ML20129E274

Nonproprietary F* & Elevated F* Tube Plugging Criteria for Tubes W/Degradation in Tubesheet Region of Prairie Island Sgs

ML20129E274
Person / Time
Site:	Prairie Island
Issue date:	09/30/1996
From:	WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML19355E409	List: ML20129E218 ML20129E245 ML20129E259 ML20129E274 ML20129E283
References
WCAP-14226, WCAP-14226-R01, WCAP-14226-R1, NUDOCS 9610010029
Download: ML20129E274 (41)

Text

__

-z,

-, e m w ;;m., _,,

%W;i)i'l Q

-v Ey?cg/jg;;

Off V:

,ty t

1,y.p 3.t d.i^..ip w ij;' d 4 Wlr?h!.((b O.] %

.t.

'7 'g; rp. -A: ;, n'yi Westinghouse N'on-Proprietary Class 3 ny.W i M. # q me,

-l, :

p::'g%).v g c

b

?q q

i,UsX

..s

,y' w,fr

• r, '

?+%' n;q M

'Y

WCAP-14226

- ((]-((l

$f $

Revision 1

> c

'StN, i i,, f,{

. '9n, $$

ww\\\\;j!r,:43:

ct M

'gjp:, pn.ing s...

s.

s *k, ',t,n;p?'A f

.#l), N O

-,n yt,Q,jp

.ek F* and Elevated F* Tube Plugging Criteria for jg pg[$Q..

Tubes with Degradation in the Tubesheet g

.,m? :

Region of the Prairie Island Units 1 and 2 ~

~

Tl Steam Generators' k

• 3. 5, '4 1,

.r y..

m.t l

. m

_ _ ? ('i,4 s

5 l ph. s y%.

[3 a -;et}Vj.;

u.,h,n,, ; )

a

,, v.m c

c.

, :n

+

' '., N-1

,w

'. tif ;p, 4

r

~ ib 6l: * '

1. . n',%

r. :. x 4

. &

N

/.

Westinghouse Energy Systems W

g-emus 9610010029 960924 PDR ADOCK 05000282 P

, PDR

~

..c..-.

i a-WESTINGHOUSE NON-PROPRETARY CLASS 3 o

~

WCAP-14226 Revision 1 F* AND ELEVATED F* TUBE PLUGGING CRITERIA FOR TUBES WITH DEGRADATION IN THE TUBESHEET REGION OF THE PRAIRIE ISLAND UNITS 1 AND 2 STEAM GENERATORS September 1996 Work Performed Under Contract:

Purchase Order No. PK4379SQ Westinghouse General Order No. MKD 77507 l

WESTINGHOUSE ELECTRIC CORPORATION NUCLEAR SERVICES DIVISION i

P.O. BOX 158 MADISON, PENNSYLVANIA 15663 C1996 WESTINGHOUSE ELECTRIC CORPORATION ALL RIGHTS RESERVED I

1 4

ABSTRACT Evaluations have been performed to develop repair criteria for tubes with degradation in the roll expansion and tubesheet region of the Prairie Island Units 1 and 2 steam generator tubes.

A criterion based on maintaining structural adequacy of the tubes to resist tube pullout forces, called F*, has been developed. This criterion specifies a distance below the bottom of the roll transition, designated F*, in which no degradation is permissible; below the F* length, all types of degradation may be left in service independent of the number or nature of eddy current indications. It is determined that the required F* length is 1.07 inches plus an allowance for NDE uncertainty.

For tubes with degradation in the intended F* region or above the original roll transition region, an additional criterion has been developed based on determining the length of additional roll expansion of sound tube needed to achieve adequate pullout and leakage resistance. For tubes with degradation in or above the factory roll transition, an additional tube roll may be applied adjacent to or above the factory roll expansion. For simplicity, the length of sound, factory rolled tube and additional tube roll immediately adjacent to the factory roll expansion is also referred to as F*

For tubes in which the additional roll expansion is between the factory roll and the neutral bending axis of the tubesheet (10.665 inches above the bottom of tubesheet cladding, or 10.885 inches above the tube end), the F*

length also applies. For tubes in which the top of the additional roll expansion is above the vertical mid-thickness of the tubesheet, the necessary length of undegraded additional roll expansion is referred to as elevated F*, or EF* The required EF* length is a function of elevation and is larger than the F* length, since upward tubesheet bending during normal operation produces relaxation of the tube-to-tubesheet radial contact pressure above the neutral bending axis of the tubesheet. Values of EF* length are calculated, excluding NDE l

uncertainty in elevation, for several elevations ranging from the top of the tubesheet to 6 inches below the top of the tul:esheet.

The evaluations demonstrate that application of the F* and EF* criteria for indications of tube degradation within the roll expansion affords a level of plant protection commensurate with that provided by Regulatory Guide 1.121 for degradation located outside of the tubesheet region.

4 i

~

p TABLE OF CONTENTS Section 19mG P_aan ~

1.0 INTRODUCTION

1-1

' l.1

Background

1-2 2.0 DEVELOPMENT OF F* CRITERION 2-1

~

2.1 Determination of Elastic Preload Between the Tube and Tubesheet 2-2 2.1.1 Radial Preload Test Configuration Description 2-2 2.1.2 Preload Test Results Discussion and Analysis 2-3 2.1.3 Residual Radial Preload During Plant Operation 2-4 2.14 Increase in Radial Preload Due to Thermal Expansion Tightening 2-5 2.1.5 Increase in Radial Preload during N.O. and FLB Due to 2-5 Differential Pressure 2.1.6 Change in Radial Preload due to Tubesheet Bow 2-6 2.1.7 Net Preload in Roll Transition Region for N.O. and 2-6 FLB Conditions 2.2 Determination of Required Engagement Distance 2-6 2.2.1 Applied Loads 2-7 2.2.2 Coefficient of Friction at Tube-to-Tubesheet Interface 2-7 2.2.3 End Effects 2-7 2.2.4 Calculation of Engagement Distance Required, F*

2-8 2.3 Limitation of Primary to Secondary Leakage 2-9 2.3.1 Operating Condition Leak Considerations 2-9 2.3.2 Postulated Accident Condition Leak Considerations 2-10 2.3.3 Operating Plant Leakage Experience for Within Tubesheet 2-11 Tube Cracks 2.4 Tube Integrity Under Postulated Limiting Conditions 2-11 2.5 Chemistry Considerations 2-12 2.5.1 Tubesheet Corrosion Testing 2-12 4

2.5.2 Tubesheet Corrosion Discussion 2-13 2.6 Summary of F* Evaluation 2-14 l'

3.0 DEVELOPMENT OF RETROFIT F* AND EF* CRITERIA 1 j

3.1 Details of Determination of EF* Based on Structural Requirements 3-2 3.2 Limitation of Primary to Secondary Leakage 3-3 l

3,3 Summary of EF* Evaluation 3-3

4.0 REFERENCES

4-1 i

i J

~. - -

- -. ~. -. -

j LIST OF TABLES NE Iidt Eagt 2-1 SG Tube Roll Preload Test - Test Data 2-15 2-2 SG Tube Roll Preload Test - Stress Analysis Results 2-18 i

2-3 Calculation of F* Length for T/TS Hardroll Interface - Prairie Island 2-19 Units 1 and 2 2-20

)

2-4 Preload Analysis Summary 3-1 Evaluation of Required Engagement Length, EF*, for Prairie Island -

3-4 Units 1 and 2

4 LIST OF FIGURES Fit No.

1 1113 Pm 1-1 Configuration for Tubesheet Factory Roll Region F* Altemate Repair / Plugging 1-4 Criterion for Partial Depth Roll-Expanded Steam Generator Tubes (Single Band of Degradation) 1-2 Configuration for Tubesheet Above-Factory-Roll Region F* Alternate 1-5 Repair / Plugging Criterion for Partial Depth Roll-Expanded Stearn Generator Tubes 1-3 Configuration for Tubesheet Region EF* Alternate Repair / Plugging 1-6 Criterion for Partial Depth Roll-Expanded Prairie Island Steam Generator Tubes 2-1 "End-Effect" Regions in a Tube Postulated to be Severed at the Bottom 2-22 of F* Length i

iV

NOMENCLATURE (Terms are listed in alphabetical order)

Linear indication inclination angle from the tube axis, a.k.a., crack angle BRT Bottom of roll transition dpm Drops per minute DRE Degraded roll expansion ECI Eddy current indication ECT Eddy current test FLB Feedline break ID Inside diameter i,

Indication length LOCA Loss of Coolant Accident LTL Lower tolerance limit MA Mill annealed MBD Multiple band degradation N

Number of sound roll expansion portions NBA Neutral bending axis NDD No Detectible Degradation NDE Non-destructive examination N.O.

Normal Operation OD Outside diameter PCWG Power Capability Working Group PLRL Pullout load reaction length I

RE Roll expansion RPC Rotating pancake coil RT Roll transition RTL Resistance to leakage SBD Single band degradation S/G Steam generator i

SLB Steam line break s,

T/TS interfacial radial contact pressure SRE Sound roll expansion T

Torque TS Tubesheet TTS Top of tubesheet T/TS-Tube to tubesheet X

Distance from top ofindication to BRT V

=

i j.*

j

1.0 INTRODUCTION

i 1

This report documents the development of criteria for repairing panial depth hardroll expanded steam generator tubes for degradation in the expanded region of the tube within the tubesheet and for degradation above the factory roll transition but within the tubesheet.

4 j

Existing Prairie Island Units I and 2 Technical Specification tube repairing / plugging criteria i

apply throughout the tube length, but do not take into account the reinforcing effect of the tubesheet on the extemal surface of the tube. Two repair criteria, called F* and EF*, are developed for tubes with degradation within the tubesheet region.

For tubes with degradation in the factory roll transition region, a criterion based on maintaining structural adequacy of the tubes to resist tube pullout forces, called F*, has been developed. This criterion specifies a distance below the bottom of the roll transition, designated F*, in which no degradation is permissible; below the F* length, all types of

]

degradation may be left in service independent of the number or nature of eJ3y current i

indications. The F* length is depicted in Figure 1-1.

5-An altemate repair / plugging criterion has been developed for tubes which exhibit degradation in or above the intended F* distance, i.e., above the bottom of the factory roll transition. This l

criterion uses a field-applied, additional roll expansion adjacent to or above the factory roll, as depicted in Figure 1-2. With an appropriate qualification program (not included in the scope

~

of this analysis), the F* length also applies in the additional roll expansion, provided the tdditional roll expansion is below the neutral bending axis of the tubesheet. The applicability of the F* length in this region is dependent upon the results of a separate qualification or j

evaluation showing the structural equivalency of the retrofit roll expansion to the factory roll.

This qualification program shall also demonstrate that the additional roll expansion of F*

length exhibits negligible leakage.

For tubes which exhibit degradation between the mid-thickness of the tubesheet and the top of the tubesheet (TTS), a criterion referred to a3 elevated F*, or EF*, has been developed. It rJso uses a field applied, additional roll expansion, depicted in Figure 1-3.

The applicability of the EF* length is dependent upon the results of a separate qualification or evaluation j

showing the structural equivalency and leakage resistance of the elevated, retrofit roll expansion relative to the factory ro'.l.

Development of the F* and EF* criteria is summarized in Sections 2 and 3, respectively. In i

Section 2.0, an evaluation is performed to determine the length of hardroll engagement, F*,

required to resist tube pullout forces during normal operation, test, upset and faulted conditions. The evaluation uses the results of previously conducted analyses and tests aimed et quantifying the residual radial preload of Westinghouse 51 Series steam generator tubes i

hardrolled into the tubesheet holes. It is postulated that the radial preload would be sufficient to significantly restrict leakage during normal operation and faulted conditions. The necessary isngth of undegraded tubing, called F*, required to resist tube pullout forces is calculated. On this basis, tubes in the Prairie Island Units 1 and 2 steam generators with no degradation i.

within the F* distance, measured from the bottom of the roll transition, are sufficient for continued plant operation, regardless of the extent or nature of tube degradation below F*,

I-1

4 1 -

In Section 3.0, an evaluation is performed to determine the length of additional hardroll expansion required to resist tube pullout forces during normal operation, test, upset and j

faulted conditions. For tubes with degradation adjacent to the original factory hard roll or i

below the mid-thickness of the tubesheet, it is shown that the F* length calculated for the I

factory hard roll region also applies for the additional roll expansion region adjacent to or above the factory roll. However, for degradation between the mid-thickness of the tubesheet and the TTS, the necessary length of elevated additional roll expansion, called EF*, required to resist tube pullout forces must be calculated. The EF* distance is larger than the F*

l distance. This is because F* and EF* are directly related to radial preload, a.k.a. interference j

fit contact pressure. The EF* criterion is applied above the neutral bending axis of the tubesheet; tubesheet upward bending effects cause a reduction in the radial preload during the

]

bounding conditions, normal operation (N.O.) and feedline break (FLB). In the case of F*,

{

applied within the factory roll or above and adjacent to it, the tubesheet upward bending l

effects are beneficial or have no effect on the contact pressure. The evaluation ofinterference j

fit contact pressure for EF* is also provided in Section 3.0.

These evaluations demonstrate that the F* and EF* criteria for tube degradation within the tubesheet afford a level of plant protection commensurate with that provided by Reg. Guide s

I (RG) 1.121 for degradation located outside of the tubesheet region.

I

1.1 Background

l Existing plant Technical Specification tube repair / plugging criteria which have been applied throughout the tube length do not take into account the reinforcing effect of the tubesheet (TS) on the external surface of the expanded portion of the tube. The presence of the TS will constrain the tube and will complement tube integrity in that region by essentially precluding tube deformation beyond the expanded outside diameter. The resistance to both tube rupture and tube collapse is significantly strengthened by the TS. In addition, the proximity of the l

TS significantly affects the leak behavior of throughwall tube cracks in this region. Based on j

these considerations, the establishment of alternate plugging criteria specific to the roll region of the tubes is justified.

J The roll expanded length of the tube can be considered to consist of two zones, the roll transition (RT) region and the roll region. The roll transition region is defined as that portion 4

of the tube where the roll expanded length transitions to the unexpanded length. The RT is approximately [

]* inch above the tube end at the bottom of the TS (i.e, above the initial tack roll applied during tube bundle assembly). The roll region is also referred to as the roll expansion (RE) or hardroll. Refer to Figure 1-1.

l Taking credit for the reinforcing effect of the tubesheet and the radial contact load between the expanded region of the tube and the tubesheet, the F*, EF* and L* plugging criteria are developed. The F* criterion permits operation with any amount of tube degradation below a 1-2 4

s calculated distance, F*, below the bottom of the factory roll transition (BRT). No degradation is permissible within the F* distance. The F* distance is shown to provide sufficient frictional force between the expanded tube and tubesheet to resist pullout due to normal operation and postulated accident conditions. The use of the F* criterion does not require any assessment of tube degradation other than elevation, and any type of degradation below the F* distance is acceptable.

Eddy current indications (ECI's) in the top portion of the roll expansion (within the intended F* distance) and above~ the factory BRT but below the neutral bending axis (NBA) may also be addressed by F*

Above the factory tube expansion, F* criteria will apply based on achievement of an additional roll expansion of sufficient engagement such that pullout forces that develop during normal or accident operating conditions would be successfully resisted by the elastic preload between the tube and the tubesheet. The additional roll expansion will serve as a continuance of the original roll expansion.

Tube degradation not meeting the elevation requirements of F* can also be addressed by the alternate repair criteria defined as EF*.

The EF* criterion provides for sufficient engagement of the additional, elevated tube-to-tubesheet hardroll, such that pullout forces that develop during normal or accident operating conditions would be successfully resisted by the elastic preload between the tube and the tubesheet. The elevated additional roll expansion is performed above the mid-plane of the tubesheet. The necessary engagement length applicable to the Prairie Island steam generators was determined based on mechanical interference fit (MIF) analysis.

l 1-3

i i

i 4

s a

Figure 1-1.

Configuration for Tubesheet Factory Roll Region F* Alternate Repair / Plugging Criterion for Partial Depth Roll-Expanded Steam Generator Tubes (Single Band of Degradation) 1-4 t

e

-m i

1

^

l l

1

.,c,.

l Figure 1-2.

Configuration for Tubesheet Above-Factory-Roll Region F* Alternate Repair / Plugging Criterion for Partial Depth Roll-Expanded Steam Generator Tubes i

1-5

W t

l l

Figure 1-3.

Configuration for Tubesheet Region EF* Alternate Repair / Plugging Criterion for Partial Depth Roll Expanded Steam Generator Tubes l

1-6

s 1

2.0 DEVELOPMENT OF Fa CRITERION 4:

I The F* tube plugging criterion is based on a semi-empirical method of quantifying the axial loadbearing capability of the rolled joint, resulting from the radial contact preload pressure i

and the associated friction between the tube and TS. The presence of the tube-to-tubesheet

}

(T/TS) radial pressure, s,, which consists of the as-manufactured pressure as changed by operating loads and temperatures, also causes significant resistance to the leakage of

. primary-to-secondary and secondary-to-primary water It has been determined by previous

~

j tests, described in Appendix A of Reference 7, that sound roll expansions of F* length are l

essentially leak tight. The use of the F* criterion obviates the necessity of determining the ECI depth, number, inclination, length and circumferential spacing. Only the distance from the uppermost part of the ECI to the BRT needs to the determined. In short, the nature and I-extent of tube degradation need not be determined.. Refer to Figure 1-1.

Tube rupture in the conventional sense, as characterized by an axially ori ented " fishmouth" opening in the side of the tube, is not possible within the tube /tubesheet roll expansion (RE).

4 The reason for this is that the tubesheet material prevents the wall of the tube from expanding I

outward in response to the intemal pressure forces. The forces which would normally act to cause crack extension are transmitted into the walls of the tubesheet, the same as for a non-1 i

degraded tube, instead of acting on the tube material. Thus, axially oriented linear i

indications, e.g., cracks, cannot lead to tube failure within the RE and may be considered on the basis of leakage effects only.

4 Likewise, a circumferentially oriented tube rupture is resisted because the tube is not free to deform in bending within the roll expansion. When degradation has occurred such that the remaining tube cross sectional area does not present a uniform resistance to axial loading, l

bending stresses are developed which may significantly accelerate failure. When bending forces are resisted by lateral support loads, provided by the tubesheet, the acceleration mechanism is mitigated and the tube separation mode is similar to that which would occur in a simple tensile test. Such a separation mode, however, requires the application of significantly higher loads than for the unsupported case.

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

the location determined acceptable for the postulated condition.

When the tubes have been hardrolled into the tubesheet, any axial loads developed by

~ pressure and/or mechanical forces acting on the tubes are resisted by friction forces developed

. by the elastic preload that exists between the tube and the tubesheet. For some specific length of engagement of the hardroll, no significant axial forces will be transmitted farther l

along the tube, and that length of tubing, i.e., F*, will be sufficient to anchor the tube in the tubesheet. In order to determine the value of F* for application in 51 Series steam generators,

. a testing program was conducted to measure the elastic preload of the tubes in the tubesheet.

4 v

2-1 i

+.,..

s The presence of the elastic preload also presents a significant resistance to flow of primary to secondary or secondary to primary water for degradation which has progressed fully through the thickness of the tube. In effect, no leakage would be expected if a sufficient length of hardroll is present. This has been demonstrated in steam generator sleeve to tube joints made j

by the Westinghouse hybrid expansion joint process.

1 1

1 2.1 Determination of Elastic Preload Between the Tube and Tubesheet i

Tubes were installed in the Prairie Island steam generators using a hardrolling process which expands the tube to bring the outside surface into intimate contact with the tubesheet hole.

The roll process and roll torque are specified to result in a metal to metal interference fit j

between the tube and the tubesheet.

4 A test program was conducted by Westinghouse to quantify the degree ofinterference fit

' between the tube and the tubesheet provided by the partial depth hardrolling operation. The

]

' data generated in these tests have been analyzed to determine the length of hardroll required to preclude axial tube forces from being transmitted farther along the tube, i.e., to establish j

the F* criterion. The amount of interference was determined by installing tube specimens in collars specifically designed to simulate the tubesheet radial stiffness. A hardroll process representative of that used during steam generator manufacture was used in order to obtain l

specimens which would exhibit installed preload characteristics like the tubes in the tubesheet.

j Once the hardrolling was completed, the test collars were removed from the tube specimens j

and the springback of the tube was measured. The amount of springback was used in an analysis to determine the magnitude of the interference fit, which is representative of the

{

residual tube to tubesheet radial load in Westinghouse 51 Series steam generators.

i t

j 2.1.1 Radial Preload Test Configuration Description i

The test program was designed to simulate the interface of a tube to tubesheet partial depth hardroll for a 51 Series steam generator. The test configuration consisted of six cylindrical collars, approximately [ ]* inches in length, [

]"' inches in outside diameter (OD), and

}

[

]* inch in inside diameter (ID). A mill annealed, Alloy 600 (ASME SB-163), tubing j

specimen, approximately [ ]* inches long with a nominal [

]"' outside diameter before l

rolling, was hard rolled into each collar using a process which simulated actual tube B

installation conditions.

1 The design of the collars was based on the results of performing finite element analysis of a section of the steam generator tubesheet to determine radial stiffness and flexibility. The i

inside diameter of the collar was chosen to match the size of holes drilled in the tubesheet.

The outside diameter was selected to provide the same radial stiffness as the tubesheet.

I The collars were fabricated from AISI 1018 carbon steel similar in mechanical properties to l

the actual tubesheet material. The collar assembly was clamped in a vise during the rolling process and for the post roll measurements of the tube ID. Following the taking of all post l

2-2 L

t i.

roll measurements, the collars were saw cut to within a small distance from the tube wall.

The collars were then split for removal from the tube and tube ID and OD measurements 4

were repeated.

Two end 'ooundary conditions were imposed on the tube specimen during rolling. The end was testrained from axial motion in order to perform a tack roll at the bottom end, and was.

allowed to expand freely during the final roll.

I 2.1.2 Preload Test Results: Discussion and Analysis All measurements taken during the test program are tabulated in Table 2-1. The data recorded were employed to determine the interfacial conditions of the tubes and collars.

These consisted of the ID and OD of the tubes before and after rolling and after removal from the collars, as well as the inside and outside diameters of each collar before and after tube rolling Two orthogonal measurements were taken at six axial locations within the collars and tubes. Additional data ofinterest were calculated from these specific dimensions. The calculated dimensions included wall thickness, change in wall thickness for both rolling and removal of the tubes from the collars, and percent of spring back.

Using the measured and calculated physical dimensions, an analysis of the tube deflections was performed to determine the amount of preload radial stress present following the hardrolling. The analysis consisted of application of conventional thick tube equations to account for variation of structural parameters through the wall thickness. However, traditional application of cylinder ' analysis considers the tube to be in a state of plane stress. For these tests, the results implied that the tubes were in a state of plane strain elastically. This is in agreement with historical findings that theoretical values for radial residual preload are below those actually measured, and that axial frictional stress between the tube and the tubesheet increases the residual pressure. In a plane stress analysis such stress is taken to be zero (References 2 and 3). Based on this information, the classical equations relating tube deformation and stress to applied pressure were modified to reflect plane strain assumptions.

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

u = C

• r + C: / r (1) i

where, u = radial deflection r = radial position within the tube wall, and the constants, C and C are found from the boundary conditions to be functions of the i

2 elastic modulus of the material, Poisson's ratio for the material, the inside and outside radii, i

and the applied intemal and external pressures. The difference between an analysis assuming plane stress and one assuming plane strain is manifested only in a change in the constant C -

2 The first constant is the same for both conditions. For materials having a Poisson's ratio of 0.3, the following relation holds for the second constant:

2-3

i

~

C (Plane Strain) = 0.862

• C (Plane Stress)

(2) 2 2

The effect on the calculated residual pressure is that plane strain results are higher than plane stress results by slightly less than 10 percent. Comparing this effect with the results reported 1

in Reference 2 indicated that better agreement with test values is achieved. It is to be noted J

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

i By substituting the expressions for the constants into Equation (1), the deflection at any radial j

location within the tube wall as a function of the intemal and external pressure (radial stress i

at the ID and OD) is found. This expression was differentiated to obtain flexibility values for

}

the tube deflection at the ID and OD respectively, e.g., dU/dP,is the ratio of the radial deflection at the ID due to an OD pressure. Thus, dU/dP, was used to find the interface pressure and radial stress between the tube and the tubesheet as:

b S,, = - P, = - (ID Radial Springback) / (dU/dP )

(3) i j

The calculated radial residual stress for each specimen at each location is tabulated in

]

Table 2-2. The mean residual radial stress and the standard deviation were found to be j

[

]"' psi and [

]"' psi, respectively. In order to determine a value to be used in the l

analysis, a tolerance factor for [

]"' percent confidence to contain [

]"' percent of the population was calculated, considering the [

]"' useable data points, to be [

]"'.

Thus, l

a[

]"' lower tolerance limit (LTL) for the radial residual preload at room temperature is

[

]"' psi.

1 2.1.3 Residual Radial Preload During Plant Operation i

During plant operation, the amount of preload will change depending on the pressure and temperature conditions experienced by the tube. The room temperature preload stresses, i.e.,

radial, circumferential and axial, are such that the material is nearly in the yield state if a l

comparison is made to ASME Code (Reference 4) minimum material properties. Since the coefficient of thermal expansion of the tube is greater than that of the tubesheet, heatup of the plant will result in an increase in the preload and could result in some yielding of the tube.

In addition, the yield strength of the tube material decreases with temperature. Both of these effects may result in the preload being reduced upon return to ambient temperature conditions, i.e., the cold condition. However, as documented in Reference 5 for a similar investigation, tube pullout tests which were preceded by a very high thermal relaxation soak showed the l

analysis to be conservative.

The plant operating pressure influences the preload directly based on the application of the pressure load to the ID of the tube, thus increasing the amount ofinterface loading. The pressure also acts indirectly to increase the amount ofinterface loading by causing the tubesheet to bow upward, i.e., placing the roll expansions near the bottom of the tubesheet in compression for normal operating and feedline break (FLB) conditions (FLB results in a higher primary-to-secondary AP than steamline break, hence FLB is used to bound the FLB and SLB cases for this analysis). For the loss of coolant accident (LOCA) event, the tubesheet bows in the opposite direction, producing dilation of the tubesheet holes and 2-4

~

reducing the amount of tube to tubesheet preload. Each of these effects may be quantitatively treated.

?

The maximum amount ofincrease in preload due to tubesheet bow for primary-to-secondary 4

pressure differential will occur at the bottom, central part of the tubesheet. Since F* is measured from the bottom of the hardroll transition (BRT) and leakage is to be restricted by the F* region of the tube, the potential for the tube section within the F* region to experience a net tightening or loosening during operation is evaluated._ However, the central location case is not the most stringent case for normal operation and FLB; rather, the most stringent case for normal operation and FLB involves a peripheral tube, which experiences little or no j

increase in radial preload due to tubesheet bending.

J The effects of the three identified mechanisms affecting the preload are considered in the following sections. In order to obtain a value of radial preload which accounts for the current operating conditions as well as future variations in normal operating temperatures and i

pressures, the radial preloads due to the three identified mechanisms have been considered for several levels of steam generator plugging. Operating parameters from the Power Capability Working Group (PCWG) for plugging levels of 0%,3.67% (representing the current operating prameters at Prairie Island), and 15% have been used.

2.1.4 Increase in Radial Preload Due to Thermal Expansior2 Tightening i

For conservatism in determining the total residual preload for normal operating conditions, tightening of the tube /tubesheet joint due to differential thermal expansion is minimized by applying the SG outlet temperature to the tubing. From the cases identified in Section 2.1.3 j

and Table 2-3, this corresponds to a cold leg temperature of 528*F. The mean coefficient of thermal expansion for the Alloy 600 tubing between ambient conditions and 528*F is approximately 7.74*104 in/in/"F. That for the steam generator tubesheet is 7.30*104 in/in/ F.

These values were reconciled as conservative with respect to the 1965 ASME Miler and Pressure Vessel Code, which was the code of construction for the Prairie Island Units 1 and 2 SGs. Thus, there is a net difference of 0.44*10 in/in/*F between the expansion properties of

)

4 the two materials. Considering a temperature difference of (528 - 70) = 458'F between ambient and operating conditions, the increase in preload between the tube (t) and the tubesheet (ts) was calculated as:

S,r = (0.44E-6)*(458)*(Collar ID) / 2 / ((dU/dP), - (dU,/dP ),)

(4)

The results indicate that the increase in preload radial stress due to thermal expansion is

[

]* psi. Note that this value applies for both normal operating and faulted conditions.

2.1.5 Increase in Radial Preload during N.O. and FLB Due to Differential Pressure The normal operating (N.O.) differential pressure from the primary to secondary side of the steam generator during the most limiting PCWG condition is 1593 psi. The intemal pressure acting on the wall of the tube will result in an increase of the radial preload on the order of the pressure value. The increase was found as:

2-5

o.

l S,,, = - P, ? - P, (dU/dP,) / ( (dU/dP,),, - (dU/dP,))

(5) l In actuality, the increase in preload will be more dependent on the internal pressure of the tube since water at secondary side pressure would not be expected between the tube and the tubesheet. However, the primary to secondary AP is used for conservatism.

The increase in radial contact pressure due to differential pressure was evaluated for both normal operating (AP = 1593 psi) and faulted (AP = 2650 psi) conditions. The results indicate that the increase in preload radial stress is [

]"' psi for normal operating conditions and

[

]* psi for faulted (FLB) conditions.

2.1.6 Change in Radial Preload due to Tubesheet Bow An analysis of the 51 Series tubesheet was performed to evaluate the change in preload stress j

that would occur as a result of tubesheet bow for interior tubes. The analysis was based on

. performing finite element analysis of the tubesheet and SG shell using equivalent perforated plate properties for the tubesheet (Reference 3). Boundary conditions from the results were then applied to a smaller, but more detailed model, in order to obtain results for the tubesheet holes. Basically the deflection of the tubesheet was used to find the stresses active on the bottom surface and then the presence of the holes was accounted for. For the location where i

the increase of preload is a maximum, the radial preload stress would be increased by

[

]"' psi during normal operation and [

]"' psi during faulted (FLB) conditions.

However, the interior tubes are not the limiting case for primary-to-secondary pressure differential. The limiting case involves peripheral tubes where tubesheet bowing has a negligible effect on tube-to-tubesheet preload. Therefore, the N.O. and FLB analyses address only tubes in the peripheral region of the tubesheet. During LOCA, the differential operating pressure is from secondary to primary. Thus, the radial preload will decrease by [

] psi as the tubesheet bows downward. However, the action of the differential pressure is such that the tube is pushed toward the tube-to-tubesheet well. This case is of no consequence to the determination of F*

2.1.7 Net Preload in Roll Transition Region for N.O. and FLB Conditions Combining the room temperature hardroll preload with the thermal and pressure effects results in a net operating preload of [

]"' psi during normal operation and

[

]"' psi for faulted conditions. In addition to restraining the tube in the tubesheet, this preload should effectively retard leakage from indications in the tubesheet region of the tubes.

2.2 Determination of Required Engagement Distance The calculation of the value of F* recommended for application to the Prairie Island Units 1 i

and 2 steam generators is based on determining the length of hardroll necessary to offset the applied loads during the maximum normal operating conditions or faulted conditions, whichever provides the largest value. Thus, the applied loads are balanced by the load 2-6

o i

i i

carrying ability of the hardrolled tube for both of the above conditions. In performing the j

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

2.2.1 Applied Loads I

j The applied loads to the tubes which could result in pullout from the tubesheet during all normal and postulated accident conditions are predominantly axial and due to the internal to cxternal pressure differences. For a tube which has not been degraded, the axial pressure load 4

is given by the product of the pressure with the intemal cross-sectional area. However, for a tube with intemal degradation, e.g., cracks oriented at an angle to the axis of the tube, the intemal pressure may also act on the flanks of the degradation. Thus, for a tube which is.

i conservatively postulated to be severed at some location within the tubesheet, the total force l

acting to remove the tube from the tubesheet is given by the product of the pressure and the cross-sectional area of the tubesheet hole. The force resulting from the pressure and internal area acts to pull the tube from the tubesheet and the force acting on the end of the tube tends to push the tube from the tubesheet. For this analysis, the tubesheet hole diameter has been l

used to determine the magnitude of the pressure forces acting on the tube. The forces acting to remove the tube from the tubesheet are [

]'" pounds and [

]'" pounds respectively i

for normal operating and faulted conditions. Any other forces such as fluid drag forces in the

/

U-bends and vertical seismic forces are negligible by comparison.

2.2.2 Coefficient of Friction at Tube-to-Tubesheet Interface In order to determine the coefficient of friction between hard-rolled tubes and the tubesheet, pull tests and hydraulic proof tests were conducted on 3/4" diameter Alloy 600 tubing, hard rolled into carbon steel collars with an OD to simulate tubesheet rigidity (similar to the tests described in Section 2.1), After rolling, an inside circumferential cut was machined through the wall of the tube at a controlled distance from the bottom of the roll transition. The samples were heat soaked at [

]'" to simulate the possible effect of reduced j

preload force due to tube yielding during manufacturing heat treatment. Two sets of pullout tests were conducted, on a tensile testing machine in air at room temperature, and with intemal pressure as the acting force on the tube. The pressure tests were performed at room t:mperature using deionized water The pull tests performed with the tensile testing machine showed a static coefficient of friction of [

]'"

For samples which were expelled from the collars during the hydraulic proof tests, the coefficient of friction was determined to range j

from [

]'".

For tubes that leaked before tube expulsion, resulting in termination j

of the tests before expulsion, the lower bound coefficients of friction were determined to range from [

]'". On the basis of these results, the use of a coefficient of friction of [

]'" is considered to be conservative for the 7/8" tubes at Prairie Island Units 1 and 2 for application in determining the required engagement distance to resist tube pullout i

forces.

i 2.2.3 End Effects j

For a tube which is postulated to be severed within the tubesheet there is a material i

discontinuity at the location where the tube is severed. For a small distance from each 2-7

-y l~

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

m / u,, = e**

• cosine ( A

• x )

(6) u

- where, A=[

]'" =. end effect constant.

i x = distance from the end of the tube.

i For the rs,dially preloaded tube, the distance for the end effects to become negligible is the location where the cosine term becomes zero. Thus, for the roll expanded 51 Series tubes the i

distance corresponds to the product of "A" times "x" being equal to (x/2) or [

]'" inch.

Figure 2-1 shows a roll expansion which is postulated to be severed at the bottom of the F*

i region. For a distance of [

]'" inch above the severed end and below the bottom of the roll transition, the expanded joint has a reduced radial load carrjiing capability relative to the remainder of the F* length. The effective radial preload car.ied by these "end-affected" regions is calculated as follows.

i The above equation can be integrated to find the average deflection over the affected length to be 0.384 of the end deflection. This means that on the average the stiffness of the material i

over the affected length is 0.616 of the stiffness of the material remote from the ends.

Therefore, the effective preload for the affected end lengths is 61.6 percent of the preload at regions more than [

]'" inch from the ends. For example, for the normal operating net j'

preload of [

]'" psi or [

]'" pounds per inch of length, the effective preload for a distance of [

]'" inch from the end is [

]'" pounds per inch or [

]'" pounds.

2.2.4 Calculation of Engagement Distance Required, F*

The calculation of the required engagement distance is based on determining the length for preload frictional forces to equilibrate the applied operating loads. The axial friction force i

was found as the product of the radial preload force and the coefficient of friction between the tube and the tubesheet. The value assumed for the coefficient of friction was [

]'",

i from Reference 5. For normal operation the radial preload is [

]'" psi or [

]' "

pounds per inch of engagement. Thus, the axial friction resistance force is [

]'" pounds per inch of engagement. It is to be noted that this value applies away from the ends of the tube. For any given engagement length, the total axial resistance is the sum of that provided by the two ends plus that provided by the length minus the two end lengths. From the preceding section the axial resistance of each end is [

]'" pounds. Considering both ends of the presumed severed tube, i.e., the hardroll transition is considered one end, the axial l

resistance is [

]'" pounds plus the resistance of the material between the ends, i.e., the totallength of engagement minus [

]'" inch. For example, a one inch length has an axial resistance of, 7

[

]'"

i 2-8

o Conversely, for the maximum normal operating pressure applied load of [

]'" pounds, considered as [

]"' pounds with a safety factor of 3, the length of hardroll required is given by, F* = [

J'"

Similarly, the required engagement length for faulted conditions can be found to be 0.75 inch using a safety factor of 1.43 (corresponding to a ASME Code safety factor of 1.0/0.7 for i

allowable stress for faulted conditions).

i The calculation of the above values is summarized in Table 2-4. The F* value thus determined for the required length of hardroll engagement below the BRT, for normal operation is sufficient to resist tube pullout during both normal and postulated accident condition loadings.

Based on the results of the testing and analysis, it is concluded that following the installation 3

j of a tube by the standard hardrolling process, a residual radial preload stress exists due to the q

plastic deformation of the tube and tubesheet interface. This residual stress is expected to restrain the tube in the tubesheet while providing a leak limiting seal condition.

l 2.3 Limitation of Primary to Secondary Leakage 1

The allowable amount of primary to secondary leakage in each Prairie Island Units 1 and 2 steam generator during normal plant operation is limited by plant technical specifications to a 4

total of I gpm for both SGs. This limit, based on plant radiological release considerations and implicitly enveloping the leak before break consideration for a throughwall crack in the free span of a tube, is also applicable to a leak source within the tubesheet. In evaluating the j

primary to secondary leakage aspect of the F* criterion, the relationship between the tubesheet region leak rate at postulated FLB conditions (which bound SLB for primary-to-secondary pressure differential considerations) is assessed relative to that at normal plant operating conditions. The analysis was performed by assuming the existence of a leak path; however, no actual leak path would be expected due to the hardrolling of the tubes into the tubesheet.

1 2.3.1 Operating Condition Leak Considerations 1

In actuality, the hardrolled joint would be expected to be leak tight, i.e., the plant would not be expected to experience leak sources emanating below F*. Because of the presence of the tubesheet, tube indications are not expected to increase the likelihood that the plant would experience a significant number of leaks. It could also be expected, that if primary to secondary leakage is detected in a steam generator it will not be in the tube region below F*

Thus, no significant radiation exposure due to the need for personnel to look for tube /tubesheet leaks should be anticipated, i.e., the use of the F* criterion is consistent with ALARA considerations. As an additional benefit relative to ALARA considerations, precluding the need to install plugs below the F* criterion would result in a significant reduction of unnecessary radiation exposure to installing personnel.

4 2-9

s The issue of leakage within the F* region up to the top of the roll transition (RT) includes the consideration of postulated accident conditions in which the violation of the tube wall is very extensive, i.e., that no material is required at all below F*. Based on operating plant and laboratory experience the expected configuration of any cracks, should they occur, is axial.

The existence of significant circumferential cracking is considered to be of very low probability. Thus, consideration of whether or not a plant will come off-line to search for leaks a significant number of times should be based on the type of degradation that might be expected to occur, i.e., axial cracks. Axial cracks caused by primary water stress corrosion cracking have been found both in plant operation and in laboratory experiments to be short, about 0.5 inch in length, and tight.

From field experience, once the cracks have grown so that the crack front is cut of the skiproll or transition areas, they arrest.

Axial cracks in the free span portion of the tube, with no superimposed thinning, would leak at rates compatible with the technical specificatien acceptable leak rate. For a crack within the F* region of the tubesheet, expected leakage would be significantly less. Leakage through cracks in tubes has been investigated experimentally within Westinghouse for a significant number of tube wall thicknesses and thinning lengths (Reference 6). In general, the amount of leakage through a crack for a particular size tube has been found to be approximately proportional to the fourth power of the crack length. Analyses have also been performed which show, on an approximate basis for both elastic and elastic-plastic crack behavior, that the expected dependency of the crack opening area for an unrestrained tube is on the order of j

the fourth power, e.g., see NUREG CR-3464. The amount of leakage through a crack will be proportional to the area of the opening, thus, the analytic results substantiate the test results.

l The presence of the tubesheet will preclude deformation of the tube wall adjacent to the i

crack, i.e., the crack flanks, and the crack opening area may be considered to be directly l

proportional to the length. The additional dependency, i.e., fourth power relative to first j

i power,is due to the dilation of the unconstrained tube in the vicinity of the crack and the bending of the side faces or flanks of the crack. For a tube crack located within the tubesheet, the dilation of the tube and bending of the side faces of the crack are suppressed.

l Thus, a 0.5 inch crack located within the F* region up to the top of the roll transition would be expected to leak, without considering the flow path between the tube and tubesheet, at a rate less than a similar crack in the free span, i.e., less than the Prairie Island Units I and 2 administrative leak rate limit of 150 gpd (0.1 gpm). Additional resistance provided by the tube-to-tubesheet interface would reduce this amount even further, and in the hardroll region 8

the residual radial preload would be expected to eliminate it. This conclusion is supported by the results of the preload testing and analysis which demonstrated that a residual preload in excess of [

]* psi exists between the tube and the tubesheet at normal operating conditions.

l 2.3.2 Postulated Accident Condition Leak Considerations For the postulated leak source within the RE, increasing the tube differential pressure increases the driving head for the leak and increases the tube to tubesheet loading. For an initial location of a leak source below the BRT equal to F*, the FLB pressure differential results in an insignificant leak rate relative to that which could be associated with normal plant operation. This small effect is reduced by the increased tube to tubesheet loading 9

2 - 10 J-

~

n associated with the increased differential pressure as well as the tightening contribution of the tubesheet bending. Thus, for a circumferential indication within the RE which is left in service in accordance with the pullout criterion (F*), the existing technical specification limit is consistent with accident analysis assumptions. For postulated accident conditions, the preload testing and analysis showed that a net radial preload of about [

]'" psi would i

exist between the tube and tubesheet.

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

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

2.4 Tube Lock-up Considerations For the postulated condition in which one or more tubes become locked at the tube support plates due to the buildup of corrosion deposits, the evaluation has considered the axial forces on locked tubes relative to the strength of the F* region. In this postulated condition, tube i

lockup occurs during operation. Therefore, the axial force on locked tubes is essentially zero during operation.

1 For a single tube, locked at the lowest tube support plate, the tensile force during cold shutdown is approximately 220 lbs. If tube lockup were to occur, it is more likely to result in lockup of more than one tube; in this case, the tensile force per tube is reduced significantly from the single-tube value. For instance, for ten locked, adjacent tubes, the tensile force per tube reduces to approximately 24 lbs. at the shutdown condition.

Because the maximum locked-tube forces occur only during shutdown, there is no need to add them to the axial load of 2973 lbs., corresponding to 3 times normal operating AP, which

'is used to calculate the F* length. The 220 lbs. of axial force for a single locked tube during the cold shutdown condition will easily be accommodated by the F* length of tube joint. At this condition, i.e., without the beneficial, pressure-tightening and thermal growth contributions to the joint strength, the joint can accommodate approximately 1987 lbs. of axial force. This is approximately nine times the force that can be exerted on the joint by a single, locked, tube.

2 - 11

s 2.5 Tube Integrity Under Postulated Limiting Conditions The final aspect of the evduation is to demonstrate tube integrity under the postulated loss of coolant accident (LOCA) condidon of secondary to primary differential pressure. A review of tube collapse strength characteristics indicates that the constraint provided to the tube by the tubesheet gives a significant margin between tube collapse strength and the limiting secondary to primary differential pressure condition, even in the presence of circumferential or axial indications.

The maximum secondary to primary differential pressure during a postulated LOCA is

[

]* psi. This value is significantly below the residual radial preload between the tubes and the tubesheet. Therefore, no significant secondary to primary leakage would be expected to occur. In addition, loading on the tubes is axially toward the tubesheet and could not contribute to pullout.

2.6 Chemistry Considerations l

The concem that boric acid attack of the tubesheet due to the presence of a through wall flaw within the hardroll region of the tubesheet may result in loss of contact pressure assumed in the development of the F* criterion is addressed below. In addition, the potential for the existence of a tubricated interface between the tube and tubesheet as a result oflocalized primary to secondary leakage and subsequent effects on the friction coefficient assumed in the development of the F* criterion is also discussed.

2.6.1 Tubesheet Corrosion Testing Corrosion testing performed by Westinghouse specifically addressed the question of corrosion rates of tubesheet material exposed to reactor coolant The corrosion specimens were i

l assembled by bolting a steel (A336) coupon to an Alloy 600 coupon. The coupon dimensions were 3 inches x 3/4 inch x 1/8 inch and were bolted on both ends. A torque wrench was used to tighten the bolts to a load of 3 foot-pounds. The performance of A508 in testing of this nature is expected to be quite comparable to the performance of A336 (Gr. F-1) steel (a material used for tubesheet construction prior to A508). The arguments used in supporting the F* case relative to corrosion of tubesheet material due to minute quantities of primary coolant contacting the carbon steel were so conservative and had such margin that minor differences in material composition or strength would not change the conclusion.

1 The specimens were tested under three types of conditions:

j 1.

Wet-layup conditions 2.

Wet-layup and operating conditions 3.

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

2 - 12

(

l

~ _.

o While lithium hydroxide is normally added to the reactor coolant as a corrosion inhibitor, it was not added in these tests in order to provide a more severe test environment. Previous testing by Westinghouse has shown that the presence oflithium hydroxide reduces corrosion -

cf Alloy 600 and steel in a borated solution 'at operating temperatures.

Another set of specimens were used to simulate startup conditions with some operational-exposure. The specimens were exposed to a 2000 parts per million boron (as boric acid) solution for one week in the wet-layup condition (140'F), and 4 weeks at operating conditions (600'F,2000 psi). During wet layup, the test solution was aerated but at operating conditions the solution was deserated. The high temperature testing was performed in an Alloy 600 autoclave. Removal of oxygen was attained by heating the solution in the autoclave to 250'F and then degassing. This method of removing the oxygen results in oxygen concentrations of less than 100 parts per billion.

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

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

These corrosion tests are considered to be very conservative since they were conducted at maximum boric acid concentrations, in the absence of lithium hydroxide, with no special precaution to deaerate the solutions, and they were of short duration. The latter point is very l

significant since parabolic corrosion rates are expected in these types of tests, which leads one to overestimate actual corrosion rates when working with data from tests of short duration.

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

i 2.6.2 Tubesheet Corrosion Discussion i

At low temperatures, e.g., less than 140 degrees F, aerated boric acid solutions comparable in strength to primary coolant concentrations can produce corrosion of carbon steels. Demerated solutions are much less aggressive and deserated solutions at reactor coolant temperatures i

produce very low corrosion rates due to the fact that boric acid is a very much weaker acid at high temperature, e.g.,610 degrees F, than at 70 degrees F.

In the event that a crack occurred within the hardroll region of the tubesheet, as the amount L

ofleakage would be expected to be insufficient to be noticed by leak detection techniques and i

is largely retained in the crevice, then a very small volume of primary fluid would be 2 - 13

N involved. Any oxygen present in this very small volume would quickly be consumed by surface reactions, i.e., any corrosion that would occur would tend to cause existing crevices to narrow due to oxide expansion and, without a mode for replenishment, would represent a very benign corrosion condition. In any event the high temperature corrosion rate of the carbon steel in this very local region would be extremely low (significantly less than 1 mil per year).

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

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

2.7 Summary of F* Evaluation On the basis of this evaluation, it is determined that tubes with eddy current indications in the tubesheet region below the F* pullout criterion of 1.07 inches can be left in service. Tubes with circumferentially oriented eddy current indications of pluggable magnitude and located a distance less than F* below the bottom of the hardroll transition should be removed from service by plugging or repaired in accordance with the plant technical specification plugging limit. The conservatism of the F* criterion was demonstrated by preload testing and analysis commensurate with the requirements of RG 1.121 for indications in the free span of the tubes.

For tubes with axial indications, the criterion which should be used to determine whether tube plugging or repairing is necessary should be based on leakage since the axial strength of a tube is not reduced by axial cracks. Under these circumstances it has been demonstrated that significant leakage would not be expected to occur for through wall indications greater than

[

]"' inch below the bottom of the hardroll transition. Therefore, an F* distance of 1.07 inches achieves a leak limiting condition.

The application of F' in this report addresses existing roll expansions, performed in the factory. The approach may also be applied to roll expansions which are added to the SG T/TS joints of the partial depth roll expansion design, such as the design of the steam generators of Prairie Island Units 1 and 2. Based on past test and analytical evaluations for the retrofit-expansion case, the F* length of additional roll expansion is expected to be l

essentially equal to the F* length of factory roll expansion determined in this program. The j

F* distance for the additional roll expansion case would address the actual conditions of hard or soft sludge, partially or completely filling the tube-to-tubesheet hole crevice before the additional rolling process. The F* distance for the additional roll expansion case would provide both the resistance to pullout and resistance to significant leakage.

2 - 14

l Table 2-1 Model 51 SG Tube Roll Preload Test - Test Data Test Lesagen CeNer ID PreJtes Coller OD Pre Red Tube ID Beiere Roll Tube OD Before ReH No.

No.

g Dag. N Dog.

Avg.

O Dog. N Dog.

Avg.

O Deg. N Dec.

Avg.

O Dog. M Dog.

Avg.

a,s.e 1

4 2

3 4

5 6

Average 2

1 2

3 4

5 6

Average 3

1 2

3 4

5 6

Average 6

1 2

3 4

5 6

Average 7

1 2

3 4

5 6

Average 6

1 2

3 4

5 6

Average Col. Avgs' 2-15

Table 2-1 (Continued)

Model 51 SG Tube Roll Preload Test - Test Data TubeID Post-Rsu Test Lecosen Cellar OD PeeWteE Tube ID PeoWteN Collar Removed Pr Jtes Cellar TubeID No.

No.

Thichases e Deg. 90 Deg.

Aug.

Delts 0 Dog. 90 Dog.

Avg.

Growgi 0 Des. to Dog.

Ave.

a,c,s 1

1 2

3 4

5 6

Average 2

1 2

3 4

5 6

Average 3

1 2

3 4

5 6

Average 6

1 2

3 4

5 6

Average 7

1 2

3 4

5 6

Average 8

1 2

3 4

5 6

Average Col Avgs:

2-16

s 1

Table 2-1 (Continued)

Model 51 SG Tube Roll Preload Test - Test Data Test Laceses Tube oO Poe64tes Post-Thick.

Coller Redil Tube ID Coller Removed Roll nose Fles.

Rose Sprieg.

No.

No.

O Deg.

90 Deg.

Avg.

Thick.

Red.

du/dP, (4

Beck a.c.e 1

1 2

i 3

i 4

5 6

Average 2

1 2

3 4

5 6

Average 3

1 2

3 4

5 6

Average 6

1 2

3 4

a 5

6 Average 7

1 2

3 4

5 6

Average 8

1 2

3 4

5 6

Average Col Avgs Notes.

1.

All measured emensens are m mehes.

2.

The 00 stress e calculated using the measured ID sonngback j

3 The radi rate o a term that appears frequently in the analysis and s found as (OD*2+1D^2y(00^2 ID*2).

2-17

\\

e i

Table 2-2 Model 51 SG Tube Roll Preload Test - Stress Analysis Results Thennel Oper.

Test Lecases Tebe ID Tebe Tebe FleK.

00 OD 00 Exp.

Tebe Freesere Total Total Spring-FleK.

dU)dP, Radial Heep Aziel Radial Flak.

Radial Radial venuises No.

No.

Sect IU,IdP, Strees Strees Strees Strees dU)dP, Strese Strees Strese a,c.e 1

1 2

3 4

5 i

6 A erege 2

1 3

2 3

4 5

6 i

Average l

3 1

2 3

4 5

4 6

k:erage l

6 1

2 3

4 5

6 Average 7

1 2

3 4

5 6

Average 8

1 2

3 4

5 6

Average Cot Avgs-Notes' 1,

The 00 stress a cahulated immg the measured ID m 2-18 i-

~

e t>

Tebte 2-3 Calculation of F* Length for TITS tierdros Interface - Preme estand Uruts 1 and 2 Mentmwn flCS Presswee for e5 PCWG Casee

% Tube IICS Steem P+ Soc thneting i13 Dette P Tcold N.O.

Herdrea Total N O.

N O.

N O.

bef.-C.r se Plugging Pressure Presswo Dette P That Tcold Deff Temp flotto (2) T ratio (3)

Sep (4)

Srt151 Preioed Se tel Pr 17) Pond (88 Pe (9) 3Pa F* t10)

Ipetal Ipelet tpoiel Ideg. F1 Ideg. F1 (dog. F1 3,C t*

0%

2250 750 1500 599.1 535.5 2*

15 %

2250 857 1593 592.1 527.9 3**

3.67%

2250 700 1550 590 530

~

Man F * =

1.07 heferences:

Ceoes t end 2 represent design beels peremeters for 0% and 15% pkgging levate beoed on Westinghouse thermeWhydreulle performance cedoe.

Case 3 represente the current operet% peremotore et Prairie ledend Unite 1 and 2 es suppeled by Northern States Power.

9 l

I l

2-19 I

m w

le Table 2-4 Preload Analysis Summary Material Properties:

Tube /Tubesheet Dimensic.'s (Tested):

F

~

Elastic Modulus 28.7E+06 Init. Avg. Tube OD Poisson's Ratso 0.30 Init. Avg. Tube Thickness 1600 Coeff. of Therm. Exp.

7.74E-06 in/in/'F Init. Avg. Tubesheet ID TS Coeff. of Therm. Exp.

7.30E-06 inTm/'F Actual Thmnmg Operating AT 460*F Apparent Thincirg N.O. AP 1593 psi Faulted AP 2650 psi Additional Analysis Input:

Coefficient of Priction Tubesheet Bow Stress Reduction N.O.

FLB End Effects:

Mean Radius (Rolled)

Lower Tolerance Limit Factor Thickness (Rolled) 95/95 LTL 2.16 (N=36)

A End Effect Length Load Factor EVALUATION OF REQUIRED ENGAGEMENT LENGTH, F' Elastic Analysis:

N.O.

FLB a.c RT Preload (LTL)

Thermal Expansion Preload Pressure Preload Tubesheet Bow Loss Net Preload Net Radial Force Net Axial Resistance Applied Load Analysis Load End Effect Resistance Net Analysis Load End Effect Length i

Add. Length Required Total Length Requimd, F' 2-20

)

-o Table 2-4 (continued)

Preload Analysis Summary NOTES:

1) 95/95 Lower Tolerance Lunit rolled preload used.

2) For Normal Operation, a safety factor of 3.0 is used.

3) For Faulted Conditions, a safety factor of 1.43 is used (corresponding to ASME Code use of 0.7 on ultimate strength.

4) The required length does not include eddy current inspection uncertainty for the location of the bottom of the hard roll.

5) Preload stresses used were for the most stringent case, i.e., cold leg,' peripheral location.

This minimizes the thermal expansion preload and eliminates the preload due to tubesheet bowing.

l t

i 2-21

~.

+

e Top of Tubesheet

/

/

t

_ x

=-

w. _-

Top of Roll Transitionk

, Bottom of Roll Transition f

4

[

J,c.e a

Bottom of Roll Transition F*

g y.c.e a

o if j

l Postulated Severed End j

j

/

i

\\

1

}

/

J 1

i Bottom of Tubasheet i

4 Figure 2-1.

"End-Effect" Regions in a Tube Postulated to be Severed at the Bottom of F* Length 2-22 4

d 3.0 DEVELOPMENT OF RETROFIT F* AND EF* CRITERIA As stated earlier, the Prairie Island staam generators are of the partial-depth roll-expanded tube joint design. Using the F* criterion, operation is permissible with degradation in the factory roll expansion, as long as it is below a calculated distance, F*, below the factory roll transition. For tubes which exhibit degradation within the intended F* distance, within the factory roll transition, or within the unexpanded portion of the tube in the lower half of the tubesheet thickness, the roll expansion can be extended upward by a field-applied, modified.

after-operation (" retrofit") process called additional roll expansion. The additional roll

)

expansion of sound (undegraded) tubing is ideally equal to the F* length. In actuality, the additional roll must be slightly longer than F* to account for uncertainties in the nondestructive examination processe, which are used to determine crack tip elevations and the roll expansion elevations. The top of the additional roll expansion must be beneath the neutral bending axis (NBA) of the tubesheet, which is approximately 10.665 inches above the i

bottom of tubesheet cladding (10.885 inches above the tube end). It is required that application of additional roll expansion and F* above the factory roll and below the NBA be consistent with a separate qualification, not included in this document, showing equivalency of the factory and retrofit roll expansions or a means of relating the performance of the retrofit roll to the factory roll. Performance consists of adequate pullout and leakage resistance. Development and qualification of the retrofit roll expansion at the factory roll top and below the neutral axis was outside of the scope of this program. The presence of foreign matter such as sludge or denting corrosion in the tube-to-tubesheet crevice may affect the additional roll expansion length; it may also affect acceptability of elevated additional roll expansion with respect to pullout and leakage resistance.

For tubes which exhibit degradation above the NBA, an elevated additional roll expansion and elevated F* (EF*) criterion can be used. The EF* criterion can be applied to the retrofit roll expanded portions of the tube joint above the NBA. The applicability of EF* is dependent upon the results of a separate qualification or evaluation, not included in this document or project, showing equivalency of, or a means of relating, the tube pullout and leakage resistance of the elevated additional roll expansion to the factory roll. Theoretically, the projected elevation of the elevated additional roll expansion and the associated EF* can range from the NBA to the top of the tubesheet. As in the use of the additional roll expansion below the neutral bending axis, the presence of foreign matter such as sludge or denting i

corrosion in the tube-to-tubesheet crevice may affect the elevated additional roll expansion i

elevation; it may also affect acceptability of the elevated additional roll expansion with cespect to pullout and leakage resistance. It is expected that if denting corrosion is present, it v ould occur primarily in the vicinity of the tubesheet top, probably affecting or preventing retrofit roll expansion there.

The total normal operation primary-to-secondary leakage allowed from both s: cam generators is specified in the Prairie Island Units 1 and 2 Technical Specification as I gpm (1440 gpd).

This is further interpreted to allow leakage of up to 1 gpm in either SG, not to exceed a total of I gpm in both SGs. However, both units have adopted a more conservative administrative leak rate limit of 150 gpd (~ 0.1 gpm) during normal operation. The applicability of F* and EF* is dependent upon the results of a separate qualification or evaluation, not included in this document or project, showing that the additional roll expansion and elevated additional 3-1

4 roll expansion provide negligible leakage, such that the total leakage (including that from F*

and EF* tubes in addition to other sources) is below the permissible leakage of 150 gpd.

Tube-to-tubesheet contact pressure in the additional roll expansions and elevated additional roll expansions follows the same considerations as discussed in Section 2.0, except that the effect of tubesheet bow loosening must be considered in the calculation of elevated additional roll expansions follows for the tubesheet-upward-bending conditions. Unlike the condition for F* length requirements below the NBA, the reduction cf contact pressure above the NBA must be considered for EF* for N.O. and other tubesheet-upward-bending conditions such as feedline break /steamline break (FLB/SLB). The contact pressure reductions due to upward bending conditions must be subtracted from the beneficial effects of thermal growth mismatch, differential pressure tightening and the as-installed interference fit contact pressure.

Due to the reduced tube radial loading, a greater length of sound tube expansion is required above the NBA than below it. More relaxation of the radial contact pressure due to upward tubesheet bending as elevation increases; consequently, the maximum calculated EF* occurs at the top of the tubesheet.

In the calculation of EF* at a given elevation, the tubesheet bow relaxation effect is a function of radial location from the tubesheet vertical centerline. The most relaxation occurs at the location of tubesheet maximum rotation. This occurs at a radius of approximately 7.00 inches from the tubesheet vertical centerline. At the bundle periphery, the loosening effect is a minimum. Therefore, EF* ranges from the minimum EF" value at th periphery, at a given elevation, to the maximum value at the seven-inch radius point. Due to the variation in relaxation as a function of distance from the tube bundle center line, the EF* value could be calculated as a function of location. However, in the interest of simplicity of application,it is recommended that a constant EF* length, at a given elevation, be used; the maximum l

tubesheet bow relaxation is assumed to apply for the entire tubesheet region. As with F*,

EF* is calculated for the cold leg, where the thermal growth mismatch is also minimized.

Therefore, the calculated values of EF* are conservative for application to the hot leg, where it will primarily be applied.

During LOCA, the secondary side pressure is approximately 1000 psi; the primary side pressure approximates zero. Thus, the tubesheet bends downward and the tube-to-tubesheet contact pressure will increase for most of the tubes at the tubesheet top and proportionately less for lower elevations, down to the NBA. The action of the differential pressure is such that the tube is pushed toward the tube-to-tubesheet weld; hence, the LOCA case is of no consequence to the determination of EF*

3.1 Details of Determination of EF* Based on Structural Requirements Calculations for EF* use the same methodology as the F* calculations in Section 2; hence, they are not shown here. Figure 1-3 depicts the EF* length. Calculated values of EF* for several elevations above the NBA are shown in Table 3-1. The only difference between the F* and EF* calculations is the inclusion in the EF* calculations of a relaxation in radial preload due to bending of the tubesheet. (For the F* evaluation, the addition of radial preload below the NBA due to tubesheet bending was conservatively ignored, i.e. assumed to be zero). As shown in the F* evaluation, the normal operation condition results in the limiting 3-2

s values of EF* (longer EF* lengths are calculated for FLB/SLB conditions); hence, the EF*

values shown in Table 1 are for the more limiting normal operating conditions, and FLB results are not shown. The calculated values of EF* range from 1.46 inch for when the top of the elevated additional roll expansion is 4 - 6 inches below the top of the tubesheet, to 1.78 inches when the elevated additional roll expansion is at the top of the tubesheet. These values do not include an additional length for eddy current uncertainty.

3.2 Limitation of Primary to Secondary Leakage The allowable amount of primary to secondary leakage from both Prairie Island steam generators is limited to 150 gpd (-0.1 gpm) during normal operation. In evaluating the number of tubes which can be dispositioned by EF*, the average leakage per tube may be considered if the additional roll expansion / elevated additional roll expansion process, excluded from this program, produces finite leakage. (Leakage is not an issue for the F* region of the factory roll expansions because factory roll expansions of F* length have exhibited essentially zero leakage.)

3.3 Summary of EF* Evaluation This evaluation determined that tubes with indications of degradation in the as-fabricated unexpanded portion of the tube joint may be modified by additional roll expansion above the factory roll and kept in service. Lengths of EF* were determined for retrofitted roll expansions which produce the same or greater interference fit radial contact pressure between the tube and tubesheet as the factory roll and which have the same coefficient of friction between the tube and tubesheet. The EF* lengths for additional roll expansion / elevated additional roll expansion processes which produce different contact pressures and/or pullout resistances may be determined by using the same methodology as shown in the EF* tables.

The values for EF* for three elevations above the tubesheet NBA are shown in Table 3-1.

The "4.0 - 6.0 inches down" elevation is probably the most practical elevated additional roll expansion location because it will generally avoid denting corrosion at the tubesheet top but it will still address tube indications up to 15.40 inches above the tube ends. For this elevation, the EF* length is 1.46 inches, not including eddy current uncertainty. (The elevated additional roll expansion length of 2.00 inches was assumed in this case; the actual length would be determined in the elevated additional roll expansion program, which was outside of the scope of this program.)

3-3

i o

a Table 3-1 Evaluation of Required Engagement Length (U, EF*

for Prairie Island Units 1 and 2 15% Plugging Level 1

l EF* (Applies when top of EF* (Applies when top of EF* (Applies when top of ELASTIC elevated additional roll elevated additional roll elevated additional roll ANALYSIS expansion is 4D to 6.0 expansion is 2.0 to 4.0 expansionis at top of inches down from top of inches down frosa top of tubenheet and tubesboet, and tubesheet, and conservatively at any lower conservatively at any lower anservatively at any lower elevation above or below elevation above or below elevation above or below NBA*)

NBA*)

- NBA*)

a.c RT Preload (LTL), poi Thermal Expansion Preload, psi

]

l Pressure Preload, psi Tubesheet Bow loss, psi Net Preload, psi Net Radial Force, lbs/ inch Net Axial Resistana, Ibelinch Applied lead, lbs Analysis load, lbs (N.Op.: SF=3; FLB:

SF=1.43)

End Efect Resistance, Ibs Net Analysis load, Ibs.

End ENect Length, inch ( x 2 for total)

Additional Length Reg'd, inch Total bagth Re(d.

1,46 1.62 1.78 EF*, anches Notes from table:

(1) EF* distances determined do not include NDE uncertainty for elevation of ECT indications.

(2) Neutral bending axis (NBA) is 10.515 inches below top of tubesheet, or appmximately 10.665 inches above bottom of tubesheet cladding (10.885 inches above tube end).

(3) Limiting EF* is determined by Normal Operation condition.

3 1

s v

4.0 REFERENCES

1.

United States Nuclear Regulatory Commission, Regulatory Guide 1.121, " Bases for Plugging Degraded PWR Steam Generator Tubes," August,1976.

2.

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

3.

Grimison, E.

D., and Lee, G.

H., " Experimental Investigation of Tube Expanding,",

Transactions of the A.S.M.E., July,1943, pp. 497-505).

4.

ASME Boiler and Pressure Vessel Code,Section III, " Rules for Construction of Nuclear Vessels," The American Society of Mechanical Engineers, New York, New York, 1965 Edition.

5.

WCAP-ll241, "Tubesheet Region Plugging Criterion for the Central Nuclear de ASCO, ASCO Nuclear Station Units 1 and 2 Steam Generators", October,1986. (Proprietary) 6.

WCAP-10949, "Tubesheet Region Plugging Criterion for Full Depth Hardroll Expanded Tubes," Westinghouse Electric Corporation, September,1955. (Proprietary) 7.

WCAP-14225, Revision 0, "F* and L* Tube Plugging Criteria for Tubes With Degradation in the Tubesheet Region of the Prairie Island Units 1 and 2 Steam Generators", December,1994. (Proprietary) 1

b e

Exhibit F Prairie Island Nuclear Generating Plant i

License Amendment Request Dated September 24, 1996 Westinghouse Authorization Letter, CAW-96-1010, and Accompanying Affidavit Proprietary Information Notice Copyright Notice i