tube assembly located in the tube sheet.

Response

AmerGen revised Section 2.7 of draft ECR #02-01121 to clarify that any tubes with flaws detected in the sleeves, or In the parent tube adjacent to the sleeve between the lower sleeve end and the parent tube kinetic expansion transition, will be "plugged-on-detection".

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7. Other Issues

a. The proposed reporting requirements should be supplemented to include the KE length and the tube sheet radius associated with each tube with degradation in the KE region.

Response

AmerGen has revised Section 5.9 of draft ECR 02-01121, "Reporting Requirements", to require that this information be submitted to the NRC staff.

b. Discuss the axial loads simulated during insitu pressure tests for the purposes of demonstrating structural and leakage integrity.

Response

AmerGen has drafted a revised Section 4.1.4 and has provided a new Table 5 in the ECR in order to provide summary Information regarding steam generator tube in situ pressure tests that have been performed at the plant. In situ pressure tests of indications within kinetic expansions have not been performed, with or without axial loads, because conservative test results would not be obtained due to the presence of the tubesheet. (Refer also to Section 4.1.2 of the revised ECR.)

c. The number of tubes in Section 1.3.2.39 of the revised Updated Final Safety Analysis Report pages does not add up (i.e., the number of tubes in-service versus the number of tubes plugged do not correlate). Discuss this inconsistency.

Response

AmerGen has proposed a revision to the subject section in order to clarify it.

Refer to the proposed UFSAR changes enclosed with this document.

The author(s) of the subject sentences did not intend to necessarily correlate between the number of steam generator tubes in service, the number of tubes plugged, and the total number of tubes, since the subject sentences were written about different sets of tubing at different time periods. The inconsistency existed because some tubes were counted in more than one group of tubes. In order to prevent any potential future confusion, the subject sentences were eliminated.