ML17219A262

I Steam Generator Allowable Tube Wall Degradation.

ML17219A262
Person / Time
Site:	Saint Lucie
Issue date:	10/31/1986
From:	ABB COMBUSTION ENGINEERING NUCLEAR FUEL (FORMERLY
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ML17219A259	List: ML17219A258 ML17219A261 ML17219A262
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CENC-1747, NUDOCS 8612150346
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ST. LUCIE I STEAM GENERATOR ALLOWABLE TUBE WALL DEGRADATION CENC-1747 October 1986 8612~503+ 8gg212 5000335 pgp ggQCK 05 pgR p

TABLE OF CONTENTS PAGE

SUMMARY

AND CONCLUSIONS ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ 1 II. INTRODUCTION ...... 2 III. GEOMETRY ....... 3 IV. DEVELOPMENT OF HYDRAULIC LOADS A~ Intl oductl On ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~

B. Loss-of-Coolant Accident 1 ~ Thermal-Hydraulic Model 5

2. Assumptions ....... 7

3. Operating Conditions ......... 8 Results ............................................

C. Main Steam Line Break

1. Thermal-Hydraulic Model ............................ 9

2. Assumptions ........................................ 9

3. Operating Conditions ..................... ~ ~ ~ t ]O

4. Results ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~

10 V. DEVELOPMENT OF MECHANICAL LOADS A. Safe Shutdown Earthquake ......

B. LOCA and MSLB Impulse Response ~ ~ ~ ~ ~ ~ ~ ~ ~ 11 C. Piessure ...... ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~

12 VI. ALLOMABLE STRESS DETERMINATION A. Allowable Stresses for Tubes ........................... 13 VII. LOCA PLUS SSE STRUCTURAL ANALYSIS A. Finite Element Model ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ 14 B. LOCA Analysis .... ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ t ~ ~ ~ 15 VIII. MSLB PLUS SSE STRUCTURAL ANALYSIS A. Loadings

1. Hydraulic Flow Loads ............................... 17

2. SSE Loads ..... ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~

17

3. Pressure Loads ..................................... 17

PAGE B. Finite Element Model 17 C. Stress Results

1. Hydraulic Flow 18

2. SSE .. 18

3. Pressure .. 18 D. Resultant Stresses 19 IX. NRC STAFF CRITERIA FOR MINIMUM ACCEPTABLE TUBE WALL THICKNESS 20 X. ALLOWABLE TUBE WALL DEGRADATION ............................ 21 XI. REFERENCES 22 APPENDICES A. STRUCTURAL GEOMETRY AND FINITE ELEMENT MODEL ....-....'.. A,O 8 THERMAL HYDRAULIC MODEL AND RESULTS . . . - B.O C. LOCA + SSE STRUCTURAL ANALYSIS RESULTS ...-............. C;0 D. RESULTS FROM EPRI/CE PROJECT S144 . ~ ~ ~ ~ ~ ~ o D~ P E. EVALUATION PER NRC STAFF CRITERIA .. ~ ~ ~ ~ ~ ~ ~ E~O F. FATIGUE EVALUATION . .... .... ..................... F.O 11

I.

SUMMARY

AND CONCLUSIONS This analysis establishes the allowable tube wall degradation for the St. Lucie I steam generator based on the requirements of Reg. Guide 1.121 (Reference 1). The analysis considers the tube loadings due to normal operation, Loss of Coolant Accident (LOCA), Main Steam Line Break (MSLB) and Safe Shutdown Earthquake (SSE). Allowable tube wall degradation is established in accordance with the ASME Code Section III allowables (or margins) consistent with the provisions of Reg.

Guide 1.121.

The following summarizes the results of this analysis:

A. The criteria on normal operating differential pressure of Reg.

Guide 1.121 controls the allowable tube wall degradation for all regions of the tube bundle with the exception of the upper re-gions (above the uppermost support plate) of Tube Rows 117 through 120 which. are governed by LOCA + SSE loads.

B. The allowable tube wall degradation is 63K for all regions except for the regions above the uppermost support plate for following Tube Rows:

a) Tube Row 117 - 59%

b) Tube Row 118 - 60K c) Tube Row 119 - 61%

d) Tube Row 120 62%

C. The allowable tube wall degradation determined from consideration of NSLB and SSE is a minimum of 66K and is not controlling.

I I. INTRODUCTION The analysis presented herein is performed to establish the maximum allowable tube wall degradation for the St. Lucie 1 Steam Generator tubes per the requirements of NRC Reg. Guide 1.121. The results of this study will be used as the bases for a steam generator technical specification change to the tube plugging criteria.

This report addresses the requirements of NRC Reg. Guide 1.121 for maximum allowable differential pressures during normal operation and accident conditions as well as the ASME Code Section III Appendix F requirements for faulted load conditions of Loss of Coolant Accidents (LOCA) plus Safe Shutdown Earthquake (SSE) loads and Main Steam Line Break (MSLB) plus SSE loads.

During the performance of this analysis, consideration is given to the tube support plate modifications wherein the outer rims and lugs of the upper tube support plates were removed. This allows the upper portion of the tube bundle to move laterally through a gap and con-tact the shroud during a LOCA event. Tube rows 87 through 140 are evaluated in this analysis with the shorter rows considered to be non-controlling due to decreasing LOCA forces resulting from shorter horizontal spans and decreasing moment arm lengths between the point of load application and the support locations.

III. 'GEOMETRY The St. Lucie I steam generator tube bundle is constructed from 0.75 inch O.D. X 0.048 inch wall tubes and is supported by grid type (eggcrate) tube supports in the axial flow region. (See Figure A.1.).

The tube bundle is restrained in the cross-flow region by several different types of tube supports. The vertical straight leg portion of the tubes is restrained by two partial eggcrates and two drill plates. The horizontal tube span inside the 90 arc 10 inch radii bends is supported by slotted vertical strips containing. horizontal strips welded into aligned slots. This support arrangement provides vertical in-plane and out-of-plane motion restraint and is an inte-gral part of the diagonal supports which extend across the 10 inch radius bend of all tubes. This support assembly is unique to the CE design and, due to its firm attachment to the restraining I-beams, provides excellent lateral as well as vertical support to the tubes.

Past modifications to the two drill plate supports consisted of re-moving their outer rims and part of the attaching lugs leaving a gap whereby lateral motion can occur.

The plates were "staked" by expanding a tube sleeve into the tube at the plate junction at nine locations on the lower plate and five places on the upper plate.

IV. DEVELOPMENT OF HYDRAULIC LOADS A. Intr oducti on A Loss-of-Coolant Accident (LOCA) produces a rarefaction wave which propagates at the speed of sound away from the break location. As the rarefaction wave passes through the tubes in the bend region of the steam generator, it imparts a lateral pressure loading on the tube bundle. The pressure loading on a particular tube is propor-tional to the pressure differ ence acting between the midpoints of the bends. Fluid friction and the centrifugal forces generated as the r

fluid negotiates the bends also contribute to the lateral loading on the tube bundle. The net force on a particular tube is the algebraic sum of the pressure, friction, and centrifugal forces.

A Main Steam Line Break (MSLB) produces a transient pressure loading on steam generator internals. The pressure loading results from the relative rates at which the secondary fluid leaves adjacent regions.

In general, the blowdown rate following a main steam line break de-pends upon the steam generator geometry, the secondary pressure, the

.secondary mass, and the nozzle area.

The thermal-hydraulic response of the primary (LOCA) and secondary (MSLB) systems during the postulated accidents is analyzed using Combustion Engineering's CEFLASH-4A computer code (Reference 4).

CEFLASH-4A is a one-dimensional, two-phase, thermal-hydraulic code which calculates the time-dependent behavior of the fluid state re-sulting from a flow line rupture or an operational transient. The solution proceeds by numerically integrating the momentum equation applied to the flow paths while maintaining mass and energy balances in the nodes. Heat transfer between the primary and secondary fluids is calculated.

The code contains its own water property tables and' number of user-selected options. Critical flow models include the Moody, the Henry-Fauske, and the modified Henry-Fauske/Moody correlations.

Friction factors may be user-specified or code-calculated to conform to initial conditions. The friction factors may be constant or Rey-nolds Number dependent during the transient. Available two-phase multipliers include those of Martinelli-Nelson and Thorn. The momen-tum flux term, which accounts for the pressure drop resulting from spatial changes in density and velocity, may be selectively included in the momentum equation. Finally, both homogeneous (mixed phase) and heterogeneous (separated phase) nodes may be chosen.

In 1982, an experimental program, sponsored in part by the Electric Power Research Institute, was completed in which the hydraulic load-ing predicted by the CEFLASH-4A code was compared with experimental measurements for a five tube model of a steam generator experiencing a LOCA. This analytical/experimental comparison provided validation of the code's thermal-hydraulic modeling and recommended that - for LOCA events which focus on the steam generator internals - the break be modeled as close to the steam generator as is physically possible (Reference 5).

B. Loss-of-Coolant Accident

1. Thermal-H draulic Model It has been established that maximum hydraulic loading on a steam generator tube in the bend region is realized during a double-ended guillotine break in the cold leg pipe. In the present analysis, the break is modeled at a steam generator primary outlet nozzle (Refer-ence 5). For a cold leg break, the hydraulic forces on a steam gen-erator tube in the bend region are illustrated in Figure B.l. The

pressure force resulting from the rarefaction wave is defined by the pressure difference between the midpoints of the bends (at the 45 positions) acting over the- cross-sectional flow area of the tube.

The time-varying average fluid friction force is based on the mass flow rate in the center of the horizontal span and is defined by:

'2 2fW vL

~0g where, f is the friction factor, dimensionless M is the flow rate in the tube, ibm/sec v is the specific volume, ft /ibm D is the tube inside diameter, ft.

L is the length between the midpoints of the bends, ft.

g is the universal gravitational constant, 32.17 ft-ibm/lb -sec 2 f

The resultant horizontal component of the centrifugal forces exerted by the fluid in negotiating the bends is neglected in the present analysis.

Ouring a loss-of-coolant accident, the secondary system remains in-tact except for heat transfer between the primary and the secondary fluids. Thus representation of the secondary systems (two steam generators) in the LOCA thermal-hydraulic model is remote. However, the finite speed of propagation and multiple reflections of the rare-faction wave which travels through the primary system require an accurate and complete representation of the primary system. The entire primary flow system is divided into a network of nodes and

flow paths as is shown in Figures B.2. In addition, it is necessary to provide more refined modeling for the steam generator that is the closest to the break location (Figure B.3).

The thermal-hydraulic detail shown in Figure 8.3 pertains to each of three tube rows modeled (tube rows 8140, 114 and 88). Node and flow path numbers not included in Figures 8.2 and B.3 are not utilized in the LOCA thermal-hydraulic model.

The following assumptions are utilized:

a. All nodes are homogeneous.

b. The properties of water and steam for establishing initial and boundary conditions are taken from Reference 6;

c. A double-ended guillotine break opening time of 0.020 seconds is assumed for the rupture of one 30-inch cold leg pipe at one pri-mary outlet nozzle. This break opening time is based on a reac-tor coolant system asymmetric load analysis (Reference 7).

d. The two-phase pressure drop is calculated by first specifying constant liquid phase friction factors based on the Moody diagram (Reference 8). These friction factors are then multiplied by the Thorn two-phase flow factor (Reference 9) stored within the code.

e. The momentum flux term is included for all internal flow paths.

f. Flow through the break is modeled using the modified Henry-Fauske/Moody critical flow correlation.

g. A discharge coefficient of 0.7 is assumed (Reference 10).

h. It is assumed that 15% of the steam generator tubes are plugged.

3. 0 eratin Conditions The LOCA event occurs at full power with the following plant operating conditions (Reference 11):

~Pr1mar Primary Temperature In, F 598 Primary Temperature Out, F 548 Primary Flow Rate, ibm/sec. 19367 Primary Pressure, .psia 2250

~Seconder Saturation Pr essure, psia 886 Feedwater Temperature, F 430.9 Feedwater Flow Rate, ibm/sec. 1641

4. Results The net force on a tube in each of the three tube rows modeled (tube rows 8140, 114 and 88) as a function of time is presented in Figures 8.4, 8.5 and 8.6, respectively. It can be seen that, for each of these tube rows, the maximum hydraulic loading is realized at approx-imately 0.015 seconds. The net force as a function of time for each tube row is input into the linear dynamic transient analysis.

The pressure in the middle of the bends as a function of time for each tube row is presented in Figures 8.7 through 8.12. The average flow rate in Flow Path 42 as a function of time for each tube row is presented in Figures 8.13 through 8.15.

C. Main Steam Line Break

1. Thermal-H draulic Model Previous analyses of a main steam line break for a wide range of operating conditions and different steam generator geometries indi-cate that peak pressure loads on steam generator internals are real-ized at either zero or low power operation. Peak pressure load across the tube bend region is realized at zero percent power. Our-ing the main steam line break, the rapid depressurization of the secondary fluid and its acceleration toward the break location are unaffected by the primary system. Therefore, the MSLB thermal-hydraulic model features only detail that pertains to the secondary system, as is shown in Figure 8.16. Node and flow path numbers not included in Figure B.16 are not utilized in the MSLB thermal-hydr aul ic model.

The following assumptions are utilized:

a. All calculations are made assuming no slip between the steam and the water.

• b. Proper ties of water and steam for establishing initial and bound-ary conditions are taken from Reference 6.

c. A double-ended guillotine break opening time of 0.001 seconds is assumed for the rupture of the 34-inch steam line at the steam outlet nozzle.

d. The two-phase pressure drop is calculated by first specifying constant liquid phase friction factors based on the Moody diagram Reference 8). These friction factors are then multiplied by the Thorn two-phase flow factor (Reference 9) stored within the code.

e. Flow through the break is modeled using the Moody critical flow correlation (Reference 12).

f. The momentum flux term is included for all internal flow paths.

g. A discharge coefficient of 1.0 is assumed.

3. 0 eratin Conditions The MSLB event occurs at zero power with the following plant operat-ing conditions (Reference 11):

~Seconder 0% Power Saturation Pressure, psia 900 Feedwater Temperature, F 375 Feedwater Flow Rate, ibm/sec. 0 uress

4. Results The pressure difference across the tube bend region (P9 - P6) as a function of time is presented in Figure B.17. The maximum pressure load of 28.6 psi is realized at approximately 0.075 seconds. The pressure in Nodes 9 and 6 as a function of time is presented in Fig-B.18 and B.19. The flow rate in Flow Path 7 (in the middle of the tube bend region) as a function of time is presented in Figure B.20.

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V. DEVELOPMENT OF MECHANICAL LOADS A. Safe Shutdown Ear th uake The project specification for St. Lucie I, Reference 3, states that the steam generator assembly shall be capable of withstanding a maxi-mum seismic loading equivalent to 0.5g's in both horizontal and ver-tical directions applied simultaneously through the steam generator supports. The evaluation of this loading condition is accomplished with an ANSYS (Reference 13) two-dimensional finite element model which is described in Section VII and is utilized for both SSE and LOCA evaluations. SSE stresses are obtained by the application of a static load equivalent to a +0.5g acceleration in the horizontal and vertical directions simultaneously. Two vertical load cases are considered since SSE + dead weight loadings are analyzed to determine the worst condition. However LOCA stresses are normally controlling, therefore, SSE stress locations are evaluated at the worst LOCA stress location.

B. LOCA and MSLB Im ulse Res onse The LOCA or MSLB accident produces an externally applied impulse to the steam generator caused by the fluid escaping from its respective loop. LOCA impulse stresses have been calculated for a unit of simi-lar design and have been found to be only + 2.0 ksi. (Reference 14).

However, due to removal of the drill support plates, lugs and outer rim, gaps are present which allow impacting between the drill plate tube assembly and the baffle which increases tube bending stresses.

A value of + 4.0 ksi is used in this analysis based on the parametric evaluation of a gap condition presented in Appendix C.

Main steam line break impulse loadings are conservatively estimated at +6.0 ksi for use in this analysis based on a worst case location 11

in the cross flow region for a unit of similar design and 0.75" 0.0.

X 0.008" wall tubes.

C. Pressure Ouring the LOCA event a tube is subjected to a net pressure force which produces an axial force in the vertical straight portion of the tube. With the secondary pressure remaining approximately constant during the LOCA event at 815 psia, a "blow-off" differential pressure stress is determined- based on this pressure and the primary pressure at the time of maximum LOCA stresses. The primary pressure, is ob-tained from the "CE-FLASH" program results and is 1336 psia at the time and location of maximum tube stress.

The pressure differential for MSLB is conservatively taken to be 2250 ps'his is based on the operating primary pressure with the as-sumption that the secondary pressure has decayed to zero.

12

VI. ALLOWABLE STRESS DETERMINATION The 'basis for the allowable stresses used in this analysis isSection III of the ASME Code for Nuclear Power Plant Components. Values for S , yield strength and ultimate strength at operating temperatures are taken directly from the appropriate tables in the Code.

A. Allowable Stresses for Tubes The ultimate strength for the SB-163 Inconel tubing is S u

= 80.0 ksi 0

,at the maximum operating temperature of 600 F. Appendix F of Section III gives the membrane stress allowable for the faulted conditions considered in this report as:

S = 0.7 S b

or S = 56.0 ks>

The membrane plus bending allowable is S

b f . 0.7 S

+ bend Mhere f is a function of cross-sectional shape as well as the ratio of membrane and bending stresses to yield stress. An interaction curve, Figure E.4 is shown in Appendix E for a 63 percent degraded tube and a healthy tube.

The ratio of membrane to yield stress for the LOCA analysis is ap-proximately 0.2~ Therefore, f is equal to 1.44 and the allowable stress intensity is:

S.I. = S

+b

= 1.44 (56) = 80.6 ksi 13

VII. LOCA PLUS SSE STRUCTURAL ANALYSIS A. Finite Element Model An ANSYS finite element model is developed for the tube bundle geome-try described in Section III. Figure A.2 gives a schematic of the portion of the tube bundle that is- modeled for this analysis. The model is a two-dimensional representation of the tube bundle devel-oped by combining the tubes in Rows 87 through 140 into 19 lumped parametric rows as given in Table A.l. The model includes the por-tion of the tube bundle above the uppermost full eggcrate support with all the tubes that are captured by the. lower drill plate plus the first three tube rows that are inside the lower drill plate. The finite element model is shown in Figure A.3. The tubes are modeled as two-dimensional beam, ANSYS Stif 3, elements. The mass distribu-tion of the drill plates is represented by lumped mass, ANSYS Stif 21, elements at the nodes which are at the drill plate locations.

The effect of the drill plates is modeled by coupling the nodes in the "x" horizontal direction at the drill plate locations. The hori-zontal portion of the tubes are coupled to the vertical supports with one dimensional spring, ANSYS Stif 14, elements. The vertical sup-port and batwing areas and moments of inertia are weighted for each row based on the number of vertical supports and batwings in that row. The finite element model has 723 elements of the types listed above.

)

The model boundary"conditions are as follows: The tubes at the full eggcrate are fixed in the "x" and "Y" direction. The tubes are fixed in the "x" direction at the partial eggcr ate locations. The .7 inch gap which is described in Appendix C is modeled using the ANSYS gap module described in Section 3.2.14 of Reference 13.

The analysis is performed for healthy Inconel tubes - .75 inch O.D. X

.048 inch wall thickness. The tube elements are attributed an equiv-alent density, p, formulated from:

=

P (1/AT)(PT AT + PSAS + C Pf Af) where p = Density of tube material = .305 lbs/in p = Density-of fluid in tube = .0261 lbs/in pf = Density-of displaced fluid = .0045 lbs/in AT

= Area of tube material = .106 in AS

= Inside area of tube = .336 in Af = Outside area of tube = .442 in C = Virtual mass coefficient A virtual mass coefficient of Cm = 1.7 is applied to the tubes. The resulting density becomes .4157 lbs/in and the corresponding mass density is .001077, lbs-sec /in. A structural damping of 2X of criti-cal damping is used for the small diameter piping in accordance with AEC Regulatory Guide 1.61.(Reference 14).

The modulus of elasticity is 28.8 X 10 psi for the tubes and 26.3 X 10 psi for the vertical supports and batwings.

The fluid dynamic loads versus time for Tube Rows 140, 114 and 88 are given in Figures B.4, B.5, and B.6. These correspond with Model Rows 19, 10, and 1. The loads given in the above figures are for a single tube in each row. Therefore, these loads must be scaled and combined 15

for the finite element model. This is done by assuming that the loads vary linearly between the three rows so that the total load on any model row is the total of the interpolated single tube load mul-tiplied by the number of lumped tubes in that model row. The model tube row loads are applied as concentrated forces as a function of time in the horizontal direction acting at the center nodes of the respective horizontal tube spans. The analysis is the ANSYS Reduced Linear Dynamic Analysis. The results of this analysis are the ele-ment stresses and loads as a function of .time. The maximum stress in the stress time history for each model tube row is shown in Table C.l. The location of the maximum stress for each tube row is indi-cated in Figure C. l. A time history plot of the maximum bending stress for model rows ll through 15 is given in Figure C.2. These stresses are at Nodes 259, 284, 309, 334 and 359 of elements 184, 183, 185, 187 and 189, respectively as shown in Figure C.3. Figure C.4 gives a pl'ot of the displaced geometry at the time of the maximum LOCA stresses. Figure C.5 gives a time history plot of the dis-placements of Nodes 19, 469 and 259. The location of Nodes 19 and 469 are shown in Figure C. 1, while Node 259 is given in Figure C.3.

Figure C.6 gives a time history plot of the displacements of Nodes 457 and 459. The location of Node 959 is given in Figure C.l, while 459 is shown in Figure C.3..

The SSE seismic loading is evaluated using the same model as for the LOCA analysis. The SSE loads are applied as accelerations of -1.5G vertically and + .5 G horizontally. The resulting SSE stresses at the locations of maximum LOCA stress are also shown in Table C.l.

16

VI I I. MSLB PLUS -SSE STRUCTURAL ANALYSIS A. ~Loadin a

1. H draulic Flow Loads The tubes in the cross-flow region are subjected to an external flow induced pressure drop during the MSLB event. Details of the analysis procedure and results are presented in Section IV and Appendix 8, respectively. The maximum pressure drop across the bend region of the tube bundle is 28.6 psi . This pressure load is from Figure B. 17 and is applied as a constant loading which is conservative. The loading imposed on the horizontal span of each tube is based on the assumption that the force acting is proportional to the ratio of an individual tube projected area to the total cross-flow tube area of the bundle.

2. 'SE Loads As defined in Section V.A, 0.5 g loads were applied simultaneously in the horizontal and vertical directions with dead weight loads consid-ered in the vertical direction.

3. Pressure Loads The tube differential pressure is assumed to be 2250 psi and is based on the assumption that the secondary pressure has decayed to zero.

B. Finite Element Model The FEM is described in Section VII with detailed geometry plots presented in Appendix A. Briefly, the model is planar with Rows 87 through 140 modeled in 19 lumped parameter rows. Most of the model 17

rows have-three actual tube rows with combined mass, stiffness, and loadings. Selection of vertical strip areas and moment of inertias are weighted so that the dynamic characteristics of the tube bundle are not altered.

C. Stress Results The maximum bending stresses due to MSLB flow loads are 9.77 and 9.95 ksi at node locations 212 and 217, respectively. This corresponds to Tube Rows 110, 111 and 112 which are the shortest rows attached to all three vertical strips (see Figure A.2). Node 217 is located at the left vertical strip/tube position while 212 is in the 90 degree bend region.

2. SSE The seismic stresses at Nodes 212 and 217 occurring at the time of maximum NSLB stresses are 7.6'and 6.3 ksi, respectively. The value of 7.6 ksi is also the highest SSE stress for any location in the bundle. Tabulated values are included in Table C.l for all model tube rows.

3. Pressure The stresses due to a differential pressure of 2250 psi in the axial, hoop, and radial directions are:

PR. =

i 7.7 ksi 2t 18

PR.

~hoop = i = 15.3 ksi t

oradial = P 2

= 0 D. Resultant Stresses The resultant stress intensity is determined by combining NSLB and SSE stresses using the square root of the sum of the squares (SRSS) method plus the addition of pressure stresses.

01 [(9.77) + (7.6) ] = 12.4 ksi 8 Node 212

<2 = [(9.95) + (6.3) ] = 11.8 ksi 8 Node 217 The worst case stress, a1, is used to determine a maximum allowable TWD of 66 percent. The procedure for determining this value is pre-sented in Article E.l of Appendix E.

1 oaxial radial

= 12.4 + 7.7-0 = 20.1 < 1.44 (.7) Su

= 80.6 ksi 19

IX. NRC STAFF. CRITERIA FOR MINIMUM ACCEPTABLE TUBE WALL THICKNESS

1. Tubes with detected acceptable defects will not be stressed dur-ing the full range of normal reactor operation beyond the elastic range of tube material.

2. The factor of safety against failure by bursting under normal operating conditions is not less than three at any tube location where defects have been detected.

3. Crack-type defects, that could lead to tube rupture either during normal operation or under postulated accident conditions would not be acceptable.

These criteria are from Reference 1 and are in addition to the fault-ed allowables of Section III. Figure E.l presents the minimum re-quired thicknesses for tubes based on the above criteria.

'20

X. ALLOWABLE TUBE WALL DEGRADATION A structural-thermal hydraulic analysis of the St. Lucie 1 Steam Generator tube bundle indicates that the allowable TWD is controlled, for a majority of tube rows by NRC Reg. Guide 1.121 criteria. A maximum allowable TWD of 63 percent is found to be acceptable for all tube rows with the exceptions of 117 through 120 above the top drill plate. The maximum allowable TWD for these rows is controlled by the faulted condition, LOCA + SSE, and is 59, 60, 61, and 62 percent, respectively.

The procedure for faulted loads evaluation is presented in Appendix E. A tabulation of LOCA and SSE stresses is given in Table C.l of Appendix C. It should be noted that 97.9 percent of all tubes meet the requirements of Reg. Guide 1.121 for 63 percent TWD with only 178 tubes or 2.1 percent having allowable TMD's of 59-62 percent.

A fatigue analysis of a degraded tube is presented in Appendix F with the maximum usage factor, U, for a 63 percent degraded tube deter-mined *to be zero.

'21

XI. REFERENCES

1. Nuclear Regulatory Guide 1. 121, "Bases for Plugging Degraded PHR Steam Generator Tubes".

2. ASME Boiler and Pressure Vessel Code, Section III for Nuclear Vessels, 1986 Edition.

3. Engineering Specification for Steam Generator Assemblies for Saint Lucie No. 1, Specification No. 19367-31-2, Revision 13.

4. "CEFLASH-4A, A Fortran IV Digital Computer Program for Reactor Blowdown Analysis", SA-78-223, J. M. Betancourt, Combustion Engi-neering, Inc., Department 489, June 1973.

5. EPRI NP-2652, "Loads on Steam Generator Tubes during Simulated Loss-of-Coolant Accident Conditions", Project S144-1, Final Re-port, November 1982.

6. 1977 ASME Steam Tables, C. A. Meyer, et al., Third Edition, ASME, New York, N.Y., 1977.

7. "Reactor Coolant Systems Asymmetric Load Evaluation Program-Final Report for Calvert Cliffs 1 and 2, Millstone 2, Palisades and Fort Calhoun", June 30, 1980.

8. Moody, L. F., "Friction Factors for Pipe Flow", Transactions, ASME, Vol. 66, 1944.

9. Thorn, J. R. S., "Prediction of Pressure Drop During Forced Circu-lation Boiling of Mater ", Int. J. Heat & Mass Transfer, Vol. 7, 1964.

22

10. Combustion Engineering Report CENPD-252-P-A, "Blowdown Analysis Method", July 1979.

11. Florida Power 5 Light Company, JNS-MCI-86-161, Letter from J. E.

Noaba to J. N. Mesthoven, October 3, 1986.

12. Moody, F. J., "Maximum Flow Rate of a Single Component, Two-Phase Mixture", ASME Transactions, February 1965.

13. ANSYS Engineering Analysis System, Finite Element Computer Pro-gram, Revision 4;1, March 1, 1983, John A. Swanson, Ph.D.

14. AEC Regulatory Guide 1.61.

. 15. Combustion Engineering, Inc., CENC-1264 (Revision 2), "Analysis to Determine Allowable Tube Mall Degradation for Palisades Steam Generators", March 30, 1976.

16. Combustion Engineering, Inc., CENC-1170, "Analytical Report for Florida Power 5 Light Company Steam Generator", December, 1971.

17. Desi nin b Photoelasticit , R. B. Heywood, 1952.

23

APPENDIX A STRUCTURAL GEOMETRY AND FINITE ELEMENT MODEL FIGURE A.1 STEAM GENERATOR ELEVATION VIEW FIGURE A.2 UPPER TUBE BUNDLE GEOMETRY FIGURE A.3 ANSYS FINITE ELEMENT MODEL TABLE A.1 FINITE ELEMENT MODEL TUBE DATA A.O

h JQ Qj, FIGURE A.1 STEAM GENERATOR - ELEVATION VIEW A.l

BAFFLE.

BATWING 31 75 ROW 140 10" RADIUS (TYPICAL)

ROW 116 ROI1 110 DRILL TYPICAL SLOTS

-t PLATE OW 90 22 DRILL PLATE ROW 66 22 P.E C 26.5 ROW 56 OW 28 P.E.C.

25.5 TOP FULL EGGCRAITE SUPPORT FIGURE A.2 UPPER TUBE BUNDLE GEOMETRY A.2

TABLE A.l FINITE ELEMENT MODEL TUBE DATA TUBE ROWS TOTAL NO. MOMENTS OF AREA MODEL ROW ROW NO. OF TUBES INERTIA ( IN ) (IN )

87 66 88 65 195 1.2773 20.67 89 64 90 63 91 64 187 1.2249 19.822 92 60 ~

93 62 94 61 185 1.2118 19.61 95 62 96 61 97 60 180 1.1790 19.08 98 59 99 60 100 59 177 1.1594 18.762 101 58 102 57 6 103 56 170 1.1135 18.02 104 57 105 56 106 . 165 .. 1.0808 17.49 107 54 108 53 109 105 .6878 11.13 52 110 51 ill 112 50 51 152 .9956 16.112 113 50 114 49 10 194 1.2707 20.564 115 48 116 47 A.4

TUBE ROWS TOTAL NO. MOMENTS OF AREA MODEL ROM ROW NO. OF TUBES INERTIA (IN ) (IN )

117 46 118 45 91 .5960 9.646 119 44 12 120 43 129 .8450 13.674 121 42 122 41 13 123 40 120 .7860 12.72 124 39 125 38 14 126 37 .7271 11.766 127 36 128 35 15 129 34 102 .6681 10.812 130 33 131 30 16 132 29 87 .5699 9.222 133 28 134 27 17 135 28 74 .4847 7.844 136 29 137 20 18 138 17 49 .3210 5.194 139 12 19 140 ~ --=- ~

9 .05895 .954 A.5

APPENDIX B THERMAL-HYDRAULICMODELS AND RESULTS FIGURE B.1 HYDRAULIC FORCES ON A TUBE FIGURE B.2 LOCA THERMAL-HYDRAULICMODEL FIGURE B.3 LOCA THECAL-HYDRAULIC MODEL (STEAM GENERATOR DETAIL)

C FIGURE B.4 LOCA NET FORCE VS TIME (TUBE ROW 8140)

FIGURE B.5 LOCA NET FORCE VS TIME (TUBE ROW 8114)

FIGURE B.6 LOCA NET FORCE VS TIVE (TUBE ROW 888)

FIGURE B.7 AVERAGE PRESSURE BETWEEN NODES 39 AND 40 VS TIME (TUBE ROll lj140)

FIGURE B.8 AVERAGE PRESSURE BETWEEN NODES 44 AND 45 VS. TDK (TUBE ROM 8140)

FIGURE B.9 AVERAGE PRESSURE BETtJEEN NODES 39 AND 40 VS TIVE (TUBE ROW 8114)

FIGURE B.10 AVERAGE PRESSURE BETWEEN NODES 44 AND 45 VS. TIME (TUBE ROW f114)

FIGURE B.11 AVERAGE PRESSURE BETWEEN NODES 39 AND 40 VS TIME (TUBE ROW II88)

FIGURE B.12 AVERAGE PRESSURE BETWEEN NODES 44 AND 45 VS TIME (TUBE ROW PP88)

B.O

APPENDIX B - (CONTINUED)

THERMAL-HYDRAULICYODELS AND RESULTS FIGURE B.13 FLOW RATE IN FLOtl PATH 42 VS TIME (TUBE RON t'!140)

FIGURE B.14 FLOW RATE IN FLOW PATH 42 VS TIYE (TUBE ROW //114)

FIGURE B. 15 FLON RATE IN FLOW PATH 42 VS TIME (TUBE ROW f88)

FIGURE B.16 Y~LB THERYAL-HYDRAULICYODEL FIGURE B. 17 PRESSURE DIFFERENCE VS TIME (Pg - P6)

FIGURE B. 18 PRESSURE IN NODE 9 VS TIYE FIGURE B. 19 PRESSURE IN NODE 6 VS TIYE N

FIGURE B.20 FLOW RATE IN FLOt'/ PATH 7 VS TIYE

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ll 30 10 29 33 - 34 35 31 30 33 31 S. 1. Tank fl 36 32 S. 1. Tank f2 5 6 32 29 S. 1. Tank l3 Fill Systoa N Hode Fler Path

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, LOCA THERMAL-HYDRAULIC MODEL STEAM GENERATOR DETAIL)

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APPENDIX C RESULTS OF LOCA AND MSLB STRUCTURAL ANALYSIS ARTICLE C.l DISCUSSION OF TUBE STRESS AS A FUNCTION OF GAP WIDTH TABLE C.l LOCA AND SSE SEISMIC STRESSES FIGURE C. 1 LOCA MAXIMUM-STRESS LOCATIONS FIGURE C.2 TIME HISTORY OF LOCA BENDING STRESS; NODES 259, 309, 334 5 350 FIGURE C.3 GEOMETRY AT MAXIMUM STRESS LOCATION FIGURE C.4 LOCA DISPLACED GEOMETRY FIGURE C.5 TIME HISTORY DISPLACEMENTS; NODES 19, 257 5 469 FIGURE C.6 TIME HISTORY DISPLACEMENTS; NODES 459 5 759 FIGURE C.7 MAIN STEAM LINE BREAK DISPLACED GEOMETRY WITH MAXIMUM STRESS LOCATIONS FIGURE C.8 PARAMETRIC EVALUATION OF BENDING STRESS VS. GAP WIDTH

ARTICLE C.1 DISCUSSION OF TUBE STRESS AS A FUNCTION OF GAP MIDTH As discussed earlier, a modification to St. Lucie 1 upper tube support plates was performed in 1978 to remove the outer rim and supporting lugs.

Based on photographs taken at this time a gap condition for steam genera-tors "A" and "B" existed between the outer portion of these plates and the remainder of the support lug which is welded to the shroud. This gap varied from 0.7 to 1.3 inches at that time. No further data is available pres-ently, hence a parametric evaluation of the gap condition for LOCA rarefac-tion loadings was performed with the results illustrated in Figure C.8.

The maximum stress occurs for a gap condition of 0.70" which is used in this analysis. A gap of about 1.15 inches appears to produce the smallest stress. In conclusion, one can state that the gap condition has the poten-tial for changing the maximum LOCA stresses by 25 - 30 percent. Therefore the LOCA impulse stress obtained from an analysis of a similar steam gener-ator. with rigid drill plate supports can be amplified by a factor of 2.0, as was done in this analysis, and should yield conservative results.

TABLE C.1 LOCA AND SSE SEISMIC STRESSES MAXIMUM LOCA REDUCED LOCA SSE STRESS MODEL ROW KSI NODE 1 25.9 19.43 2.8 505 2 24.0 18.0 5.7 32 3 20.8 15;6 4.9 55 19.9 14.93 4.7 80 5 18.7 14.03 4.5 105 6 16.8 12.6 4.3 130 7 16.2 12.15 3.7 164 8 18.8 14.1 1.3 682 9 22.2 16.65 7.4 213 10 23.7 17.78 6.7 238 11 37.72 28.28 7.1 259 12 34.6 25.95 5.3 284 13 28.7 21.53 4.0 309 14 21.9 16.43 3.0 334 15 20.7 15.53 2.4 859 16 21.9 16.43 1.7 884 17 22.4 16.8 1.7 909 18 22.1 16.58 .7 934 19 23.0 17.25 4.1 964 1

The LOCA Stresses are reduced by 25'X per Appendix D.

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L APPENDIX D

'IMPORTANT FINDINGS FROM EPRI/CE PROJECT S144 A brief discussion regarding the above test program is presented in this section. The objectives of the program are (i) to verify the CEFLASH Code modeling of the fluid-dynamic loads in a steam generator tube during a LOCA and (ii) to verify the predicted structural responses.

The test loop simulated the primary side thermal-hydraulic conditions in an operational nuclear steam generator . The loop consisted of five full size double 90 0 bend tubes and steam generator plena, a pressurizer, a reactor resistance simulator, a heater, a pump, and associated pipes and valves to complete the system. The tubes used were of typical length and the same outside diameter as those used in Cf steam generators. Prototypical supports were provided for the bundle of five tubes. See Figure D. 1 which is a photograph of the test stand. Cold leg guillotine breaks were simulated using quick opening valve and rupture disks. Break opening times ranged from less than 1 msec to as much as 67 milliseconds. The loop instrumentation was designed to measure the transient pressure history at various locations and monitor the structural response of the tube to the LOCA hydrodynamic loading.

A series of blowdown tests were performed for different operating and boundary conditions. The parameter variations included fluid temperature, pre-blowdown flow rate, break opening time, break opening area, and break location. Both uniform and mixed length tube bundles were used.

Analytically predicted transient pressure histories and the differential pressure history across the tube span were compared with the experimental data. See Figure G.2. Predicted structural responses in the bend region were also compared with the test data. The transient pressure histories as predicted by CEFLASM were in excellent agreement with the test data.

The calculated structural 'responses of the tube also had good overall agreement with the test data.

0.0

During the course of the test program, mechanical tests were conducted to measure the structural damping of the horizontal tube span. It was found to be 8X of the critical damping. It was also observed that the frictional and "binding" forces in the vertical support reduced the predicted bending stresses an average 34.3X. See Table D, l. As shown in Table D. 1, the analytically predicted bending stresses in this report were lowered 25. 1X (34.3X - 1 standard deviation 9.2X = 25. 1%) to take account of the friction and binding.

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-0.3

TABLE D.t PALISADES STEAM GENERATOR 1.0CA ANALYSIS BATWINCi FRICTION - EPRI TEST EPRI TEST ANSYS ANAL Break Report Open T Report Correct With FP No FP X Due Test ID Description s s s s Frict (Sec) (ksi) (ksi) (ksi) (ksi)

A-28 Smaller Break Area .017 5.5 5.7 10.7 12.2 46. 7 (Downstream Orifice)

A-31 Short Break Length .016 6.4 6.6 9.8 11. 2 32 ~ 7 A-37 Mixed Height Bundle .014 5.8 6.0 9.0 10.3 33. 3 (L'ong Tube)

A-37 Mixed lleight Bundle .014 10. 3 10.7 14. 2 16.3 24.6 (Short Tube)

Average Value 34.3 Standard Deviation 9.2 Selected Value (Avg - S) 25.1

C APPENDIX E EVALUATION OF NRC STAFF CRITERIA ARTICLE E.l PROCEDURE FOR DETERMINING ALLOWABLE,TWD TABLE E.l RELATED DATA FOR ALLOWABLE TWD DETERMINATION FIGURE E.l EVALUATION PER NRC REG. GUIDE 1.121 FIGURE E ~ 2 ALLOWABLE TUBE WALL DEGRADATION BASED ON LOCA + SSE STRESSES FOR TUBE ROWS 117 - 123 FIGURE E.3 ALLOWABLE TUBE WALL DEGRADATION FOR TOTAL TUBE BNUNDLE FIGURE E.4 MEMBRANE AND BENDING INTERACTION DIAGRAM FOR CYLINDRICAL TUBE E,O

C ARTICLE E. 1 PROCEDURE FOR DETERMINING ALLOMABLE TMD LOCA + SSE The maximum allowable stress intensity for a degraded tube for LOCA + SSE stress is determined from the following equation:

2 2 . 2 1/2 S.I. LOCA SSE LOCA X A i +. a

'XP

- e'rP = f (0.7) Su RARE IMP (Equation E.l)

The LOCA rarefaction, impulse, and safe shutdown earthquake stresses are evaluated by the SRSS procedure for a healthy tube and then amplified by the factor A. which is the ratio of the section modulus of a degraded tube to a healthy one. Tube stresses due to a delta pressure, primary minus secondary, are included and compared to an allowable, f(0.7 S ) where f is u

a function of cross-sectional geometry and the ratio of membrane and bend-ing stresses to yield stress. An interaction graph is shown in Figure E.4.

This was developed for a cylindrical tube cross-section for membrane and bending stresses. Table E.l presents data calculated for use in the above equation for S.I.

The three model rows with maximum LOCA rarefaction stresses are ll, 12, and 13 corresponding to steam generator tube rows 117-118, 119-121 and 122-124, respectively. These values are shown in Table C.l.

Model Row 11 Evaluation: (Use TMD 59% Input)

S. I. = [(28.3) +(7.1) +(4) ] (2.559)+4.3+.82 = 80.5 < Allow. = 80.6 ksi

Model Row 12 Evaluation: (Use TWD 62% Input)

S.I. = [(26) 2 +(5.3) 2 +(4) 2 ) 1/2 (2.776)+4.7+.82 = 79.9 < Allow. = 80.6 ksi Model Row 13 Evaluation: (Use TWD 68% Input)

S.I. = [(21.5) +(4) +(4) ] (3.305)+5.2+.82 = 79.6 < Allow. = 80.6 ksi Figures E:2 and E.3 show the allowable TWD for the localized region and the entire tube bundle.

NSLB + SSE Allowable TWD is determined normally by considering stresses from four sources. 1) MSLB flow loads, 2) MSLB impulse loads, 3) Safe shutdown earthquake and 4) differential pressure. Stresses resulting from 1), 3),

and 4) have been determined. NSLB impulse related maximum stress is as-sumed to be +6 ksi, based on an analysis of a steam generator of similar design, and is used in this portion of the evaluation. It should be noted that NSLB impulse related stresses are generally smaller than MSLB flow stresses and with the combining of stresses utilizing the SRSS procedure, the effect of MSLB impulse stresses on the resultant value is further de-creased.

The worst case NSLB + SSE stresses are taken from Section VIII and used to evaluate the allowable TWD.

NSLB Flow Stress = 9.77 ksi SSE Stress 7.6 ksi MSLB Impulse Stress = 6.0 ksi Resultant = (9.77 + 7.6 + 6.0 ) = 13.7 ksi E.2

Solving equation. E.l using an allowable TWO estimate of 66 percent, the stress intensity, S.I. for MSLB + SSE is 65.3 ksi.

The allowable S. I. is f (.7) S = 66 ksi based on a value of f = 1.18 ob-u tained from Figure E.4. The ratio of membrane stress to yield for Inconel tubing is 0.81 as shown below and is required to determine f.

a m

=

2tr i = 2.25 '

.327 22.6 ksi m 22.6 = 0.81 .'. f = 1.18 from Figure E.4 Hence, the maximum allowable TWD for MSLB + SSE stresses is 66 percent which is non-controlling.

E.3

TABLE E.l RELATED DATA FOR ALLOWABLE TWD DETERMINATION TWD t R ZD H A.

xP (IN) (IN) IN IN ZH/ZD (KSI) 58 .0201 .3471 .006972 .017470 2.506 4.3 59 .0197 .3467 .006828 .017470 2.559 4.3 60 .0192 .3462 .006649 .017470 2.627 4.4 61 .0187 .3457 .006471 .017470 2.700 4.6 62 .0182 .3452 .006293 .017470 2.776 4.7

.0178 .3448 .006151 .017470 2.840 4.8 66 .0163 .3433 .005618 .017470 3.110 5.2 68 .0154 .3424 .005286 .017470 3.305 5.6 69 .0149 .3419 .005124 .017470 3.409 5.7 70 .0144 .3414 .004948 .017470 3.531 5.8 P> = 1336 psi P2 = 815 psi 4PRi hP = Pq - P2 = 0.521 ksi xP ~tr Where cr = 0.815 ksi

((rremb) =

5 g7 g

= 0. 18 . . f= 1.44 (Figure E.4)

SSE

= 2.2 ksi (Healthy Tube at Time and Location of Maximum LOCA Stress) 0 LOCA 4 0 ksi (Heal thy Tube)

IMP E.4

4

1. NORMAL OPERATION 2. NORMAL OPERATION 3. STEAM LINE BREAK a < S = 27.9 KSI, 600 F 3 a < S= 80.0 KSI, 600 F a < .7 S= 56.0 ksi, 600 F NORNL 36P R' ~P Ri OPERATING r S-. P +P S U

+ 07S- . P +P PRESSURES 2.25 .327 1.435 . 327 3( 1.435 .327 r

P>

= 2250 psia 3. 65 r 8 Pz = 815 psia t = 0.0178 in. tr = 0.0179 in, tr = 0.0134 in.

hP = 1435 psi X Allow. Degradation X Allow. Degradation X Allow. Degradation STEAN LINE

(.048 - .0178) X 100

.048 -,0179 0 100 0 100 BREAK

. 048 Pg = 0 = 63 = '72 Pi = 2250 psia FIGURE E. 1 EVALUATION PER NRC REG. GUIDE 1.121

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APPENDIX F FATIGUE ANALYSIS OF DEGRADED TUBES The fatigue evaluation of a degraded tube is based upon the conservative assumption of 15,000 cycles from ambient conditions to 1005 power. The resulting alternating stress intensity is (Ref. 16, Page A-465)

S lt = 16.8 ksi A stress concentration factor for a tube that is degraded 63$ is determined by considering a strip on the tension side of the tube and treating it as a shouldered plate in tension. (Ref. 17, Page 178).

Degraded Zone Transition Area A full fillet radius is assumed R = .03 h = .03 0.018" p d d = .036 T D = .096 d= D/d h 0.65 0 ~ 6s

~h-Y = 1.33 Applying. this stress concentration factor yields

= (1.33)(16.8 = 22.4 ksi.

Salt )

This alternating stress is good for cycles in excess of 106 (see Figure I-9.2, Reference 2). f/ence, the usage factor is E.O

4 r' ATTACHMENTE DETERMINATIONOF NO SIGNIFICANT HAZARDS CONSIDERATION The standards used to arrive at a determination that a request for amendment involves no significant hazards consideration are included in the Commission's regulations, 10 CFR 50.92, which states that no significant hazards considerations are involved if the operation of the facility in accordance with the proposed amendment would not (I) involve a significant increase in the probability or consequences of an accident previously evaluated; or (2) create the possiblity of a new or different kind of accident from any accident previously evaluated or (3) involve a significant reduction in a margin of safety. Each standard is discussed as follows:

(I) Operation of the facility in accordance with the proposed amendment would not involve a significant increase in the probability or consequences of an accident previously evaluated.

The proposed steam generator tube wall acceptance criteria has been determined in accordance with Regulatory Guide l.l2I. The acceptance criteria is based on margins of safety consistent with the margins provided in Section III of the ASME Boiler and Pressure Vessel Code. The demonstrated margins of safety provide reasonable assurance that tube failure will not occur during operating or accident conditions. Therefore, the proposed change will not result in a significant increase in the probability or consequences of an accident previously evaluated.

(2) Use of the modified specification would not create the possibility of a new or different kind of accident from any accident previously evaluated.

The proposed change does not change the configuration of the plant or the way in which it is operated. Therefore, the change does not create the possibility for a new or different kind of accident from any previously evaluated.

(3) 'Use of the modified specification would not involve a significant reduction in a margin of safety.

The supporting steam generator tube stress analysis meets the criteria of Regulator'y Guide I.I2I. The proposed acceptance criteria is based on safety factors of 3 for normal operating conditions and I.S for accident conditions. The associated margins of safety are equivalent to the margins determined by the stress limits of Section III of the ASME Boiler and Pressure Vessel Code. These margins of safety assure a low probability for tube failure during operating or accident conditions. Therefore, the proposed change does not result in a significant reduction in a margin of safety.

Based on the above, we have determined that the amendment request does not (I) involve a significant increase in the probability or consequences of an accident previously evaluated, (2) create the probability of a new or different kind of accident from any accident previously evaluated, or (3) involve a significant reduction in a margin of safety; and therefore does not involve a significant hazards consideration.

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