ML20086D429

Forwards Response to NRC 950531 RAI Re Comm Ed Proposed TS Amend to Increase Interim Plugging Criteria for Byron & Braidwood Unit 1 Sgs.Proprietary Westinghouse Rept WNEP-9124, Program Tranflo... Also Encl.Rept Withheld

ML20086D429
Person / Time
Site:	Byron, Braidwood
Issue date:	06/30/1995
From:	Saccomando D COMMONWEALTH EDISON CO.
To:	NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM), Office of Nuclear Reactor Regulation
Shared Package
ML19330G145	List: ML20086D429
References
NUDOCS 9507100132
Download: ML20086D429 (52)

Text

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Dow ners Grm e. II. 60515 June 30, 1995 Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, D.C. 20555 Attn: Document Control Desk

Subject:

Response to Request for Additional Information Regarding Commonwealth Edison Company's Proposed Technical Specification Amendment to Increase the Interim Plugging Criteria (IPC) for Byron and Braidwood Unit 1 Steam Generators NRC Docket Numbers: 50-454 and 50-456

Reference:

D. Lynch letter to D. Farrar transmitting Request for Additional Information dated May 31, 1995 i

The Reference letter transmitted the Nuclear Regulatory Commission's (NRC) Request for Additional Information (RAI) pertaining to Commonwealth Edison Company (Comed) Technical i

Specification amendment request to increase the current Interim I Plugging Criteria (IPC) for Braidwood and Byron Unit 1 Steam i Generators. Attached is Comed's response to the RAI questions 1-

26. ]

1 I

Please note that included in the Attachment is the hard copy response to question 2. This TRANFLO output file for Case 1 was provided via internet E-Mail to Mr. Donoghue on June 7,.1995.

Additionally, WNEP-9124, " Program TRANFLO Version 1.0, A Computer l Code Update with Improved Input / Output," contains information I proprietary to the Westinghouse Electric Corporation, it is

} respectfully requested that the information which is proprietary

( to Westinghouse be withheld from public disclosure in accordance l with Title 10, Code of Federal Regulations, Part 2, Section 790 (10 CFR 2.790). Due to the expedited nature of this submittal, WNEP-9124 is not currently supported by an affidavit signed by Westinghouse, the owner of the information.

1 Westinghouse will provide an affidavit which sets forth the basis

! upon which the information may be withheld from public disclosure I by the Commission and addresses with specificity the considerations listed in 10 CFR 2.790(b)(4) within four weeks.

Comed and Westinghouse regret any inconvenience that the absence of this affidavit causes the Staff.

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NRC Document Desk June 30, 1995 Correspondence with respect to the proprietary aspects of.the items listed above should be addressed to N. J. Liparulo, Manager of Nuclear Safety & Regulatory Activities, Westinghouse Electric Corporation, P. O. Box 355, Pittsburgh, Pennsylvania 15230-0355.

If you have any questions concerning this correspondence,please l contact this office.

Sincerely, .

+N k m _ fa Denise M. Saccomando Nuclear Licensing Administrator Attachment cc: D. Lynch, Senior Project Manager-NRR G. Dick, Byron Project Manager-NRR R. Assa, Braidwood Project Manager-NRR J Martin, Regional Administrator-RIII Office of Nuclear Safety-IDNS I

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Ecsponse to NRC Request for Additional Information Regarding the Proposed Revisions to the Technical Specifications Related to the Interim Plugginst Criteria Byron Unit 1 and Braidwood Unit _1 J

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S.\APC\CCE95\RAI.626.WP5 June 29,1995 L-

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l The following are the responses to the request for additional information regarding the proposed revisions to the technical specifications related to the interim plugging criteria for the Byron Unit 1 and Braidwood Unit 1 as transmitted on May 26, 1995. Question numbering is keyed to that used by the NRC in the transmittal of l

the RAI to Comed.

1. Indicate which version of the TRANFLO code was used to perform the calculations cited in WCAP 14273 and provide complete documentation for this version. The documentation should describe any updates to the code since the publication of WCAP-8821 which described the version of the TRANFLO code l approved by the NRC staff. i

RESPONSE

As discussed in WCAP-14273, TRANFLO code was developed by MPR Associates. WCAP-8821 documents the original version of the TRANFLO in detail, which was approved by the NRC staff. The original version considers the steam and water mixture as a homogeneous fluid. Later, MPR Associates implemented a drift flux model to the original version of the TRANFLO. The purpose is to better simulate relative flow velocity between water and steam.

This drift-flux version is documented in MPR-663 in 1989. The document describes the mathematical model, physical model and list of the FORTRAN program. A complete document of the Drift-Flux Version (MPR-663, November 1980) is included as Attachment 1 to this response. An appendix of an MPR report, which shows the comparison between results of calculations by the j original and drift flux versions, is also included in Attachment 1.

The TRANFLO Version 1.0 (November 1991) is an update to the Drift-Flux Version (MPR 663). A complete document of the TRANFLO Version 1.0 is included in Attachment 1 to this response.

The version 1.0 has only one inlet for feedwater entering the steam generator.

The preheat steam generators, such as the Model D4 for the Braidwood and Byron Plants, require two inlets for feedwater flow. The TRANFLO Version 2.0 (January 1993) allows two feedwater inlets. This involves a change to the Subroutine EXTFLOW in the TRANFLO Version 1.0. A comparison was made between results of calculations using Versions 1.0 and 2.0. Figures included in Attachment 1 show their comparison and, as expected, they agree well.

In summary, the documents included as Attachment 1 provide the detailed information and updates for the TRANFLO version used in WCAP-14273.

S.NAPC\CCE95\RAI_526.WP5 June 29.1995

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2. Furnish the TRANFLO output files for the Case 1 calculation (i.e., a postulated MSLB at hot standby) with sufficient information to permit the staff to make meaningful comparisons with its independent calculation of spatial and temporal variaticns of pressure, temperature, steam quality, and void fraction.

RESPONSE

The TRANFLO output file for the Case 1 was provided to Mr. Donoghue on June 7,1995 by Internet E-Mail. A hardcopy of the same file is attached for reference. The file includes transients of pressure, temperature, steam quality and void fraction in each node. They will provide sufficient information for evaluating both spatial and temporal variations of these parameters.

3. For the Case 1 calculation, what steam separator performance is assumed in terms of pressure drops, exit steam qualities, and flow rates.

RESPONSE

To properly estimate the pressure drops through the primary and secondary separators, pressure loss coefficients are used for them. Based on laboratory and field data evaluation, the primary separator loss coefficient for the Braidwood and Byron Model D4 steam generator is taken to be 13.9.

Similarily, the secondary separator loss coefficient is taken to be 40.

The function of the secondary separator was simulated in its capacity to trap l and separate the water from steam. Primary separator system consists of riser, downcomer and orifice. The riser is represented by Nodes 28 and 27, and a flow connector (i.e., Connector No. 37). The riser downcomer is represented by Node No. 46, which has an upstream flow connector (No. 31) to Node 27, and a downstream flow connector (No. 32) to Node 47 (a fluid reservoir). The orifice is represented by a flow connector (No. 30). Therefore, riser flow from tube bundle can split into riser-downcomer flow and orifice flow paths.

Moisture separation by the swirlvane was not simulated in the Case 1 calculation. A good moisture separation during full power operation will deliver about 70% of the water in the riser to downcomer Node 46 and then to the fluid reservoir Node 47. The remaining 30% would go to Node 26.

The initial water level for the Case 1 calculation is such that riser below the swirlvano is essentially submerged in the water. The major mechanism of hydraulic process followed a steam line break is the water flashing due to rapid depressurization. Water in either the fluid reservoir Node 47 and the one above it (i.e., Node 26) would flash rapidly. In view of the above, whether to S.\APC\CCE95\RAI.52fLWP5 June 29,1995

4 simulate the swirlvane or not would not be critical for the thermal and hydraulic process in the tube bundle.

A calculation for Case 1 with the simulation of the swirlvane was recently performed. The results for the pressure drop through tube support plates are essentially the same as those obtained without the swirlvane. A small difference appears at the peak of the pressure drop; the run with the swirlvane yields a slightly higher peak, about a few percent more than that from the run without the swirlvane. The effects of this difference on the TSP displacements would be negligible and the difference is negligible compared to the uncertainty allowances (factor of two) applied to the TRANFLO loads for the displacement analyses.

4. The description in WCAP-14273 of the TRANFLO capabilities listed fluid temperature as one of the parameters calculated by the code. However, the results presented in WCAP 14273 are focused on TSP hydraulic loading effects. l How were the spatial and temporal variations of the fluid temperatures used to evaluate the thermal effects on steam generator (SG) internals considered in the analysis done to support your position that there will be minimal tube support plate (TSP) deflections under accident loadings?

l

RESPONSE

Table 41 is a spread sheet which lists the temperatures of the tube wall and secondary fluid as calculated by the TRANFLO for the Case 1, hot standby condition. It covers all nodes for the total four seconds. Note that the hydraulic load peaks within seconds, and then decreases to small, steady value within the four second interval. The table provides information for the evaluation of the spatial and temporal variation.

i The temporal variation is small; it is within 13*F over 4 seconds. Figures 4-1 and 4-2 illustrate the temporal variation for the tube wall and secondary fluid at the tubesheet and U-Bend, respectively. Figures 4 3 and 4 4 demonstrate the spatial variation of tube wall and secondary fluid along the hot and cold leg, respectively. The spatial variation of the tube wall and secondary fluid temperatures are limited to less than 5*F from the tubesheet to U-Bend. Thus, both the temporal and spatial difference between the tube wall and secondary l

fluid temperature are small. The tube support plate would follow the secondary fluid temperature.

Based on the above discussion, it can be concluded that effect of temperature variation is insignificant on the relative deflection of the tube support plate to the tube. The small temperature variations over the approximately four second i S \APC\CCE95\RAI 526.WP5 June 29,1995

interval that the TSP displacements are significant would not influence the displacement analysis results.

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5. When a free span SG tube burst test is conducted, the test results differ from the pattern of fracture progression that would occur if the potential burst were to be constrained within a TSP. In a low energy test where the burst of the tube will immediately lower the pressure (e.g., in a test using a hydraulic jack),

the fracture is a line-like opening tending toward a fish mouth opening. In a  ;

high energy test where the pressure continues to be applied after fracture of the wall, the resulting fracture begins to tear at the chevron marks at the ends  !

of the axial crack resulting in extensive circumferential tearing at the crack ends. The comparative results of high and low-energy SG tube pressure test fracture surfaces are shown pictorially in the book " Theory of. Flow and Eracture of_ Solids - Volume I,"' in the chapter entitled " Tests under Combined Stress."

The test results presented in Tables 9.1 and 9.2 of WCAP-14273 for SG tubes with t,hrough wall axial cracks protruding slightly outside the TSP, pressurized to the postulated main steam line break (MSLB) pressure differential of 2560 psi, show reductions in burst pressure with increasing crack exposure. In the l event that there may be larger TSP deflections than that reported in WCAP 14273, the burst pressure can significantly decrease and approach the free span burst pressure, which for 2560 psid, could result in a more

" catastrophic" fracture progression similar to that in a high energy test.

Accordingly, describe the equipment used in the SG pressurization. Discuss l the pressurization source for the burst tests, the utilization of the " bladder",

and the volume of the pressurizing medium. Additionally, indicate whether the burst pressure remains constant until the fullest possible extent of the burst fracture is achieved.

RES10NSE Additional information relative to the phenomenon referred to is provided in the same reference in Chapter 15, " Limiting States of Stress in Solids, Theories of Mechanical Strength." The information contained in the reference text is from Davis, E. A., "The Effect of Size and Stored Energy on the Fracture of Tubular Specimens," Journal of Applied Mechanics, Vol.15, American Society of Mechanical Engineers, September,1948. The comparative results of high

' Nadai, A., Theory of Flow and Fracture of Solids, Second Edition, McGraw.

Hill Book Company, New York,1950.

SAAPC\CCE95\RAL526.WP5 June 29.1995 C____.______.___

and low energy tube pressure test fracture surfaces which are shown pictorially in the reference are not those of SG tubes, i.e., the material tested was "semikilled medium ship steel" with a yield strength of 38 ksi and an ultimate strength of 60 ksi in the plate form. The fracture behavior in the high-energy {

tests was described as being initially shear and then changing to brittle l (cleavage). Additionally, illustrations of unstable propagation of cracks in Alloy l 600 SG tubes are illustrated in Hernalsteen, P., "The Influence of Testing Conditions on Burst Pressure Assessment for Inconel Tubing," International Journal of Pressure Vessels and Piping, Vol. 52,1992. However, the mode of unstable failure in the latter reference is no less dramatic than that in the former. In both cases, rupture proceeds such that the opening in the tube wall l is significantly larger than the cross sectional area of the tube. For Alloy 600 however, the additional tearing is also ductile (shear) and there is no evidence of a ductile to brittle (cleavage) transition.

The key point of each of the references, and the request for additional information, is that the unstable crack extension occurs after the initial burst pressure is reached. Considerations of free span burst have not involved assumptions regarding the nature of the tube opening following the initiation of rupture. A constrained burst is considerably different. In the latter case the flanks of the crack are not free to deform significantly (contact with the TSP hole would be expected at about one half of the burst pressure), and the crack behaves like a shorter crack, with an attendant much lower probability of burst, and hence lower probability of unstable extension.

The burst testing performed in support of the alternate plugging criteria has not been in a high energy facility. The tubes are lined with a plastic tube, also referred to as a bladder, prior to the test. The OD of the bladder is reinforced ]

with a small,2 mils thick, foil (brass) shim to prevent extrusion of the bladder prior to achieving the burst pressure of the tube. Once the crack tips start to extend and the flanks open significantly, the pressure in the pressurizing medium is released and no further energy is supplied. Hence, dramatic crack openings as illustrated in the references are not achieved. The extended crack tips for the " fishmouth" type of opening do retain a shear failure appearance.

When cracks are pressurized within or slightly offset from the TSP, such as in support of the IRB leak test program, essentially a constant pressure to maximumize crack opening is applied since there is no "burt" or significant bladder extrusion to release the pressure.

The emphasis of the analyses performed in support of the APC is to demonstrate that the probability of burst is acceptably low. This is the case regardless of considering the burst to occur in a high or low-energy loading configuration.

Jure 29,1995 S.\APC\cCE95%RAL526.WP5

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6. In the data presented in Figure 9-2 of WCAP 14273, have these clearance constrained tests been conducted in a high-energy facility? If so, describe the failed surface at the ends of the protruding crack at burst. If the tests were not conducted in a high energy test facility, describe the failed surface.

RESPONSE

The test results reported on Figure 9 2 of WCAP-14273 were not from tests performed in a high-energy test facility. The failed surface at the ends of the crack form an angle of approximately 45' with the exterior and interior surfaces of the tube. This is similar to the failed surfaces resulting from test cases where the tearing has continued as a result of being in a high energy test I situation (see the discussion of the previous request for additionalinformation).

It should be emphasized that the tests are performed to obtain the burst  ;

pressure corresponding to initial crack tearing at the tips of the crack and no l assumptions are made on the area of the opening following a burst. Bladder l pressurizations made to support the IRB leak tests for increasing the crack opening are equivalent to constant pressure tests since there is no " burst" to relieve the bladder pressure.

7. At the intersection of the SG tube U bend with the straight section of the SG l

'ube, there are transition lengths having highly strained sections of tube l I

asulting from the manufacturing process (e.g., ovality and non uniform cold working of the tube material) which could be additive to the flexural stresses resulting from differential expansion between the hot and cold leg of a SG tube, flow induced stresses, and seismic stresses. Accordingly, state whether there are manufacturing process effects which could contribute to the potential for SG tube burst at the uppermost support. If so, descr'be this interaction.

Justify your position if you believe there is not an adverse impact of the manufacturing effects on the burst strength of the SG tubes.

RESPONSE

The tangent point of the U bend with the straight leg of tubing occurs at an elevation of about 2.1" above the top of the uppermost TSP. The effects oflocal residual stresses from the forming operation, i.e., resulting in ovality and non-t uniform cold working of the material, would affect the tube over a length of about 0.3" from the tangent point. Surface stresses from the original tube straightening operation would be relieved by any blunting of the crack tip.

Overall, the load in the tube is tensile in the axial and hoop directions. The S.\APC\CcE95\RAI 626.WP5 June 29,1995

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I tensile stress in the axial direction tends to reduce the size of the plastic zone at the crack tip, thus increasing its resistance to fracture from the hoop stress.

Due to separation of the tangent point from the uppermost TSP, the manufacturing process for the U bend have no significant influence on the burst capability of the tube at the TSP. In general, burst tests have shown higher burst pressures for the bent U bend tube sections than for straight sections due to the cold working and curvature of the tube.

8. Are there operating stresses in the SG tubes other than normal tangential and l axial stresses associated with the pressure differential between the primary I and secondary side that could contribute to the bursting of the tube? For example; does the hydraulic loading on the TSPs during the blowdown phase of a postulated MSLB induce significant axial stresses in the SG tubes which are used to limit the deflections of the TSPs? If so, have you considered these.

blowdown loads in evaluating the analysis of these expanded tubes?

RESPONSE

l Relative to the effect on burst strength of axially oriented cracks, axial tensile stresses in a tube will tend to close such cracks, while axial compressive stresses will tend to open the crack. Thus, compressive stresses have the

potential to reduce the burst capability of the tube due to the crack opening.

l Table 8-13 of WCAP 14273 summarizes the maximum axial forces in the expanded tubes (the tubes that are used to limit the deflections of the TSPs).

The maximum compressive force is -458 pounds. The cross sectional area of a nominal tube is 0.0955 inch 8. Thus, the 458 pounds converts to an axial stress of 4.8 ksi.

Section of 8.3 of WCAP 13494, Revision 1, Catawba Unit 1 Technical Support for Steam Generator Interim Pluaaine Crite"la for Indications at Tube Suovart Plates. provides a discussion of test results to investigate the effects of bending stress on the burst capability of tubes with axially oriented cracks. The test results show that a compressive stress on the order of the yield strength of the tube is required before any significant effect is realized in the burst pressure capability of a cracked tube. In the vicinity of the crack, the bending stress acts like a membrane stress, so it is judged that these results are also applicable to the case being considered here.

In addition, the expanded tubes experiencing this stress are to be removed from service such that burst is not an issue. If any additional, active tubes, are +

fixed in the plates due to the presence of corrosion products, then the axial load will be distributed among these tubes, further reducing the stress. Thus, it is SAAPC\cCE95\RAl.526.WP5 June 29.1995

concluded that any axial stresses induced in the tubes as a result of their being locked in the plates will be small, and will not have any significant effect on tube burst strength.

9. In a free span test, as the flawed tube tends to burst, a crack opens and develops into a fishmouth shape, followed by further tearing at the crack ends.

In the constrained tests presently being conducted with a portion of the crack outside the simulated TSP, has the crack extended beyond the original pre-cracking dimensions? What is the fracture appearance of the SG tube after bursting within the constraint? Why have the tests been restricted to small crack offsets (i.e.,0.10 inches for the 0.75-inch diameter SG tubes?

RESPONSE

If the constrained tests are conducted to a high enough pressure to result in bursting of the tube, the free end of the crack extends in shear, i.e.,

intersecting the surface of the tube an an angle of approximately 45*. If the tube has not burst, extension of the crack occurs by blunting of the crack tip, which is on the order of about 40 mils in Alloy 600 SG tubes. Since the volume of material must remain constant, some apparent extension takes place.

The pressurizttion tests in support of the IRB leak test program have been limited in crack offset to the maximum TSP displacement conservatively predicted for a SLB event with tube expansion to " lock" the TSPs.

Conservatisms in the displacement analyses are documented in WCAP 14273.

These maximum displacements of 0.10" for Model D4 SGs (3/4" tubing) and 0.15" for Model 51 SGs (7/8" tubing) apply to the most limiting tube location on any TSP. Nearly all SG tubes have much smaller maximum TSP displacements. Thus, the IRB leak tests are performed at the bounding TSP offsets and there is no need to test at larger TSP offsets in support of the tube expansion based APC.

There has been no discernable crack length extension found, for data available to date, as a result of flow or bladder pressurization to the free span burst pressure. Crack length measurements prior to leak testing were obtained by dy penetrant tests. Length measurements after flow testing are made with a toolmaker's microscope. Differences between pre test and post-test length measurements have been within about 0.020", which is too small a difference to separate measurement differences from 2 rack tearing to extend the crack and the differences may be due to tip blunting. Final pre leak test corrosion  !

I lengths will be determined by destructive exam following test completion.

These data may permit a better determination of crack extension although it is clear that crack tearing has not significantly increased the total crack length.

I S.\APC\CCE95\RAL626.WP6 June 29,1995 1

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10. Figure 9 3 of WCAP44273 shows the effect of the gap between the SG tube and the TSP on the probability of tube burst with a 0.7 inch crack extending outside the TSP between 0.3 and 0.7 inches. State whether these probability assessments are based on data from burst experiments. If not, provide justification for your probability assessments.

RESPONSE

Section 9 4 describes the methodology used to develop the probability of burst of a single, free-span indication as a function of crack length based on the results of the regression analysis of burst pressure as a function of crack length based on the results of over 200 burst tests. The probability of burst of a single indication as a function of crack exposure is found as the probability of randomly obtaining a Student's t distribution variate as large as that calculated using equation (9.10) developed in Section 9.5 of the report from the burst testing of partially exposed specimens. It is stipulated that the discussion contained in Section 9.5 should have made reference to the development of the curve labeled "PoB for Large Clearance" on Figure 9-3. The origin of the value 0.425 used in equation (9.10) is discussed in Section 9.3 of the report and is considered to be conservative for the smaller crack exposures predicted by the plate displacement analyses. This adjustment to the free span burst pressure is based on burst tests performed for different tube to TSP gaps as a function of the crack length extending outside the TSP as shown in Figure 9-2. It is also considered adequate for larger exposures up to 0.5", and likely conservative for exposures greater than 0.5" since the constraining effect of the TSP must asymptotically approach zero when the crack is entirely exposed.

11. In the test program currently underway with SG tubes having cracks extending outside the simulated TSP, the burst pressure has been assumed to be equal to that of the free span burst pressure with a crack whose length is equal to the length protruding from the tube support. What is your justification for this assumption? Is the constrained part of the crack assumed to have the same effect as a crack tip at the edge of the support plate?

RE: PONSE The calculation of burst pressures based on the length of the crack equal to the length protruding from the tube support plate is only applied to calculate the SG tube burst probability following implementation of tube expansion. This i methodology is not applied for the IRB leak test program currently underway. I l

S \APC\CCE95\RAL626.WP5

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For the IRB leak test program, leak rates are initially measured with the primary flow pressurizing the tube and tests are performed with the tip of the l crack at the edge of the TSP and the tip of the crack offset (0.10" for 3/4" tubing,0.15" for 7/8" tubing) from the edge of the TSP. To further bound the potential leak rates by opening of the crack within the limits permitted by the TSP constraint, leak rate measurements are also performed after pressurizing the imlication with a bladder to the calculated free span burst pressure of the .

indication. It has been demonstrated in the test program that this bladder pressurization step opens the crack face to contact the ID of the TSP hole and tends to maximize the crack opening. This demonstration was achieved by,in one test, progressively increasing the bladder pressure by about 1000 psi ,

increments and repeating the leak rate measurements after each step. For two {

tests, a bladder pressurization step was also performed at less than the  ;

calculated free span burst pressure. The bladder pressurization steps are typically performed with the crack tip offset from the edge of the TSP. I Between the combinations of pressurization by primary flow and by bladders, I the bounding leak rate has been well defined for each specimen in the test program and no assumptions are made about the true burst pressure for an indication extending outside the TSP.

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12. In the SG tube burst tests presently being conducted, the pre-cracked tubes are expanded prior to insertion in the simulated TSP. Explain why the tubes with cracks are pre-expanded outside the simulated TSP instead of bursting or measuring leakage from the tube entirely within the support constraint, i

RESPONSE

As described in the Question 11 response, the IRB leak rate tests are performed with primary flow pressurization prior to bladder pressurization.

Only Test 4-1 was performed with all leak tests following bladder pressurization. When bladder pressurization is applied, the indication for most specimens was located with the crack tip offset from the TSP. The offset condition for bladder pressurization, rather than the indication within the TSP, was selected as it might be expected that the offset condition would potentially maximize the crack opening at the tip of the crack. For the one bladder l pressurization test performed with the crack tip within the TSP and repeated j offset from the TSP, there was no significant difference in the crack opening i area. In general, the results tend to indicate that the crack opening area is l limited by contact of the center of the crack with the ID of the TSP hole and the opening area is not likely to be dependent on offsets from the TSP as long as the offset distances are less than about half the crack length, l

8 \APC\ccE95\RAL526.WP5 Jo. .r. 25

. l Leak rate measurements are performed for each specimen with the crack tip at the edge of the TSP and offset from the TSP. These measurements are performed even though the bladder pressurization step was performed with the offset crack location. Thus, the test results permit comparisons ofleak rates for the crack within and offset from the edge of the TSP.

Note: Due to the_ similarity ofQuestions 13 and 14, a_shule response ig.orovided.

13. The calculated TSP displacements are based on an elastic analysis. However, it is uncertain that the TSPs remain elastic under all postulated accident condition loadings. If the TSPs become inelastic, even locally near support rod intersections, the clastic analysis could be invalidated. Provide additional justification and/or alternate bounding calculations to demonstrate that the elastic analyses of TSP displacements remain valid.

14. Describe in detail the development of the solid plate model, stated in WCAP-14273 to be equivalent to the actual TSPs, which simulates the behavior of the TSPs with SG tube penetrations, flow holes and various cutouts.

Provide additional information demonstrating that the stresses in the equivalent solid plate provide a conservative bound for those areas of the actual TSP which have the most limiting stress levels.

RESPONSE TO QUESTIONS (13) AND (14)

The overall finite element model used in the analysis is shown in Figure 7-14 of WCAP 14273, with geometry plots for each of the individual plates shown in Figures 7 21 through 7-28. Relative to the various cutouts in the plates (other than tube and flow holes), it is apparent from these plots that the crescent shaped cutout on the hot leg side of the FDB has been accounted for, as well as the flow slots for the TSP along the tube lane and, in some cases, at the plate periphery.

To account for the tube and flow holes in the plates, equivalent plate properties were calculated for Young's Modulus and Poisson's ratio. The properties for the FDB, which does not have flow holes, were calculated using the paper by Slot and O'Donnell, Effective Elastic Constants for Thick (and Thin) Perforated Plates with Square and Triangular Penetration Patterns, Journal of f Engineering for Industry, November 1971. For the TSP, which has a double square penetration pattern of tube and flow holes, the equivalent properties were calculated using a Westinghouse Proprietary Research Paper, The Analysis of Perforated Plates with Square Penetration Patterns of Circular

/ Holes and Square Holes, May 1981. Although different formulations are used SAAPC\cCE95\RAI.526.WPS June 29,1995

for the FDB and TSPs, the calculations use the same methodology in each case. Due to square penetration patterns, different properties exist in the pitch and diagonal directions. The first step is to establish equivalent parameters for Young's modulus and Poisson's ratio in the pitch and diagonal directions s (E,*/E, E4 */E, va *, v,*), respectively. The equivalent Young's modulus for the j overall plate is taken as the average of the pitch and diagonal directions. The next step in the process is to determine an equivalent value for the shear modulus, G'/G, for the plate. This is done in a similar manner as for Young's modulus, starting with values in the pitch and diagonal directions, and then taking an average of the two values. The final equivalent value for Poisson's ratio is determined from the relationship between Young's modulus and the shear modulus. The effective clastic constants were assumed to apply from the centerline of the tube lane out to the last row of tubes. Referring again to Figure 7-21, the untubed region of the plates corresponds to the last two rows of elements. Thus, effective elastic constants were specified for all but the last two rows of elements for each plate.

Relative to the plate stresses, the stresses reported in WCAP-14273 in Figures 8 27 through 8 34 are calculated as described in Section 8.10 of WCAP-14273. Briefly, the displacements from the dynamic solution at the times of maximum plate displacement are applied to the plate models as boundary conditions, and the corresponding equivalent plate stresses are calculated. For the elements where effective plate properties have been used,  ;

the equivalent plate stresses were multiplied by the inverse ratio of Young's i modulus (E / E* = 10.55) to get stresses for the perforated plates, and correspond to the stresses shown in the above figures. These stresses do not necessarily correspond to the peak ligament stresses, however. In order to get a relationship between the equivalent plate stresses and peak ligament stresses, some additional supplementary calculations have been performed.

The supplementary calculations have been performed using two finite element models of the actual hole and ligament geometry for the Braidwood 1 and Byron 1 steam generators. One modelis used to evaluate the plate response to applied loads in the pitch direction, and the second to evaluate loads applied in the diagonal direction. The orientation of the two models is shown in Figure 131 (attached) relative to the overall hole geometry. The model geometries for the two plate segments are shown in Figures 13 2 and 13 3 (attached) for the pitch and diagonal directions, respectively. These models were used to calculate the ratio of peak stress intensity for the perforated plate to an equivalent solid plate. Since the plates are loaded out of plane by the steam j line break pressure drops, the response is in terms of plate bending. Thus, the j ligament models were analyzed tubject to the application of edge moment loads. Peak stress factors were calculated as a function of bi axiality of the applied moments for ratios ranging from -1.0 to 1.0. The resulting peak stress j SAAPC\CCE95\RAI.526.Wr5 June 29.1995 i

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intensities were then compared to the peak stresses for an equivalent solid plate subject to the same edge loads, and ratios of the stresses were calculated.

The results of the calculations are summarized in Table 131 (attached), and show a maximum stress intensity ratio of 7.43. (Note that the results in Table 13-1 have been compared to results of similar analyses performed for other plate geometries and show good agreement both in distribution and magnitude.)

Comparison of the peak stress ratio of 7.43 calculated with the ligament models to the inverse ratio of Young's Modulus used in the analysis of 10.55, shows that the stresses reported in WCAP 14273 are conservative by a factor of 1.42. If the results from the above calculations are used, the stresses for Plates C(3H), and J(7H), will be elastic essentially throughout, as will the stresses for Plate P(11H), except for the center tab. Further review of the results for the center tab element for Plate P(11H) shows that the factor of 10.55 was also applied to this element. Since there are no flow holes in this region, the equivalent plate stresses of approximately 5 ksi apply directly.

There is one remaining location where the plate stresses exceed the ASME Code minimum yield stress of 23.4 ksi. Referring to Figure 8-30 for Plate C(3H) along the left vertical edge of the plate, stress contour number 20 will exceed yield by less than 2 ksi. Recall that this is a surface stress, and yielding will only occur over a small portion of the plate thickness. Also note that this is not in the location of maximum displacement. Furthermore, the maximum calculated displacement is -0.0581 inch (see Table 810 of WCAP 14273), versus a limit of 0.100 inch. Thus, yielding of the plate surface over a very localized region will not significantly affect the maximum plate displacement.

Thus, it is concluded that the plate stresses reported in WCAP 14273 are conservative (by a factor of 1.42). Based on the results of the supplemental calculations, it is further concluded that the stresses exceed the ASME Code minimum yield for one plate over a very small area, and that a significant margin, in terms of plate displacement, exists for this plate. Yielding of the plate surface over a very localized region will not significantly affect the maximum plate displacement. Therefore, based on these results, it is concluded that the use of elastic analysis to predict plate displacements under the applied loads is acceptable.

15. Provide a description of the special purpose computer program "pltdym".

Indicate how extensively this program has been used, and whether it has been validated and verified. Also, indicate whether the code has been benchmarked in accordance with NRC Standard Review Plan guidelines.

8,\APC\CCE95\RAL526.WP5 June 29,1995 i

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RESPONSE

Computer program "pitdym" is a special purpose computer code that was developed specifically for evaluating the tube bundle response to a time history pressure loading. The fundamental building block of the code, the algorithm i for solving the differential equation for displacements, has been used (

extensively in other codes related primarily to tube dynamic response.

In addition to the basic equation solving capability of the code, the following capabilities have been built into the code; l

1. Incorporation of non linear support interaction. Depending on the steam generator model, wedges may be welded to interfacing members l such that they will provide resistance to vertical plate motion. (The primary function of the wedges is to provide in-plane alignment of the support plates.) The non-linear support interaction has been programmed in such a manner that it can act in either the up or down directions (positive or negative), or can be incorporated as a linear support that is always active.

2. Incorporation of non linear plate / spacer interaction. The support system for the tube support plates is composed of several types. One type of support is a tierod / spacer combination. The tierods are solid bars that are threaded into the tubesheet at the bottom of the tube bundle, run the full height of the bundle passing through each of the plates, with a nut on the top surface of the top plate. On the outside of the tierods are cylindrical spacers that are situated between the support plates that serve to align the plates vertically. The plate /

spacer interface is non-linear in nature. Thus, the capability for non-linear plate / spacer interaction has been incorporated in the code.

3. If sufficient plate displacement and rotation occurs under the applied loads, it is possible that interaction between the plates and tubes may occur. If the plate rotates locally such that the top surface of the plate contacts the tube on one side while the bottom surface of the plate contacts the tube on the other side, then the tube will bind up in the plate and restrict further deflection of the plate. The non linear interaction of the tube and plates has been inccrporated in the dynamics code.

4. In addition to the capability to solve the time history response of the structure, the code can also solve for an initial set of statically applied forces.

SMPC\CCE95\RAl.626 WP5 June 29,1995 ;

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"Pltdym" has been verified and validated consistent with Westinghouse procedures for verifying any special purpose code that is used in the course of performing a structural analysis. Separate verification problems have been evaluated for each of the above features of the code, as well as the basic equation solving capability of the code. Documentation of the code listing, sample input and output, and results from each of the verification probleme is documented and has been checked by an independent analyst.

16. Compare the amount of SG tube expansion you are proposing in WCAP-14273 with that experienced by the alloy 600 parent tube material in a hybrid expanded joint (HEJ) sleeved tube and in a laser welded sleeved tube. Provide comparative residual stress plots derived from a finite element analysis showing the residual stress of the expansion for an HEJ, a laser welded sleeve, and your proposed TSP expansion. Include the residual stresses at the various transition / weld steps.

RESPONSE

The proposed hydraulic expansion at the TSPs results in a diametral increase of about 57 to < 90 mils with the increase within this range dependent on the yield strength of the particular tube. A HEJ expansion consists of a hydraulic expansion of the tube diameter about 0-10 mils, followed by a hardroll expansion of about an additonal 17 mils. A laser weld sleeve hydraulic expansion is about 0-3 mils on the diameter. Other relevent expansions include tubesheet expansions and hydraulic expansions performed in the cold leg of preheat SGs to limit tube vibration. Tubesheet expansions include hydraulic, mechanical and explosive expansions and are nominally about 16 mil diametral increases. In the early hydraulic tubesheet expansions, an occasional expansion was bulged above the top of the tubesheet. Tubesheet bulges up to about 30 mils were left in service. The preheat expansions typically ranged from about 14 to a maximum of about 40 mils. However, preheat expansions with an additional bulge (resulting from occasional misalignment of the expansion mandrel relative to the TSP) of up to about 30 mils outside the TSP were left in service.

Additional information on each of these types of expansions and the residual stresses obtained from stress indexing tests is given in the response to Question 20. Finite element analyses have not been performed for the TSP expansion of WCAP 14273 due to difficulties in converging the analysis model and modeling the stress / strain curve for the larger expansions. Relative corrosion susceptibility has been evaluated based on the stress indexing tests discussed in the Question 20 response.

SAAl'C\CCE95\RAL526.WP5 June 29,1995 i.- .

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17. Discuss the operating experience for any similar type of expanded joint (similar with respect to percent expansion and residual stress distribution) in Alloy 600 SG tubes in similar service. Identify the heat treatment of these similar applications using Alloy 600, including the stress corrosion cracking (SCC) susceptibility ranking.

RESPONSE

Operating experience and residual stresses for similar types of expanded joints in mill annealed Alloy 600 SG tubes are addressed in the response to Question

20. Only Westinghouse mill annealed Alloy 600 tubes are used for the service life and SCC susceptibility estimates in the Question 20 response.

Historically, only recent LWS expansions of the similar expansions discussed in the Questions 16 and 20 responses were heat treated following application of the expansion process. Relative to the ID environmental conditions, none of these expansions is similar to those planned for application at Byron 1 and Braidwood 1 since the tubes to be expanded are to be removed from service.

The nominal environment will be entrapped air unless the expanded tube includes a throughwall indication, and the temperature will be on the order of 544*F, whereas the sleeve and tubesheet expansions include exposure to primary water on the inside of the tube and operate at inside temperatures on the order of 580 to 620*F. .

l Relative to comparing the expansions with those associated with HEJ sleeves, a discussion of the sleeve / tube upper joint is appropriate. The joint consists of a hydraulic expansion (HE) of a length of about 4" followed by a hardroll of about 1" slightly above the center of the HE. The oldest HEJ sleeved tubes have been in service for about 12 years without incidence of outoide diameter SCC. Recent advances in ECT technology, and operating experience at Doel 4, I have resulted in the detection ofindications at the top and bottom transitions of the IIE, and at the bottom transition of the hardroll. In the US, the number ofindications at the top of the HE has been ~1% of the total number of detected indications. Indications have been detected in a small number of sleeved tubes after four years of operation. A destructive examination of two tube sections removed from one of the Kewaunee SGs is ongoing, but, the indications have been confirmed by destructive examination to be of ID origin, as were the HE upper transition indications at Doel 4. No indications of OD origin hase been detected on the tubes. These results are in concert with SCC tests which have indicated that the highest level of residual stress occurs on the ID surface of the tube.

S.\APC\cCE95\RAL526.WP6 June 29,1995 I

For the tubesheet expansions in which OD indications have been found, the indicationashave generally been found in sludge pile locations which aid contaminant concentration and increase the metal temperatures. Significant sludge concentration is not expected on top of the TSPs at the areas of the expansions. Some sludge may remain in the crevice at the TSP edge of the expansion although the expansion process is expected to substantially close this gap in the Braidwood 1 and Byron 1 SGs which have no indications of significant TSP corrosion.

Based on the lower temperature of operation and the nominally dry environment, PWSCC of the expanded tubes would not be expected to occur unless a throughwallindication was present in the expanded tube. Even with throughwall penetration, the potential for circumferential cracking is low as shown in the Question 20 response. Based on the lower temperature of operation, lower residual OD stresses, the absence of a sludge pile and the experience base of similar expansions, ODSCC at the TSP expansions would be expected to significantly lower than that for IDSCC.

18. Discuss the relative performance in accelerated laboratory corrosion tests (e.g.,

doped steam), of the proposed expansion with that of HEJ and laser welded joints and other corrosion data.

l RFSPONSE The corrosion behavior of the proposed expansions is based on the information presented on Figures 1017 and 1018 of the report. The doped steam results are given in Figure 1017 and were performed for thermally treated Alloy 600.

For the range of expansions applicable to the TSP expansions,66 and 94 mil expansions developed axial cracks in > 600 hours0.00694 days <br />0.167 hours <br />9.920635e-4 weeks <br />2.283e-4 months <br /> in 400 'C doped steam. This extrapolates to > 30 years at 616 *F as shown in the figure. Although thermally treated tubing times to crack can be expected to exceed that for mill annealed tubing, the data support significant times to crack consistent with the service life estimates given in the response to Question 20. In addition, it is important to note that these tests resulted in axial cracking which is not a concern for the TSP expansions. Recent estimates of service life for laser welded sleeved (LWS) tubes at another plant resulted in a lifetime estimate to the onset of SCC of ~23 years at a temperature of 609 'F for a tube locked at the first TSP (based on an estimated onset of roll transition cracking in 3 years at 620 'F). Based on an activation energy of 35 kcal/ mole, the effect of a change in temperature from 609 'F to 544 *F would result in an increase in the life expectancy by a factor of about 6.6. Although the far field residual stresses would be expected to be similar, a 50% greater residual stresa could be tolerated without reducing the lifetime estimate to less than that of the LWS SMI C\cCE95\RAI_626.WPS June 29.1995 L

tube. Furthermore, the data of the response to RAI 20. indicate that the time to the onset of roll transition cracking at Braidwood 1 and Byron 1 was on the order of 6 years at 619 *F. This would effectively double the expected lifetime of the above calculation. A detailed comparison oflifetime estimates provided in the response to RAI 20. demonstrates that the expected lifetime to the onset of cracking of the exp nsions significantly exceeds the planned operating life of the SGs following the tube / TSP expansions.

19. Discuss the possibility for rolled [ expanded] joint stress reduction by means of heat treatment, if feasible, and the consequent impact on SCC performance.

RESPONSE

The stress corrosion cracking of the expansion would be expected to be enhanced by performing a stress relief operation if the overalllength of the expansion could be adequately heat treated without inducing significantly large far field residual stresses in the tube . This stems from the large amount of data which has demonstrated an enhancing effect of the stress relief of sleeve installation welds. Since the joint is expected to function at a reduced temperature, such enhancement is not considered to be necessary and feasibility has not been established. The magnitude of the deformation associated with the expansion would necessitate that the stress relief he performed over a relatively long length. In order to avoid buckling of the locked tubes during the stress relief, the length being stress relieved is maintained relatively short. Thus, a stress relief would have to be effected in several steps. Each independent stress relief would add residual stresses to the previously heat treated section and the far field stresses. Significant process development efforts would be required to establish a feasible process with a significant reduction in overall residual stresses. The time consumed in performing multiple operations would be significant and is not considered practical.

20. Discuss the bases for an assessment of the anticipated service life for the proposed expanded joint.

RESPONSE

i The anticipated service life of the expanded joint is based on the field operating experience of similarly stressed tube expansions. The proposed expanded joint can be compared with a number of tube expansions performed in the original manufacture of the steam generators, i.e., hardroll, explosive and hydraulic l expansions at the tubesheet. Comparisons can also be made with expansions SAAPC\CCE96\RAI.626.WP5 June 29,1995

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performed as part of a repair process, e.g., hybrid hydraulic /hardroll expansions utilized in HEJ sleeve tube repair, and hydraulic expansions performed in preheater SGs to reduce the potential for tube wear due to flow induced vibration Laser welded sleeve tube repair also utilizes tube expansions as a part of its implementation, but is an inappropriate comparison basis here because of the welding and heat treating processes applied.

Table 201 provides a comparison of the proposed tube expansion properties with those of the processes noted above. The expected stress levels in the proposed expanded tubes are based on Polythionic Stress Indexing test results shown in Figure 1018 of WCAP 14273. These tests for bulged hydraulic expansions represent the bulge AD above the nominal 16 mil hydraulic tubesheet expansion. For comparison with the TSP expansions, which are measured as total diameter increases,0.016" should be added to the AD of the x-axis in the graphs. In Figure 10-18, OD tests extended up to 65 mil AD and ID tests extended up to about 150 mil AD. For the test results in the figure, only expansions with AD > 100 mils exceeded about 40 ksi.

The test results show that the OD residual stresses are effectively bounded by a stress of about 20 kai for bulged tube hydraulic expansions with no significant dependence on the bulge diameter. OD degradation is strongly environmentally dependent especially at these modest stress levels. The absence of a sludge pile at the TSP expansions to concentrate potential contaminants and increase the metal temperatures reduces the potential for the proposed expansions to have OD cracking compared to tubesheet expansions with similar residual stresses. The tube expansion process can be expected to significantly close the tube to TSP crevice for the Braidwood 1 and Byron 1 SGs and thus reduce crevice concentration at the expansion locations.

Thus, in addition to the relatively low residual stresses and the low temperatures for the proposed expansions, the potential for OD cracking is lower at the TSP expansions than for the top of tubesheet expansions due to environmental considerations. It can be noted that only ID indications have been confirmed for HEJ sleeves that have been pulled and destructively examined and no indications have been reported for the cold leg Model D4 preheater expansions.

ID stresses for hydraulic expansions were shown by these tests to be bounded by the minimum to nominal stresses found in hardrolled tube expansions for AD values up to the 90 mil maximum expansion. Up to about 0.060" AD (0.081" total expansion), the indicated ID stresses are in the range of stresses in hydraulic tubesheet expansions. Between 0.080 and 0.100" AD total expansion, there is a lack of data to clearly identify the trend of increasing residual stresses in this expansion range. It can be expected that the residual stresses would be less than 60 ksi at 0.090" AD or comparable to hardroll l

s \APC\cCE95\RAl_526.WP5 June 29,1995 t

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expansions at this maximum expansion diameter. Therefore, for the expected l range of AD for the proposed tube expansions,0.057" to 0.090", the expected stresses are in a range that is between the maximum stresses of hydraulic TS i expansions and the minimum to nominal stresses of hardrolled tubesheet l expansions. This range of ID stresses is comparable to the stresses in l WEXTEX expansions and HEJ sleeve expansions which were determined in l Magnesium Chloride stress indexing tests. Consequently, the most I conservative comparisons, based on the stress level in the expanded tube, is time to cracking adjusted for temperatures of the proposed TSP expansions i against the mechanical hardroll tubesheet expansion followed by the HEJ l sleeve and WEXTEX expansions. The applicable lower bound comparison for the proposed expanded joint is the hydraulic tubesheet expansion. There are approximately 17,000 Alloy 600MA tubes currently in operation in Plant FC that were hydraulically expanded.

More than 100,000 WEXTEX and 100,000 hardroll expansions are in operation, and, in aggregate, more than 10,000 HEJ sleeves are in operation. Among these, the earliest detection of circumferential cracking was in an HEJ sleeve after 1 year of operation in plant E. The cracking in this sleeve initiated on the primary side of the tube, and was shown to occur in a material heat highly susceptible to SCC. The observed degradation of the HEJ sleeves in the plant that experienced the early failure was characteristically different than other operating experiences, indicating that the conditions leading to the rapid onset of the degradation were unique to this plant. Additionally, the operating temperature in this plant is among the highest in operating plants. Therefore, the operating experience of the HEJ sleeve in other operating plants is considered the best basis of comparison for the proposed tube expansion.

In the same plant that experienced cracking in the HEJ sleeve after 1 year of operation, the tube hardrolls experience cracking at the same time as the HEJ sleeve expansion, indicating that the stresses in both the hardroll and the HEJ sleeve expansion were approximately the same. Cracking of the hardrolls in both the Byron and Braidwood plants was first observed after approximately 6 years of operation; therefore the Byron and Braidwood hardroll experience is another basis to estimate the aniticipated service life of the proposed tube expansion.

Table 201 shows the expected factors on time, based on the bounding values of activation energies and the operating temperatures of the plants experiencing circumferential cracking for each of the types of tube expansions discussed above. An operating temperature of 544 'F, the saturation temperature for Byron and Braidwood, was used as the basis of comparison because the expanded tubes will be plugged and will be operating at this temperature.

These factors, applied to the minimum operating time to cracking for each of 8.\APC\CCE95\RA1.526 WP6 June 29,1995

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the comparisons, is the basis for estimating the expected life of the proposed expansions. Based on the minimum field experience time for reported  !

circumferential cracking multiplied by the time factor of the table, excluding )

the single HEJ comparison that is considered to be a unique event, a I conservative estimate of the minimum time for anticipated onset of cracking in the proposed tube expansions is greater than 20 years for the most conservative estimate of activation energy. The minimum time occurs for estimates based on the ID cracking of HEJ sleeves while ID cracking of  ;

hardroll expansions projects to about 35 years to cracking for the TSP expansion. Projections based on OD cracking in hardroll transitions would be about 28 years for the TSP expansion. However, based on lower residual OD stresses, the absence of a sludge pile and the experience base of similar sleeve expansions as well as the lower temperature of operation, ODSCC at the TSP expansions would be expected to be significantly lower than that for IDSCC l and times to cracking would be expected to exceed the 28 years projected from the hardroll expansion field experience. It can be noted that even the most conservative single HEJ occurrence implies a time of 9.2 to 23.7 years for cracking at the plugged tube temperatures.

Estimating the anticipated onset of circumferential cracking based on operating experience is considered valid since, in general, inspections for circumferential cracking on, at least, a sample basis is a normal part ofin-service inspections.

If a significant population of tubes experienced circumferential cracking, even a sampling inspection would identify this degradation and the sample would be expanded. Therefore, the data in Table 10-5 of WCAP 14273 confirm that circumferential cracking has been relatively rare within the operating experience of these plants, and also, slow in progression. Only the plants with the least operating time to detection of circumferential cracking were used in Table 20-1 to estimate the time to anticipated onset of circumferential cracking in the proposed tube expansions. The remainder of the plants noted on Table 10 5 provide confidence that rapid circumferential degradation would not be expected.

Additional corrosion tests have been performed in support of the tube expansion for the preheater support plates which are also supportive of the proposed expansion. While these tests do not directly enter the service estimates of this response, they provide supplemental data on the acceptability of tube expansions at TSPs. A brief summary of the tests and results is given below:

Denting Dependence on Tube to TSP Clearance The proposed expansion process, like the preheater process, essentially closes the tube to TSP gap. The test was performed to determine if reductions in the gap influenced the susceptibility to denting and was SA.APC\ccE95\RAL626.WP5 June 29,1995

- performed using a reference denting chemistry. The test results show that the' dent size at a given time is reduced by about a factor of five for a 3 mil diametral gap compared to a 25 mil gap and a 2 mil gap has about have the dent growth of a 5 mil gap. Thus the expansion process reduces the 4 potential for denting at the TSPs. l Dependence Upon Open, Partially Filled and Packed Crevices 4 Expansions could be performed into crevices that vary from open to fully  !

packed crevices. Polythionic acid tests of sensitized Alloy 600 tubing were performed to compare the times to crack for specimens expanded into open, partially sludge filled and packed crevices with initial 40 mil tube to TSP diametral gaps. The results showed no significant differences in time to.  ;

crack between the different crevice conditions. Additional tests were l l performed for expansions performed into previously corroded TSPs and these results also showed no differences in time to cracking compared to uncorroded TSPs. 1 I

Comparisons of Time ;o Crack of Hydraulic and Hardroll Expansions In addition to stress indexing tests that show lower residual stresses in hydraulic expansions compared to hardroll expansions, time to crack comparisons were made in various test mediums for diametral expansions up to 40 mils. The test mediums included magnesium chloride (SS304),

polythionic acid (Alloy 600) and controlled potential tests (Alloy 600) in - ,

10% sodium hydroxide. All test results demonstrated that the hydraulic  !

expansions have less cracking susceptibility than mechanical expansions. l

- 21. Discuss the consequences for postulated through wall circumferential cracks in the parent SG tube material in the rolled area above and/or below the TSP during an MSLB.

RESPONSE ,

Tube expansion has been implemented very conservatively for the Model D4 SGs to accomodate postulated through' wall circumferential cracks even if it is further postulated that the cracks grow to cause severing of the tube at the expansions. As shown in WCAP 14273, Section 10.6, and responses to other questions in this RAI, significant circumferential cracking is not expected at ,

the plugged tube temperatures of the expanded tubes.

- The expandcd tube regions at the TSP intersections include a sleeve stabilizer.

Assuming that the tube becomes severed at the expansion, the sleeve functions as a stabilizer and prevents the severed tube end from contacting the adjacent SMPC\CCE95\RAI.526.WP5 June 29,1995 ,

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tube. Thus, the potential for propagation of damage to adjacent tubes is prevented.

Throughwall circumferential cracking at the expansions does not impact the TSP " locking" function unless the cracking propagates to near or complete severing of the tube. The tensile load in an expanded tube during the SLB event is < 400 lbs (Table 8-13 of WCAP 14273). Thus, the SLB event will not cause severing of the tube unless the circumferential crack has essentially caused severing prior to the event. Early detection of moderately sized circumferential cracks is therefore not critical to the " locking" function.

The largest displacements of TSPs 3 and 5 (lower two TSPs above the FDB) are independent of circumferential cracking even if the cracking progresses to severing of the tube. The SLB displacements for these two TSPs are in the downward direction toward the tubesheet (Table 8-9) and the sleeve prevents lateral displacement of the severed expansion. Therefore, the expansion performs it's locking function even for the postulated severed condition. The Braidwood 1 and Byron 1 indications are dominantly at these lower two TSPs with 85% and 89% of the indications, respectively.

In addition to the above benefits of the sleeve stabilizer for reducing the consequences of postulated circumferential cracks, the tube expansion design .

has conservatively included redundant tube expansions to further accomodate postulated severed expansions. The following summarizes the TSP displacement insensitivity to postulated severed expansions:

Severing of all expansions on the 5 redundant tubes (5 of 21 expanded tubes) results in negligible changes in TSP displacements and the maximum TSP displacement remains at 0.094".

Severing all expansions at TSPs 3,5 and 7 on 6 of the 8 tubes expanded at these elevations results in negligible changes in displacements (maximum displacement < 0.11"). This analysis did not take credit for the sleeve preventing tube separation at the lower TSPs, in which case the maximum displacements would not increase at the lower two TSPs.

Severing all expansions at TSPs 8,9,10 and 11 on 12 of the 17 tubes expanded at these elevations results in maximum displacements of < 0.2" which is less than the acceptance limit of 0.31" for tube expansion to limit  ;

the tube burst probability to < 10-8 even if all TSP intersections are assumed to have throughwall indications.

Severing of all expansions except for the 7 tube locations with duplicate expansions results in maximum displacements of < 0.2".

SAAPC\cCE95\RAI.626.WP6 June 29,1995 i- -

In summary, the sleeve stabilizer prevents propagation of damage to adjacent tubes if an expansion is postulated to sever, the lower two TSP displacements  !

are independent of postulated severed tubes and the redundancy in the tube  ;

expansion design permits large numbers of severed expansions while still l maintaining acceptable TSP displacements.

i i

22. Discuss the requirement, including the rationale, for periodic inspection methods, sample size, and frequency schedule for the proposed tube expansions.

RESPONSE

The periodic inspection requirements for the expanded TSP intersections are that a minimum of three expanded tubes shall be deplugged and inspected at every third planned inspection following tube expansion. If a circumferential ,

crack is found at an expanded intersection, the inspection should be extended to other expanded tubes in the SG with the inspection scope extension dependent on the severity of the indications found in the base inspection.

Following identification of circumferential cracks, the adequacy of the tube expansion matrix for limiting TSP displacements should be evaluated under the assumption that a circumferential crack further develops to a severing of the j tube at the expansion location. )

l l

The estimates of the times to potential circumferential cracking in the plugged j expanded tubes are in the range of 15 to 48 years as given in Section 10.6 of i WCAP 14273 and the response to Question 20 of this RAI. As noted in the response to Question 21, an expansion must essentially sever to affect TSP displacements and the consequences relative to limiting TSP displacements are minimal even for a large fraction of the expansions postulated to sever due to circumferential cracking. While plugging the expanded tubes reduces the temperatures which results in the long times to potential cracking, an inspection of the plugged tubes requires deplugging, inspection and replugging of the tubes to be inspection. These operations result in significant personnel exposure and outage time. Based on the reasonably long time to potential circumferential cracking and the minimal consequences of a significant number of large circumferential cracks, a periodic sampling plan for inspecting the expanded intersections in the plugged tubes is adequate. A period of three operating cycles has been selected for inspecting the expansions. This corresponds to the shortest time period that circumferential cracking has been l observed in operating SGs and has occured in hardroll expansions, which bound the rant t of residual stresses for the hydraulic TSP expansions and operate at temperatures of about 618'F compared to about 540 F for the plugged tubes.

Thus the inspection period providea large margins against anticipated times for circumferential indications at the TSP expansions.

SAAPC\CCE95\AR526.WPS - June 29,1995

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• j The three tubes selected for inspection are anticipated to be tubes with the most expansions in the tube, which is six expansions at four tube locations.

These locations include generic tube expansions at R6C7 and R5C110 (see Table 21 of WCAP 14273) which are the more important locations for limiting TSP displacements. The inspection would then encompass inspection of 18 of the 72 expanded intersections in the SG that is inspected. If an abnormal expansion is identified during implementation of the expansion process, the tube with this expansion could be substituted for one of the tubes with six expansions. In this case, a minimum of 14 expansions would be inspected.

This minimum sampling plan of about 5% of the expansions is adequate for detection of initiation of circumferential indications given the minimal sensitivity of the expansion design to the consequences of circumferential ,

cracking. I If a circumferential crack is identified in the sample inspection, the inspection would be extended to other expanded tubes based on the severity of the indications found in the base inspection. For example, if a single small crack is found at the lowest two TSPs (hottest TSPs that are most likely to have initial cracking), there is no impact on the TSP displacements and the extension of the inspection would be limited to two tubes titat would include the redundant tube if present in the expansion matrix. On the other extreme, if circumferential cracks are found in both the reference tube and its redundant tube above the second TSP, the inspection would be extended to all expansions in all SGs since this represents a potential impact on the TSP displacements if the indications would grow to a severed tube. Due to the deplugging and replugging ,

operations required for inspection, it is important that the extension of the '

inspection be related to the severity of the indications found. Due to the large number of combinations of indications with varying severity implications, the selection of tubes for an extension of the inspection is best made at the time of the inspection. Similarly, the time frame for performing additional tube expansions if circumferential indications are found in an inspection can best be determined based on the results of the inspection and the severity of the  ;

indications. l Overall, the proposed inspection plan provides adequate inspection for the I expanded tubes given the low temperatures with associated long times to crack l initiation and exr.ected low growth rates, the insensitivity of the TSP l displacements to severed expansions and the redundancy in the number of tubes expanded.

23. Provide a more comprehensive discussion and analysis of a postulated circumferential severance than that provided in Section 10.5 of WCAP-14273, including clarification of the following areas:

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a. The discussion in WCAP 14273 is apparently limited to " lower" TSP -

elevations. - What is the basis for this limitation and what are the limits to be imposed on your proposed SG tube expansions?

b. This discussion assumes that the SG tube is severed at the upper edge of the expansion. Discuss the case of a tube severed at the lower expansion point below the TSP and the case for a double circumferential severance where a detached tube segment is left between two TSPs.

c. It is stated in WCAP 14273 that the potential for fluid elastic excitation is ' minimal'. Provide a discussion of exactly under what conditions this is possible. Provide a more quantitative analysis.

RESPONSE

The discussion in Section 10.5 of WCAP-14273 is intended to be a general discussion of the performance of the sleeve in the expanded tube as a tube stabilizer. For discussion purposes, the section postulates a tube severed at the top edge of the TSP (bottom edge of bulge) or at the upper edge of the bulge above the TSP. These specific locations were chosen only to focus the discussion. The points made for the sleeve stabilizer apply to the bulge above or below the TSP and at any TSP.

In response to point a. above, the discussion applies at any expanded TSP and is not limited to the " lower" TSP elevations.

In response to point b. above, the discussion includes considerations of a severed parent tube at the TSP edge of the bulge, at the bulge end away from the TSP or anywhere within the bul.e. The discussion focused on the bulge above the TSP only as an intended means to simplify the discussion with a specific mentalimage of a bulge location. The effectiveness of the sleeve as a stabilizer applies to all expanded locations including the lower expansion point below the TSP as a symmetric location to the upper expansion point'above the -

TSP. For the case of a double circumferential severance where a tube segment is left between the two TSPs, the discussion is still applicable. The segment is trapped between the two TSPs by the sleeve with an essentially zero gap--

between the sleeve and the ends of the tube segment at the bottom of the upper TSP bulge and the top of the lower TSP bulge. The sleeve continues to prevent lateral motion of the tube segment and thus functions as a stabilizer.

c.

In response to point c. above, the discussion applies the word " minimal" to .

> mean that fluid elastic excitation is not expected. To function as an effective tube support point for fluidelastic vibration, a point of contact is required but a SAAPc\ccE95\RAL526,WP5 - June 29,1995

significant preload at the point of contact is not required. At the edges of the expansion bulge or within the bulge, the sleeve to parent tube gap is essentially zero due to the plastic deformation at the expansion. Thus tube to sleeve contact is expected at the parent tube locations where a postulated severing of the parent tube could occur. The negligible tube to sleeve gap and the typical small changes in lateral alignment between plates due to heatup effects provide a support point for the severed parent tube which would prevent fluidelastic vibration. The use of the words minimal and essentially eliminate fluidelastic vibration are only intended to recognize the fact that tube vibration is a function of tube support probabilities and there may be some small probability of a gap between the sleeve and tube that could permit fluidelastic excitation.

Even if the postulated parent segment is assumed to be fluidelastically excited, the small sleeve to parent tube gap would limit the vibration amplitudes to negligble levels relative to considerations for wear of the sleeve.

24. You state in WCAP 14273 that it is critical to limit the degree of SG tube expansion and to modify the expansion depending on variables such as the tube yield strength. Accordingly, what would be the consequences of a failure of the computer controlled SG tube expansion monitoring process? What inspection measures do you propose to detect any failures in the expansion process?

ItESPONSE As discussed in Sections 10.4 and 12.4 of WCAP 14273, all expanded TSP intersections are inspected by bobbin coil profilometry following application of the expansion process to verify that the required bulge sizes have been achieved. Verification data on the accuracy of bobbin profilometry for sizing the expansions is given in Section 10.4.2 and a discussion of the application of the bobbin profilometry data for expansion process verification is given in Section 10.4.3. The bobbin profilometry data provides the diameter of the bulge. The yield stress for the tube is obtained from the expansion process pressure versus time as exceeding yield causes an inflection point in the curve which is related to the yield stress. From the yield stress for the tube, the required expansion diameter for comparison with the measured data is obtained from the process requirements developed as a part of process qualification. Pull force as a function of bulge size and material yield stress, such as Figure 10 4 of WCAP 14273, is developed from the process qualification data to define the required bulge sir.e as a function of yield stress.

From the above profilometry data, an expansion that is smaller than the requirements can be readily identified to indicate a probable failure of the computer controlled expansion system.

SMPCNCCE95%RAI.526.WP5 June 29,1995

An undersized bulge will be evaluated for appropriate corrective action. If the diameter is small enough, reexpansion would be attempted. If reexpansion is not successful, the specific joint location would be evaluated to determine the need for further action such as expansion of another tube. Not all expansion locations require the minimum stiffness requirement used to establish the bulge diameter requirements as the stiffness requirement was developed for the limiting TSP expansions. On a case by case basis, the measured bulge diameter and associated stiffness can be evaluated for limiting TSP displacements to acceptable values as an alternative to expanding another tube.

In summary, the post-expansion bobbin profilometry inspection provides the basis for accepting the expansion or defining the need for corrective action if the required expansion diameter is not achieved.

25. Discuss the types of additional stresses that would resuL from various configurations of SG tube " lock up" conditions that would result from the tube expansions with respect to stress corrosion cracking in the tubes and stressing of the TSP, wrapper, and shell supports.

RESPONSE

Section 8.11 of WCAP-14273 provides a lengthy discussion of the effect of expanding tubes on the structural integrity of the SG components. Accordingly, much of the response to this RAI is a summary of the information presented in that section. The assessment of the effect of the expansions considers two limiting conditions, the case where the active tubes in the SG are not already locked to the TSPs, and the case where the active tubes are locked to the TSPs.

For either case, the expansion of the tube during the loching operation must be accompanied by Poisson contraction of the tube in the axial direction. The expansion process will be applied to the uppermost tube / TSP intersection first, with subsequent tube expansions progressing down the tube. This is necessary since the expansion joint includes an integral sleeve to act as a stabilizer and to increase the axial stiffness of the tube. The Poisson contraction of the tube during the expansion process results in a residual tensile stress being developed in the tube (s) being locked in each tube span between the TSPs.

A discussion of the effect of the residual stresses on the expected SCC performance of the expanded tubes is provided in the responses to several other of the RAIs, e.g., RAIs 18 and 20.

Case 1 - Existing tubes not locked to the TSPs -

S.\APC\CCE95\RAL6t3 WP5 June 29,1995

l This is the easiest case to address, although it is not likely to be the case in the Braidwood 1 and Byron 1 SGs. Relative to stressing cf the TSP, wrapper, and shell supports, the expansions would be expected to induce less loading on the SG structures than that from locked tubes during the shutdown (cold) condition. This is because the number of tubes to be purposely locked is much less than the number of tubes typically seen to be locked to the TSPs through crevice packing or denting. In essence, it is believed that the position of the TSPs at shutdown conditions, except at locations adjacent to the stayrods, in a SG in which the tubes have become locked to the TSPs by natural causes, e.g.,

denting, is dictated by the thermal contraction of the tubes relative to the TSP, wrapper and shell supports. This is because locking / denting of the tubes in the TSP holes occurs at the hot condition. The coefficients of thermal expansion of the carbon steel support structure and the Alloy 600 tubes are 7.76 10'8 and 8.16 10'8 respectively considering the carbon steel at a temperature of 550 F and the tubes at a temperature of 600*F. In addition, since locking / denting occurs during operation, the tubes are axially strained by the internal pressure at the time of locking / denting. The net effect of both conditions can result in a cold preload of the tubing on the oider of 1100 lbr per tube. This is about the same order of magnitude expected for each expanded tube in the hot condition.

Thus, the locking of tubes to the TSPs due to natural causes would be expected to result in significantly larger loads than the application of the expansion process to the small number of tubes planned for at Braidwood 1 and Byron 1.

In other words, "the addition of expanded tubes does not add any significant new loading mechanisms to the TSPs" during normal operation of the SG.

Furthermore, the interaction of the TSPs with the wrapper will continue to be governed by the twelve (12) stayrods, which each have an order of magnitude greater stiffness than an expanded tube, and the backup bars and wedges.

Case 2 - Existing tubes already locked to the TSPs The field inspection data of the Braidwood 1 and Byron 1 SGs indicates that several, i.e., most, of the tubes are likely to be already locked to the TSPs. This condition occurs as a result of crevice packing (the mechanism that leads to ODSCC of the tubes at the tube / TSP intersections), which may be considered as incipient denting, and/or light or heavy denting of the tubes at the locations of the TSPs. The extent oflocking may be less at Byron 1 than at Braidwood 1 because of the recent chemical cleaning operations. Field experience has demonstrated that this phenomenon does not occur to isolated tubes, but does involve the participation of many tubes. As discussed in the previous paragraph, the expansion of a tube at the TSPs will result in Poisson contraction in the axial direction. Thus, the expanded tubes will experience a higher tensile stress in the cold condition than that being experienced by its locked neighbors. When the plant returns to power, the TSPs are returned to their normal position in the hot condition by the locked / dented tube population.

SAAPC\CCE95\RAL526.WP5 June 29,1995 1

Now, however, loads imparted by the expanded tubes are resisted by the stayrods, the backup bars and wedges, and the locked / dented tube population.

Again, "the addition of expanded tubes does not add any significant new loading mechanisms to the TSPs" during normal operation of the SG.

In summary, the application of the tube expansions would not be expected to introduce significant new loading mechanisms to the TSPs, or to the TSP's supporting structure. This conclusion is independent of considerations of the active tubes being free or locked / dented at the TSP intersections.

26. Data are presented in Table 10-5 of WCAP 14273 to demonstrate the lack of service cracking in hydraulically expanded SG tubes. Exactly how many mill annealed SG tubes in these plants from which this data was taken had received a fully effective inspection for circular cracking and for how many cycles?

RESPONSE

The data of Table 10 5 have been reduced to the most limiting plants for each type of expansion and updated to the latest information in the responne to Question 20. Table 201 identifies the numbers of tubes for each type of expansion that are in service and all data in the table are for mill annealed tubing. At the years given in Table 201 for minimum time to detection of circumferential cracking, effective RPC inspections were performed that encompassed most hot leg expansions. When circumferential cracks were detected, large extensions ofinitially planned sample inspections were performed and the extensions typically resulted in 100% inspections of the hot leg expansions. Prior to detecting the circumferentialindications, the inspections were typically sample inspections ranging from a few hundred tubes to 20% inspections.

The preheat SG TSP expansions have been most extensively inspected at Doel-4, which has the highest operating temperature. Sample inspections have included UT inspections and no crack indications have been found. Other effective inspections of the preheater expansions have been limited since the guidance for inspecting these indications is based on finding cold leg cracking at the hardroll tubesheet expansions with higher residual stresses and no cold leg indications have been reported in Model D4 expansion transitions.

l l

l l

1 SAAPC\cCE95\RA1.526.WF5 June 29,1995

p.

600.00 Tube Node 1 (Near Tube Sheet) 550.00 - - N -

_hw -- . ,;

2nd Fluid Node 37 500.00 - -

(Just Above Tube Sheet) 450.00 - -

S i

o 3 400.00 - -

E.

E 350.00 - -

300.00 - -

250.00 - -

200.00 0 1 2 3 4 5 Time, second Figure 4-1 Transient of Tube Wall and 2nd Fluid Temperature (Tube Node 1,2nd Fluid Node 37)

NRCBWC1N.XLS 6/29/951:12 PM

I 600.00 Tube Node 9 (at U-Bend, Hot Leg) 550.00 - - '--- , _

2nd Fluid Node 29 l 500.00 - - (U-Bend) l 450.00 - -

l u. .

! O l e.

a 3 400.00 - -

8. l E

350.00 - -

300.00 - -

250.00 - -

l 200.(X) 0 1 2 3 4 5 j i

Time, second j t Figure 4-2 Transient of Tube Wall and 2nd Fluid Temperature (Tube Node 9,2nd Fluid Node 29) j I

NRCBWC12.XLC 6/29/951:15 PM

600 Tube Wall 550 - -

2nd Fluid 500 - -

450 - -

u_

o E.

a 3 400 - - -

8.

e 350 - -

300 - -

250 - -

l 200 0 50 100 150 200 250 300 350 l

Length from Tube Sheet, inch Figure 4-3 Spatial Variation of Tube Wall and 2nd Fluid Temperature along Hot Leg Tube i

I NRCBWC13.XLC 6/29/951:21 PM

600 Tube Wall 550 - -

2nd Fluid 500 - - .

450 - -

E.

a 3 400 - -

8.

E O

350 - -

300 - -

250 - - !l i

l 200 0 50 100 150 200 250 300 350 l l

l Len91h from Tube Sheet, inch Figure 4-4 Spatial variation of Tube Wall and 2nd Fluid ,

Temperature along Cold Leg Tube NRCBWC14.XLC 6/29/951:18 PM

w \

I 7-i c-v_ M_ %+

N  ;

1

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Nboh N tov Figure 13-1 Model D Tube Support Plate Tube and Flow Hole Geometry

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sheet 1

...-,~.m. .m..n....-- . . . . ..a, ...E.,, .nus..~n.

Table 41 Tube Wall and Secondary ___ _

EuMTemperature along Tube Length r

r 5UTime =) 0(second

_.ECSN!g Tute Wall __2nd Fluid Temp _ _ .

z, inch TEMP TEMP Difference

. 3.375 _.557 557 . . -..0

.r

$4.75 .- . 557 . tI. - .557 1

. . . + - - _ -

0

._..___+...___._.___. . _ . . . . . . . . . . _

l30.25 557 557 0 173.25 557 $57 0 216.25 557 557 0 259.25 557 557 0 325.85 557 557 0 325.55 557 557 0 2d25 557 557 1 0 216.25 557 557 0

._17 3. 25_ ,. 557 _ _ _557_ , _ _ _0

_t 30.25 _ 557. _ _ 557_ _ .__ __0

_ _ . 9923 .. 557___, 557 __ _

0 81.75 557 557 J

_ 63.75 _ _ 557 .,_ 557_ __. O

_ 45.75 557 557 0 27.75 557 557 . 0 12.75 557 557 0 3.375 557 557 0

~ ,... .__ _ _ _ . _ . . . _ . . . .__.- [ __

At Time =, _ 0.01793 _ second l __

. _ _ _ . _ . _ . _ . _ _ a _ _._ _. k Elevation Tube Wall. 2nd Fluid. Temp __

-. r. mch { . TEMP - . . - EMP T_-.. i Diff_eren.ce. .. -

..4._.

21.75 557 557 _ f , _ _ 0 54 75 557 557

{ _ _ _0

_ 90.75 _ _ ,_ 557 _ _ _557,_[ 0

.130 25 . _. . 557 . i .-. 557 ._ .0

. .. . . . . - . . ~. - . . . . . .

216.25 557 $57 0 259.2.5. _ - - 557._ _ . _ _5_5__7. .

7____ _ .0 325.85 _ _ 55_7, _ 557 1 0 325 85 .557 i

_. 0

. ... 557 ._ f.

._173 25_ f, _ , 557_], 557 0 130.25 557 557 I O 99.75 557 -

557 _j 0 81.75 557 0

' 557 ] ___

63 75 557 _ 557 _

0 45.75 557 I 557 6 0

Page 1

Sheet 1

'4 27.75 . 557 l 557 1 0 12.75 557 I $57 b 0 3.375 557 . 557 0

. ~ . - _ . -.. .

At Time = 0.0686:1 second Elevation Tube Wall 2nd Fluid 4Temp _ _

_ y, inch TEMP TEMP Ihfference 3.375 f 557 556.984 0.02 21.75 557 556.98 0.02

_ 54.75 557 556.98 0.02 90.75 557.01 556 98 0.03 130 25 557.01 556.98 0.03 173.25 557.01 556.98 0.03 216.25 557.01 556.98 0.03 259.25 557.01 556.98 0.03 325.85 557.01 556.% 0.05 325.85 ,__ 557.01 556.% 0.05 259.25 557.01 556.98 0.03 216.25 557.01 .

556.98 0.03

_._ _3 73:25 _-,, 557.01 556.98 __,0 P3 130.25 g _557.01 j $56.98 0.03 99 75 557.01 1 556.98 0.03 Bl.75 557.01 i 556.98 0 03 63.75 ,55,7[0} 556.98 0.03 45.75 557.01 i 556.98 0.03 27.75 557.01 556.98 0.03

. _ 12]5,_, _ _,_55731 _ j56],8 _

a03 3.375 55741 556.98 0.03

~

At Time = 0.1808i second Elevation Temp Tube WallJ 2nd Muid

. .. _hj"C_h__.,,_ TEMP ._ TEMP thfference _

_ _ 556;96 _ _ _, _0 3

21.75 } 75 _. l556.99 ., 556.96, . 556_.96 _ .,_ 0,03 54.75  ! 557.01 , 556 % 0.05

. 3 ;75 .. [_ .j5]p1,,___j,, _ {563_,_ . p,9) 130 25 l $$7.0t { $56.95 0 06

._173.25. . [__ $57.0.i .. 1 556I92 d5 ]_ _'

...__.w_. p...____+...__._.

325.85 1.16 556.97 [ 555 81 ,1

._ 325;85 ., 556.97. ] 555 81 _ l.16 259.25 557 556.48 0.52 216.25 557 01 _

556.74 _ [ _ 0.27 173.25 557.01 l 556 92 0.09 f

va25T55gcn653;365

. _9 92 ! i . 557;p[ 1 __{56 1 , ,.

ROS 8375 t 61.75 d. 357pt.i_536.96_f_

557 01 J 556 96 0 05 _ _0.05 45.75 557.01 1 556.% I 0.05 Page 2

7-

.. +

1 Sheet 1 '

2 557AI_

. - .7;7_5_ 556.% J 0,0_5 12.75 557.01 556.% 0.05 0.0,5

. 3.3 75,__ .. ,_{57.01_i_,_556.% i 4._ _ ._- 4 JJjme n._._,0;34178_l._ second_[ _

..b.-..~

.._EMyp@n . _Mg_q[j 2nd Fluid { Terne __

_.2'I"'.h ,,, ., ,,,_ ]@,P ., ,,,,_TEp1 P : Difference _

3.375 556.83 556.4_7 _ 0.36 556.92 556,43 0.49

._ 21.75 54.75 556.97 556.33 0.64

_. 90.75 130.25 I 556.98 556.%

556.2 556.03 '

{ 0.78 O.93 173.55 556.92 555.8 112 216.25 556.87 555.61 1,_ _f.26 259.25 f 556 8I ..~ _555.37.--..pl . _1.44 - .

325.85 556.68 554.68

( _. _2

, 325 85 556.68 554.68 J_ __2 259.25 556.81 555.37 { _ ,. 1.44 216.25 556.87 555.61 1.26

_._I73 2(_ _ 556.92 555.8 1.12 130.25 556.% $55.95 1.01 99.75 556.98 556 06 0.92

_ 81.75 556.99 556.32 0.67 556.5.i 0.47

. dM1_...h ._557 45.75 Q57.01 556.63 ._ 0.3,8 27.75_,[ $57.01]_556.68 _ . 0.33

. )22b ._ _. 557 556.53 0 47 3.375 557 556.55 0.45 At Time = 0.57172 second

- .- . - . b E_levation_ Tube Wall _2nd Fluid Ternp r, inch TEMP TEMP Difference 3.375 556.83 $$6.47 0.36

---. . -+

21.75  ! 556 - 92.-.

556.43 . . -

0.49

. _$4_75._ ! $.5_6.97. _..556.33 _ .- _0.64

_ _90.75_g,J56 98 _ 55_6.2 . _ , _ _ _0; 7 8 130.25 ! 556.%~~j 556.03 0.93

,1525 7.4..55A,.i_2_ . f ~5_55.T.. ..... ~.~_~~. .. .~. E12 216.25 _ [ 556.87 . 555 61 { l.26 259 25 556.81 555.37 I 1.44

< 325 85 556 68- 554.68 l 2 325.85 556 68 554.68 I- ~2

_ 259.25 _ 556.81 1.44 555.37 [

21t 25_ 556.87 _555 61 _ p _ i 26 173.25 556.92 555.8 j_ l.12 1.01 130 25 ] _ 556.32

_99g5_. 4 . 55g98

.. 8!:32.._.j__556.99 I 556.96 556 06

_ 555.95 ] (

0 92 0.67

_9p5 _ .___ _551_[_ 556 53 ( 0.47 45.75 557.01 l 556.63 I 0.38 Page 3

• .~

Sheet 1 27.75 557.01 556.68 1 0.33 12.75 - - 557 556.53 0._.47

-3375 557 . 556.55 0.45 At Time = 0.80629 second i

_ _ _ _ _._.__m.___

Eleta3% Tube Wa,ll__,.,,2nd p pi , Teny, _ _

_puh _,_,._,yyP ,_ _ . TEMPgDdference_

3.375 555.19 551.98 3.21 21.75 555.31 552.02 3.29 54.75 555.4 551.99 .

3.41 90.75 555.47 551.92 3.55 130.25_ .-_,_5. 5._5.48. m .._. 551_.83_ . L, 3 ._b _.65 173.25 555.46 551.71 ,_{_ 3.75

- 21q25 .__qy,4__, _,,5)1,5,6_ _ -, , ,3;84 259.25 5553 55137 3.93 325.85 554 99 550.79 4.2 325.85 554.98 550.79 4.19 259.25 [5I. 58 551.37 5il 216.25 555 39 I 551.56 g_ _ 3;8,3 173.25 555.47 551.71 l 3.76 lj0.2.

555$5h 7 5188 ] 339 99.75 555 62 551.98 I 3 64 81.75 555.72 552.17 3 55

- f3. 75._.- .. 5. 55.78.. - .5._52_.26 .-. 3.52

........__.p.._.._...4-_.._.-_ - . _ . -

27.75 l 555.82 552.34 3.48

_12.75_.,j $55.54 ,, 551.8 1 74 3.75 3 375 _. { 555.52 _ _ 551.77 _

___...__.h_.____ _. . _ . _ _

At Time = l 1.04546 second

. __. . . _.w._

i ._.._f, . [

I Elevation Tube Wall {2nd Fluid jTemp,_ _

___* d.mh _,,,TMP,, _.. ,, _ TEMP.,,,Jffe,qnce__.

3.375 553.99 549.79 4.2 21.75 554.07 549.83 4.24

_ $4752 ,_ 554;l3 ,j 549.81 4 32 90.75 - 554.2 l . 549.75 4.45 7

.$ . . bh.$k . . '__ b& .. .. _.h 216.25 ! 4.75 554.17 J __549 42 ] _

_259g5J 554 w, ..[ __ 549 24 L _385 325 85 553.77 I 548 68 I 5.09 3 _25 8_5 _..._5.53._.75 I _g 5.48.__68 T. ._.

_. 5_.07 259.25 ._ 5,5105 1, p{9;24_,, __ _ ,,_ 4 _8_1 216 25 554.16 l 549.42 .

__ 4_.,74

_ 172 25 .,_ 554.26 L 549.56 ' 4.7 130.25 55436 ! 549.72 4.64 99.75 ~ 554 42 549.81 4 61 81.75 554.52 549.97 4.55 63.75, _ . 554.57 l 4.52 45.75 1 554 6 J 55005_I 550 09 4.51 Page 4

i I

t 1

.. Sheet 1 I 27.75 554.62 550.11 4.5_I

_ 12.75 554.37 549.63 -4.74  ;

_ 3.375 554.36 549.6 _ 4.76

- At Time = 1.28899 second

.. . ._ d ._

Elevation j, Tube Wall _2nd Fluid jTemp__

_._ r, inch _ .. TEMP TE,M,P,,,][hfference__ .

3.375 21.75 552.68 552.71 547.55 547.61

{ 5.13 5.1 54.75 552.73 547.59 6 5.14

_ 90.75.

--- . -.-. 55279 54

---._7._54. l-- .- 5 2. 5 .

130.25 552.81 54.7.46 5.35 173.25 552.81 547.36 5.45 216.25 552.78 547.24 5.54 259.25 552.7 547.07 5.63 325.85 552.41 546.55 5.86 325.85 ___ 552.39 _ 546.55 5.84

_.259,25 __ 552.68 547 07 __. 5.6_1 216.25 .

552.79 547.24 5.55 173.25 552.87 547.36 5.51 130.25 552.98 547.51 5.47

_ 993 5__. 1 .. 55133

- 547.59 _ _.__5.44 81.75 553 12 547.72 5.4 63.75 553.17 547.78 5.39

_.._.4125.__, __ _5,5 3 2_,,,_ .. 547.81 _ 5.39 27.75 553.22 547.83 5.39

..._ _ _12.75 552.98.. ,. 547.4 l 5.58

- 4_

3.375_ 552.% 547.35 5.61

. At Time =_J _ l.53692_ .

accond_, . _ _

Elevation Tube Wall 4

2nd Fluid Temp _ _ '

.gjnch,j_. TEMP _(TEMP _ jlhfference

_ 3;375 ] _ 551.34 545.35 i 5.99 21.75 551.32 545.4 5.92 '

_ . 5y5 ,_ ,, 551.29 $45.38 5.91 90,75 551.3 545.33 5.97 130 25 551.29 545.25 6 04 545.16 6.12 173.25 [ 551.28 216.25 551.24 545 04 6.2 259 25 551.18 544.88 6.3' 325 85

_ 325,85_ $50.9 ._ _,_544.39 j _

550.91 544.39 (( 6.52

_,6.51

_259 25_ j_ 55g8_ _ , 54188._L __. _ 6 3

-216.25 . . .551 28 -.- 545 04 ..}

6.24 130.25 551.46 545.3 6.16

_ 99.,75 _ 351.51 545.37 _

6.14

_ 81.75 551.59 _ 1 545.48 6.11 63.75 551.63 545.53 6.1 45.75 551.66 545.56 6.1 Page 5

Sheet 1 27.75 551.68 j_545.58 - } , _ ; _, . 6.1 12.75 551.49 545,23 6.26 3.375 551.46 545.17 6.29

_At Time = 1.78923 _second Elevation Tube Wall 2nd Fluid Temp TEMP , TEMP Di[fr nce _

_ 3.375 551.34 545.35 g_ 5.99 21.75 _ 551.32 545.4 5.92

{i

- _54.75_._ p 551.29_ ._.9 ._5._4_5.38 .g ._. _ _ . 5_.9.- 1 -. .

545.33 5.97

_ 90 75 _ 4 551.3 4 I

.l}025 551.29 _ ._,545.25 ._ _ 6.04 173.25 551.28 545.16 6.12 2_16;2jg_55,{p _ 545.04 _6.2

_ .-. 25,9_25 , _ .,_,5 51.18 _.,_, ,,$_44.88 L __6.3

_ 325 85 __,_ 550.91 544.39 ] _ 6.52 6.51

_325_81_,._,_,jR9_,_ y 544.39 j

.._.259.25 551.18_ _ 544.88 _ ._ , _ 63 216.25 551.28 545 04 6.24 173.25 55136 545.16 6.2 130.25 551.46 545.3 6.16 99.75 551.51 54537 6.14 81.75 551.59 545.48 6.11 63.75 551.63 545.53 6.1 45.75 551.66 545.56 6.1 27.75 551.68 545.58 6.1 12.75 551.49 545.23 6.26

_ . 3_3 7 5. .._ .-_5 5._1 4_6_._ . 545.17 .. ____ ___6_.29

.+ - -

At Time = l_ 2.01679 second __

Eleystion ,4,, Tube Wall 2nd Huid JTemp L inch

._{.. TEMP ._ __ TEMP]Dif,ference_

,548.73 541.1I 7.62 3.375 ] . _ , _

21.75]_ $48 6 _ __ 541.,16 4 __ _7,._44 54;75 , j _ 548 46,,,4_ 541.15_ ,i 7.31 90.75__ j $48 38 p.541.11 7.27 130.25 3 548.27 ! 541.04 7.23

( _ 173;25 _,,548 l{ __,540_96___ 5 23 216.25 548.12 i 540,85 j 7.27 259.25 548.05 540.71 1 7.34 1

32[85 _i547.79 325.85 $4h.h9 540.27

[ 7.52 7.52

)

)

540.27 }

259 25 548.05 f 540.71 I 734 7 ,_, o ,,

_ I73 25 _. ._ 540.96 _ _

7.26

_ i3a255482:

99.75

_{_.548.22 _ _54107__. _. 7.24 7.22 l'

{ _ 54836_ ._ 541_.14 _

81.75 548.41 7.2 541.21 J .

, 63.75 548.45 541.25 7.2

! 45.75 548.48 541.27 7.21 r

t

! Page 6 l

7.,.

I Sheett 27.75 548.5 j 541.29 _,._ 7 21 12.75 548.34 541.01 _ 7;33

. 3 p 5 ,,,,__ 548 31 540.94 __ ,,,,_ 7.3 7

-.- ..- L. - _. - -_ . . _ - - . . . . . . . . . - . .

_6?ll.me" L2g056 _. wwnd i

. . . _ _ _ . , . __ _ . . _ . _ . _ __d _ .. _ . ._ _

Elevation Tube Wall 2nd fluid gernp _

.r inch TEMP TEMP lihfierence 3.375 548.73 541.11 7.62 21.75 548.6 541.16 7,44 54i [

~

7 31

_ 54.75 _._548 46 90 75 548.38 541.11 7.27 130.25 548.27 541.04 I 7.23 173.25, 548 19 _ , 540.96 _._ _ _ _ _7.23 216 25 _ _ 548.12 _ 540.85 ___7J7 259.25 548.05 540.71 7.34 325.8 7.52 547. 5 ] [ 5 0 727 b 325.85 547.79 540.27 7.52 259.25 548.05 540.71 7.34 216.25 548.14 540.85 1 7.29 173.25 548.22 540 96 7.26 130 25 548.31 ] 541.0 _7 7.24 99.75 548.36 541.14 7.22 81.75 548.41 541.21 7.2

_._ p{75,_ _ __ ,,548.4 5y .

541.25 7.2 45.75 548.4R 541.27 7.21 27.75 548.5 541.29 7.21 I2 7$. _ 548 34_ ._,,_54ipl_,_,_ __7.33 3.375 548.31 540 'A 7.37

_6[Tirne = ,,,,2.j 86p6,,,,, ,,__ sewnf _(,,

_.._..-____4.._____.).________

Elevation Tube Wall 2nd Fluid Temp.

s inch TEMP ,

TEMP Diffe_rence_

_3.375 _. 54 87_ 5.'3_854, i1.02 21.75 j 544.66 534.02 l 10.64

_.544.43 53,4_14 ._,10.29

_ 90.75_ _ 544.27 534.19 ! 10 08 130.25 _{ 544fy _

,,53,4._1_9] 9.87 173.25 _ 53_4 16 _ _ 9.73 216.25 543.89 _I__ 534 099 64 543.73 .

2593k $I3.57 533,97 , [6 325.8,5 . [ 543.17 _533 53g__9.64 325 85 543.13 533.53 9.6 259.25 ._I43 .,_ 533.97 - . _ . . . . _ . . . . _ 9.49

. 21_6;25 _ _p4hj6_, ._. $} q __.,_ ,, _ _ __, 9;4]

I?M - 543 64 5396__n __ 9 48 130 25_ ___. , _fi};]3 _ J 534.27 ,_Q..,,__ _

9_46 E2! _ _.#.._. 543 ]7_,,, ,. _ _;i34;}2,_[ _ ,_ 93

__ 81.75 543.82 534.36 9.46 63.75 543.85 534.37 9.48 45 75 541.87 534.37 9.5 Page 7

? Sheet 1

_ 2735._ _543 86g _53432__ ._9.54

. 3 75 _ _ 5431._l __533.53 . _ _.._ 9.97 3375 54143 533.41_ p _ 10 02

,_._.._.p.._.._._..[_. . . . h._.. _ _  !

At Time = _ 2.79162 ] _ second,,j __ _

l I Elevauon Tube Wall 2nd Huid iTemp __

r. inch TEMP TEMP l Difference 3.375 u-- 543.32. . -.532.19 I . - -

i 1.I3 54.75 542.72 I 10 49

-. 5_32 23 3 ..--

. _.13.0 25_ .

. _ _54 2._.24._. ._53 2_5.. _35_0 li 173 25 542.03 532.05 9.98 216.25 541.84 531.97 9.87 259.25 541.68 531.88 l 9.8 325.85 541.38 531.59 l 9.79 325.85 541.32 531.59 9573 259.25 541.5 531.88 i 9.62 216.25 5El.56 531$95T 9.59 173.25 541.62 532.05 9.57 130.25 541.69 532.12 9.57 99.75 541.73 532.17 9.56 81.75 54I.76 532 2 9.56 63.75 541.78 532.23 ' 9.55 45.75 541.8 532.25 9.55 27.75 541.81 532.29 9.52

.- .l235, .__ p%!? .!32 53_j _ _9 06 3.375_, 3 _ 541.59 531.95 9.64 I

At_Tme m 3.0308_ j_second _ _

_ __. . _ _ _ . _ _ . .f __..j_.._.

. peyapon_ Tup, W,all . _,2ndygd]Jerap___,__

IMD .4_]EMP _ _TEMPjDiffere_rme _

3.375

[ _ _542.25 _ _ 530 93 __. _ __ 1 1.32 21.75, _ Q 541.85 . 530.94 _ 10.91 54 75_ j _ _5_4143 _ 5}09,2 . . l0.51 90 75 541.13 5.W89 4 10.24 130 25 _. , 540.79 ] _530;8_4 j .. 9.95

_123.2p $40 51_ l. 530.78 ,_J 9.73 216 25 _ 540;29 _ 530.71_ _ _ 9.5 8

.259.25 - 540.1 9 46

5. 30 64 . a.. . .

. . . - - .. *: .b. . .. . . -- - .- . . ' b 32p5_J _53975. . . . .1 53tt i _ 9.33 259.25 539 84 [ 510,[ 6 _ _9 2 216.25 539.87 } 530 71 l 9.16 91:3

._173225_ p_539 95539 91_.L 530 78_}l ._.

._. i3a25 L 539 84 9 ii

_99g5 _ j 539ys . j_ 53pj8 L _ _9 i

_ 8!35 J. 540 01 j 535 9 L _ .9J1 9.11 63 75 . Q40 03_ i 530 92 l 45.75 ' 540 05 1 530 94 ' 9 11 Page 8

. I Sheet 1 D

. _7 . . . _. .W.b . .._. $k _ _ ~_ b' 12.75 540.03 g 530 91_4 __ _9.12

_ .3. 37_5.. . ._. _5 39. _99_. l ._530._85..

. . . . . _.__ . . 9 14 . . .

. _ s. _

._q.____._..g.._ . . ~ .

At Time = 3.43091 second 4.. ._._.._p_..._... _ ._

_ Elevation _ { Tube, Wall _J_2nd HuidjTemp_ _ _ _

r inch l TEMP TEMP (Difference 3.375 542 25 [ 53d93 .. . I l.32 21.75y _541.85_ . _ 530 94 _, { _1_0 91 54 75._ ] 541.43 530.92 j _ _ 10.51 90.75 541.13 .

530.89 10.24 130.25 540$7[ } 53d84 9.95

_l]3.2{__ , _, f 40.5_I_ _{,_530.78_ _ _ _ 9.73 216 25 540.29 530.71 l 9 58 l 259 25 _ ._ 540.I .

, _ 53_0 64 . _ _ 9 46 325 85 . 539 83 . g 53_0 4_ ,'_ _ _ 9 4}

325 8{ _ p39 7_5_ j_ 5303__. . _ ,, _ p;35 ,

259_25 ,; _,539 84 _j . 5_30 64 .( ._ _ _ 9.2 . l 216_25 539.87 j $3_0 7_I [ 9.16 173 25 539.91  ! 530.78 _9.13

n. .

130.2 { . 539.95 ] _ 530 8_4 _ __ _ _ 9.11 99.75 1 539.98 530.88 9.1 81.'75 540 01 530.9 9.11 63.75 - - -540.03 . ._530. 92 . , ..- - . . 91 1 .-

45.75._ ._

540.05 _530;94_ j 9.1 i 27.75_ j

_ 12.75 ._ _ j ,_540_0_3 _

540.08 _j 530 97

_ $30 91_f

]. 9. I 1 9.12 3.375 I 539.99 9.14

_ 530 85 _ _.

At Tiny = 4.05475 second

.__ . _ . 1_ _ _ _ . . . _.

Elevauon 1 Tube Wall j 2nd Huid Temp _ _

r, inch i TEMP ' TEMP Difference 3.375 539.22 _ 5263 9 [ 12} 3 526.26 12.41 21.75_ _ [ 538.67 I

54.75 538.07 526.22 11.85 90.75 537.57

_526h ~ 11 [7 130.25 I 537 526.16 l 10.84 173 25 5.56.51 _

526Ill _1_0) 526 05 10 03 216 25 ] _ 536 08 259 25 .j 535 71 _ ,. 525 97 ! 9.74 325.85  ! 535 32 ! 525.73 i 9.59 325 85 ,

535 1 525 73 ,

9.37 259.25 j 535 06 . [ 525 97 ,j. 9 04 2l_6 25 _j. 534 98 { 526 05_ J _ 8.93 8 81 173.25 [ 534 92 .] 52611_ .[

13323 4 53t89 4 526 is_ ! __ _833 99 75 L 534 88 j.._526 2i_ 4 _.8 67 8I 75

.L 534 88 L 526.23 .J.__ _8 65 63 75 ; . Su 8t.. L_526.25 j _ _8.63 45.75 l 514 88 1 526.27 '

8 61 Page 9

p.,

e l

.. Sheet 1 )

_ _.~2 L?5__._ ., 534.89 i - . -.526.3- -- .. - - .- 58 9

% w% '

l

"'" . y

' "'u2 5"*/ had I 534 9 526.22 i ~ s6s 1 i

Page 10

O Table 13-1 Stress Intensity Ratio as a Function of Bi-Axiality Model D Tube Support Plate SINT Ratio Biaxiality Ratio Pitch Diagonal 1.00 3.63 3.79

-0.80 3.65 4.19

-0.60 3.74 4.69

-0.40 3.95 5.33

-0.20 4.36 6.19 0.00 5.12 7.41 0.20 5.22 7.40 0.40 5.51 7.41 0.60 6.00 7.42 0.80 6.67 7.42 1.00 7.39 7.43 8.0 0 0 0 7.0 -

6.0 - i

.E, 5.0 -

4.0 -

g -.- g - E 3.0 '

I 1.5 -1.0 -0.5 0.0 0.5 1.0 1.5 l Di Axiality a Pitch .<> Diagmal DISK 225A - SLBPLT\TBLQ01 - 06/29/95

Table 20-1 Summary of Expanded Tube Field Operating Experience Tutm ao Roskfuel humber of Min. Time for Operat6ng Time Factor for Men. Est.

Type of Expension Dia. Stress 1600 MA Circ. Cracking Plant Temp Onset of C4rc. Crack Time to Tubes in Field Experience Q=35 Q=50 Crack [8]

(in.) (in.) (ksi) Service (yrs) (*F) (yrs.)

Mecharucal 0.875 0.016 40 OD [3] >20.000 8 OD V-2 611 7.2 16.6 57.6 (Hard Rot) 40-60 ID [3]

0.750 >100,000 3. OD B-2 620 9.2 23.7 27.6 4.ID S 619 8.9 22.7 35.6

~6 Byron-1, 619 8.9 22.7 ' 53.4 Braidwood-1

, WEXTEX 0.875(1] 0.016 30 [4] >100,000 9, ID/OD A-2 611 7.2 16.6 64.8 8.ID W-2 610 6.9 15.9 55 2 -

Hydra:Aic Expanston 0.750 0.016 20 OD [3] 19,400 none cracked AC-2 626 10.8 30.1 >64.8 20-30 tD [3] 6 yrs. operation 0.688 17,160 11,OD FC 617 8.5 21.1 93.5 HEJ Sleeve 0.875 0.027 [2] 35-401D [9] -9000 4,00[7] A, 599 5.2 10.5 20[ ~

~20 OD [9]

0.750 >1750 1,lD E-4 620 92 23.7 92 TSP Hydraube Exp. 0.750 0.014-0.040 ID>OD [10] >1000 none cracked [6] Model D4 538-558 - -

(cold leg) 0.055 [5] ~25 none cracked Model D4 538-558 0.875 0.025 -15 none cracked P-1 542 Tube Expanson 0.750 0.057- 30-6010 [3] 544 0.090 20-30 OD [3]

[1] None en smaller tubes [2] After Hardroll [3] Polythionse Acid Stress indexing Tests

[4] Magnesium Chloride Stress Indexing Tests

[5] Bulges outsde TSP left in service if less than 30 rnits larger than expanson within the TSP Number in service is an estimate.

[6] Expansion performed prior to plant startup in domestic plants. Longest operating time is 11 calendar years.

[7] ID in parent tubes found by dest.W examination.

[8] For similarfy stressed expansion operating at 544 *F

[9] Estimated from doped steam tests

[10] Magensium Chloride tests. Time to cracking shorter on ID than OD and both independent of expansion up to 0.041* Ad tested.

Resu!ts judged similar to Polythionic Acid stress indexing tests.

4 RAl20TBLXLS 6/29/95

_.m. ._ _ w