ML20214T406

Safety Evaluation Supporting Util 860424 Submittal Re Turbine Sys Maint Program for Early Detection of Cracking in Low Pressure Turbine Wheels

ML20214T406
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Site:	Catawba
Issue date:	06/02/1987
From:	Office of Nuclear Reactor Regulation
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NUDOCS 8706100247
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-m ENCLOSURE fp. . .:,;'o, UNITED STATES

! yy , f i NUCLEAR REGULATORY COMMISSION u- . 3, t IVASHINGT ON, D. C. 20555

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I SAFETY EVALUATION BY THE OFFICE OF NUCLEAR REACTOR REGULATION l

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RELATINGTOPROBABILITYOFMISSILEGENERATIONIN ,

GENERAL ELECTRIC NUCLEAR TURBINES, JANUARY 1984

SUMMARY

AND CONCLUSIONS The objective of the staff's review of the subject report was to evaluate and, if appropriate, to approve of the methods and procedures utilized by the l General Electric Company, Large Steam Turbine-Generator Department (GE) to determine specific turbine system inspection and testing intervals for their respective utility customers.

During the past several years, the staff has recommended a probabilistic ,

approach to determine turbine rotor inspection intervals and turbine control rys+e- nsintenance and testing frequencies so as to maintain the as-built turbine system integrity. The GE report describes sui:h an approach generically and, to the extent possible, supports it with test and turbine

'er cperating experience data. The staff recognizes that probabilistic .

analyses based on limited statistical data, especially for a complex system, will include inherent uncertainties. Nevertheless, when the overall approach includes conservative assumptions which overcome the uncertainties, then the ultimate results can be meaningful. I The staff concludes that the methodology described in the GE report is state-of-the-art and is acceptable for use in establishing maintenance and inspection schedules for specific turbine systems.

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Applicants or licensees who accept GE's recommendations, based on this report, should confirm their commitment to the staff and provide a description of their specific mairitenance and inspection program including graphs or tables of missile probability (P ) versus service time for their specific 3

turbine.

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1.0 BACKGROUND

l Although large steam turbines and their auxiliaries are not safety-related I systems as defined by NRC regulations, failures that occur in these turbines l can produce large, high energy missiles. If such missiles were to strike and to damage plant safety-related structures, systems, and components, they could render them unavailable to perform their safety function. Consequently, General Design Criterion 4, " Environmental and Missile Design Bases," of Appendix A, " General Design Criteria for Nuclear Power Plants," to 10 CFR Part 50, " Domestic Licensing of Production and Utilization Facilities," requires, in part, that structures, systems, and compon.ents important to safety be appropriately protected against the effects of missiles that might result from such failures. In the past, with regard to construction permit and operating license applications, evaluation of the effects of turbine failure on the public health and safety followed Regulatory Guide 1.115, " Protection Against

, - Low-Trajectory Turbine Missiles," and three essentially independent Standard .

Review Plan (SRP) Sections 10.2 " Turbine Generator," 10.2.3 " Turbine Disk Integrity," and 3.5.1.3 " Turbine Missiles."

According to staff guidelines stated in SRP Section 2.2.3 and Regulatory Guide 1.115, the probability of unacceptable damage from turbine missiles (P4 ) should be less than or equal to about I chance in 10 million per year.for an individual plant, that is, P 5 10 4

per year. The probability of unacceptable resulting damage from turbine missiles is generally expressed as the product of (1) the probability of turbine failure resulting in the ejection of turbine disc (or internal structure) fragments through the turbine casing (P 3

); (2) the pro-bability of ejected missiles perforating intervening barriers and striking safety-related structures, systems, or components (P2 ); and (3) the probability

of struck structures, systems, or components failing to perform their safety i function (P3)*

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.l 3 In the past, analyses assumed the probability of missile generation (Py ) to

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be approximately 10 per turbine year, based on the historical failure rate (Ref. 1). The strfke probability (P )2 was estimated on the basis of postu-lated missile sizes, shapes, and energies and on available plant-specific information such as turbine placem'ent and orientation, number and type of ,

intervening barriers, target geometry, and potential missile trajectories (See SRP Section 3.5.1.3 for a description of the evaluation procedures previously recommended by the staff.) The damage probability (P 3) was generally assumed to be 1.0. The overall probability of unacceptable damage to safety-related, systems (P 4

), which is the sum over all targets of the product of these probabilities, w'as then evaluated for compliance with the staff safety objective. This logic places the regulatory emphasis on the i strike probability, that is, it necessitates that P be made less than or 2

cqual to 10' , and disregards all the plant specific factors that determine the actual P yand its unique time dependency.

Although the calculation of strike probability is not difficult in principle, i for the most part being not more than a straightforward ballistics analysis, it presents a problem in practice. The problem stems from the fact that numerous modeling approximations and simplifying assumptions are required to incorporate available data into acceptable models. The available data t'

includes the following: (1) properties of missiles, (2) interactions of ,

aissiles with barriers and obstacles, (3) trajectories of missiles as they interact with and perforate (or are deflected by) barriers, and (4) identification and location of safety-related targets. The particular cpproximations and assumptions made tend to have a significant effect on the resulting value of P2. Similarly, a reasonably accurate specification of the damage probability (P 3

) is n t a simple matter because of the difficulty in defining the missile impact energy required to render given safety-related systems unavailable to perform their safety functions and the difficulty in postulating sequences of events that would follow a missile producing turbine failure.

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, 4 Operating experience shows that nuclear turbine discs crack (Refs. 2 and 3),

that turbine stop and control valves fail (Refs. 4 and 5), and that disc ruptures could result in the generation of high-energy missiles (Ref. 6).

Analyses (Refs. 5 ihd 7) show that missile generation can be modeled and the probability can be strongly influenced by inservice testing and. inspection .

frequencies. -

During the past few years, the results of turbine inspections at operating nuclear facilities have indicated that cracking to various degrees has occurred at the inner radius of turbine discs of Westinghouse design. Within this period, a Westinghcuse turbine disc failure occurred at one facility owned by the Yankee Atomic Electr'ic Company (Ref. 2). More recent inspections of General Electric turbines have also discovered disc keyway cracking (Ref. 3).

Stress corrosion has been identified by both manufacturers as the operative cracking mechanism.

In view of operating experience and staff safety objectives, the staff has shifted emphasis in the reviews of the turbine missile issue from the strike and damage probability2P3P to the missile generation probability Py and, l in the process, has attempted to integrate the various aspects of the issue into a single, coherent evaluation.

1 Thrnugh experience of reviewing various licensing applications, the staff has .

j concluded that P23 P analyses provide only " ball park" or " order of magnitude" I values. Based on estimates for a variety of plant layouts, the staff further concludes that the strike and damage probability product P23 P can be )

reasonably take to fall in a characteristic narrow range which is dependent on j the turbine generator orientation; i.e., (1) for favorably oriented turbine  !

generators P 23P lies in the range of 10 to 10 per year, and (2) for l unfavorably oriented turbine generators 23 P P lies in the range of 10 to  ;

-2 per year.

10 In addition, detailed analyses such as those discussed in this )

cvaluation show that, depending on the specific combination of material l properties, operating environment, and maintenance practices, P can have values ]

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5 from 10 to 10' per turbine year depending on the turbine test and inspection intervals. For these reasons, in the evaluation of P 4 (Py23 P P ), the probability of unadeeptable damage to safety-related systems from potential turbine missiles, the staff is giv.ing credit for the product of the strike and damage probabilities of 10' for a favorably oriented turbine and 10' per- ,

year for an unfavorably oriented turbine, and is discouraging the elaborate calculation of these values.

The staff believes that maintaining an initial small value of Py through I turbine testing and inspection is a reliable means of ensuring that the objectives precluding turbine mis'siles and unacceptable damage to safety-related structures, systems, and components can be mat. It simplifies and improves procedures for evaluation of turbine missile risks and ensures that the public health and safety is maintained.

To implement this shift of empha' sis, the staff recently has proposed guidelines for total turbine missile generation probabilties (Table 1) to be used for determining (1) frequencies of turbine disc ultrasonic inservice inspections and (2) maintenance and testing schedules for turbine control and Overspeed protection systems. It should be noted that no change in safety criteria is associated with this change in emphasis.

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l Table 1. Turbine System Reliability Criteria 5 ..

Probability per year - -

Favorably unfavorably Oriented Oriented Turbine -Turbine Required Licensee Action

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(A) P <10 2 P3 <10 This is the general, minimum

. reliability requirement for loading the turbine and bringing the system on line.

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(B) 10 <P3 <10 10 < P 2< 10 If this condition is reached dur-ing operation, the turbine may be kept in service until the next scheduled outage, at which time the licensee is to take action to reduce P to meet the appropriate A criterion (above) before return-ing the turbine to service.

.3 .2 .4 .3 (C) 10 <P3 <10 10 <Pi<10 If this condition is reached during operation, the turbine is to be isolated from the stea.m supply within 60 days, at which time the licensee is to take action to re-duce P3 to meet the appropriate A criterion (above) before returning the turbine to service.

.2 .3 (D) 10 <P 10 <P 2 If this condition is reached at any

. time during operation, the turbine is to be isolated from the steam supply within 6 days, at which time the licensee is to take action to reduce P2 to meet the appropriate A criterion (above) before return-ing the turbine to service. l l

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.' 0 SCOPE OF REVIEW There are essentially two modes of turbine failure; one due to rotor material failure at approximately the rated operating speed, or one due to failure of the overspeed protection systems resulting in excessive disc speeds.

Failures of turbine discs at or below the design speed, nominally 120 percent of normal operating speed, can be caused by small flaws or cracks left during fabrication or those that initiate during operation and grow to critical size either by fatigue. crack growth, by stress corrosion crack growth, or by a combination of both of these mech'anisms. Cracks in the bore or hub region of turbine discs could eventually lead to disc failure.

Failures of turbine discs at the destructive overspeed can result from a failure of the governor and overspeed protection systems, consisting of : (1) speed sensing and tripping systems and (2). steam valves. If the turbine is v; cf control, its speed can increase until failure occurs. For unflawed ,

oiacs, destructive overspeed is reached at about 180 to 190 percent of the normal operating speed. In general, failures that occur at destructive overspeed are caused by stresses which exceed the materials tensile strength.

In the ever}t of a turbine disc burst, high velocity missile-like fragments may .

break through the turbine casing, possibly generating secondary missiles.

These missiles have a potential of damaging reactor safety systems. l Alternately, the disc fragments could be arrested and contained by the turbine itself. Hence, in evaluating the risk associated with turbine disc rupture, it is necessary to determine whether or not missiles external to the casing can be generated by postulated disc ruptures.

This evaluation considers the above possibilities and summarizes the review and evaluation of the GE report, listed above, which describes GE procedures l

for estimating (1) the design speed missile generation probability, (2) the destructive overspeed missile generation probability, and (3) the perforation ef the turbine casing by turbine disc burst fragments.

• . g 3.0 DISCUSSION / EVALUATION This' discussion delefibes an overview of the methodology in Section 2 of the GE report where three major components of the methodology are considered:

- Probability of turbine overspeed

- Probability of wheel

• burst

. - Probability of casing penetration The probability of a wheel burst and the probability that a wheel fragment will penetrate the casing will de' pend on the speed at which a wheel bursts.

The turbine speed is close to 1800 RPM under normal operating conditions, however; when an abnormal event occurs, such as load rejection and/or failure of the control system to function properly, turbine speed may reach 180 to 190 percent of the rated speed. The probability of attaining these various turbine overspeed levels, therefore, is a major component of the methodology.

Another major component of the methodology is the probability of wheel burst at various operating conditions which are defined by two important parameters; the wheel speed and temperature. The primary failure mode of the turbine wheel is assumed to be brittle fracture due to the presence of a stress corrosion crack in the keyway near the bore of the shrunk-on wheel. The ,

fracture mechanics calculations include the variations in the toughness of the wheel material, in the depth of the crack, the variation in the likelihood of crack initiation, in the ability to detect crack sizes during inservice inspections and in the rate of crack growth during subsequent service.

The third major component of the methodology is the probability of wheel fragment penetrating the turbine casing, given the wheel burst at a particular speed. The missile penetration probabilities are based on energy methods (Ref. 8) and laboratory tests. The variations involved in these calculations lead to a probabilistic estimate of casing penetration as a function of burst speed.

• Turbine wheel and. turbine disk are synonymous.

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9 Section 3 of the GE report describes overspeed protection systems. GE nuclear steam turbines are equipped with three speed-sensing de' vices against turbine

overspeed

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(1) Normal overspeed protection is achieved through the' control, valves, ,

intercept valves and check valves; ,

(2) an energency overspeed protection device is set to close all steam  ;

valves if the speed reaches 110 to 111 percent of the operating level; and

! (3, a backup overspeed protection device is set to close all steam valves if the speed exceeds the emergency trip set point (112%).

Both mechanical hydraulic control (MHC) and electro-hydraulic control (EHC) l systems are employed. Failure model- for MHC and EHC systems are analyzed by a fault tree method and the probability of attaining a given speed is calculated.

Section 4 of the GE report considers wheel burst in both brittle and ductile modes. Operating experience shows that the primary failure mode of the turbine wheels is assumed to be brittle fracture due to the presence of stress .

corrosion cracks in the keyway near the bore of the shrunk-on wheel. After ascertaining the fracture toughness property at various depths, calculations 1

are made to determine crack length at a particular time from the initial service. Considerations in the probability analysis are given to variations l in the likelihood of crack initiation, in the ability.to detect and size l cracks during inservice inspections and in the rate of crack growth during ,

l subsequent service.

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The statistical distribution is applied to crack initiation and growth  ;

behavior data obtained from inservice inspections performed on the majority of ,

wheelsofoperatin&'GEnuclearlowpressureturbines. The relevant information can be extracted from.these statistical distributions to arrive at a given value for any assigned probability or vice versa. Because of various ,

parameters involved, the time to crack initiation varies significantTy from wheel to wheel.

Tests on wheels with laboratory produced stress corrosion cracks and those retired frcm service were utilized to define the ability of ultrasonic testing (UT) methods to detect and size wheel cracks. The GE analysis of the data shows that the crack depth is .07 inch longer on the average than the measured value for the wheel hub. The crack initiation distribution is influenced by the oxygen concentration in the steam and the type of locking ring that covers the keyways. The Weibull distribution was fitted to the observed field data and the characteristic life, a Weibull parameter, is approximated. Due to limitations of UT equipment, undetected cracks'might have initiated. These undetected cracks are considered and combined with i... o.erage crack growth rate from the initiation then the actual distribution of crack depths can be estimated.

The GE report synthesizes stress corrosion crack growth with fracture appearance transition temperature (FATT) and excess temperature versus the critical stress' intensity factor, Kg , and the calculated stress intensity factor, K3 , at.

various operating conditions. FATT is determined from the test results of retired wheels and other laboratory generated test data. The FATT value increases with' distance from the surface to the interior of the wheel. The ,

prediction of deep-seated (i.e. the wheel center) FATT values is based on the regression analysis which considers the range of three Nickel alloys. The distributions of points about the median are normal with a standard deviation of 28'F. The overall standard deviation is 35'F when the cooling rate, ultimate tensile strength, percent carbon, and percent Nickel error-distributions are considered.

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-The tcughness of the wheel material can be ascertained from toughness curves based on excess temperature (material temperature minus FATT) and the data generated from valf8' ASTM specimens. A semi-log relation fits the data below 100*F excess temperature. The data are more widely scattered at a lower excess temperature than at the hig'her values. Here, the natural logarithm of ,

standard deviation is a linear function of excess temperature. A log-normal relation is used for all upper shelf values utilizing the Rolf-Novak relation for all the GE shrunk-on wheel service data.

The stress intensity factor, K y, is determined from a relationship involving a crack shape factor, the stress, the crack depth and the geometry of the part

r. ear the crack. The general shape of stress corrosion cracks is assumed to be elliptical. The cracks are quarter-elliptical at corners and l semi-elliptical in the interior. An average aspect ratio of depth of crack (half the minor axis of the ellipse) to the half-length along the surface (half the major axis) is assumed to be about 0.4, based on a study of three i wheels that had several stress corrosion cracks. The average crack shape factor for corner cracks in the keyway under the hub was calculated to be 1.85. The average shape factor of 1.71 for semi-elliptical cracks under the web was calculated. The log standard deviation for both of these factors is

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taken to be 0.02. The corrosion crack branching factor distribution is l derived from test data on retired wheels and other data reported in the .

literatnre. The nominal bore stress is assumed to be log-norma) with a standard deviation of 0.02. The keyway geometric function is based on weight function method applied to the results of finite element analyses.

After obtaining the probability of a crack initiating at time t and knowing crack depth "a" at inspection time tt , and using Weibull distribution for growth rate, the probability of having a crack depth "a" at time t 2 regardless of when it initiates is obtained by multiplying these two probabilities and integrating over the range from zero to time t2 . Multiplying the probability of having a crack depth "a" at time tg with probability of not detecting a.

crack depth "a" and integrating the product from zero to infinity for all possible crack depths, the probability of missing a crack of any depth at t 2 is obtained. Thus, dividing the probability of missing a crack of depth "a"

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.t it: probability of missing a crack of any depth at time t3 will give the density function of any undetected cracks at time t . 2 Various combinations of temperatures and 13cking devices result in a median value of 0.03 inch with a log normal standard deviation of 0.24. This shows that there is a 50%

probability"of the undetected crack size being less than 0.03 inch. , .

The probability of cracks existing when no indication is found is computed fro- dividing the probability of missing any depth crack at time t3 by the sum of the probability of missing any depth crack and the probability of no crack existing at time 1 .3 This probability is the same as the probability of missing a crack at inspection time t . 3 The difference between the true crack initiation distribution and the observed crack initiation distribution is considered as a percentage difference for a given time. Thus, the percentage

'of cracked wheels is higher for the true crack initiation distribution than for the observed distribution.

The field data was adjusted based on (1) true crack initiation distribution, (2) Weibull distribution as a function of temperature, and the reciprocal to temperature, (3) reactor type, and (4) type of locking ring. The data then showed that the Weibull slope was close to unity and that the characteristic growth rate distribution for the third iteration remains indistinguishable from that of the first iteration. .

The probability of wheel burst at any time is a function of speed and temp- -

rature during a cycle between two refueling outages. The cumulative probability of burst increases in time since the last inspection. The probability of wheel burst at time t ,2P (tB).2can occur either at normal operation, P BN(t2 ), or at abnormal operation, P BA(t2). However, wheel burst at abnormal speed will occur only if there is no burst at a normal speed. The annual rate of oissile generation during normal operation is calculated by multiplying PBN(t2 )

with the probability of a missile given a burst at a normal speed. Probability l of burst for abnormal events is derived on l

a 13 "e assumption that a burst will not occur until the cumulative burst probability exceeds the level attained during normal operation. An . abnormal event occurs at a given temperature and a give'n maximum speed. This probability difference is added for all temperature levels for this abnormal ,

event. Further, summing the probabilities of all abnormal events givss the probability of generating a turbine missile external to the casing, P,(t2 ).

which depends on the speed at which the wheel bursts. Hence, the event (missile) must be integrated over the speed ranges for a given temperature.

This difference must be multiplied by the probability of speed and temperature occurring, and summed for all temperatures that can occur for the abnormal event. This probability must be again multiplied by the annual probability of an abnormal event occurring and summed for all possible abnormal events.

l Thus, the probability of a missile due to abnormal events is obtained. The

final missile generation probability P is the sum of the probability of a

missile due to normal and abnormal events.

J Tne second mode of failure is ductile fracture of the wheel during an .

abnormally high overspeed occurrence. Failure occurs when the average t c.;;"ial stress across the wheel section exceeds th'e tensile strength of the material. Since both brittle and ductile modes are statistically independent, a ecmbined probability of failure is determined.

Section 5 of the GE report discusses the values of casing escape probability of each shrunk-on wheel of GE nuclear turbines. Earlier analyses assumed that the energy absorption was due to a gross deformation of many components of the low pressure turbine casing. However, present tests show that the absorption is a local " punching" mechanism. EPRI full scale casing penetration tests

, consisted of accelerating a 120 degree segment of an actual turbine wheel at 180% speed of the turbine. Test results show that empirical formula are overly conservative (Ref. 9). The actual penetration of the missile is only half-way through the wall when a 8300 lb missile at 450 ft/sec strikes the wall. The range of final energy variation (energy remaining after absorption) -

is based on a normal distribution.

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14 Section 6 of the GE report gives an overall determination of a wheel burst probability which is a function of time, temperature, and speed. During a typicalnormalopeNiingcycle,thesteam/wheeltemperaturevariesfrom50*F at the start (0-speed) to 220*F (for wheels most susceptible to SCC) at full-loading (1800 RPM), then to 120*F after the coast down. The probabi]ity of' annual failure rate is calculated for both the normal and abnormal operating conditions. By combining these two probabilities, the probability of missile generation, P2 , is derived.

Section 7 of the GE report discusses typical results of calculations.

To provide further insight into the influence of various factors involved in the '

method, missile probability calculations have been made for a typical GE turbine used with a boiling water reactor. Two tables summarize the wheel .

t information for each of the 32 wheels used on the turbine. 'The median of ,

calculated deep-seated FATT is given for each wheel together with the reasured values of surface FATT and tensile strength. The type of locking ring used with axial key is also noted. Tables also describe the wheel temperature under full load conditions and the median of crack growth ,

rate which is calculated using this design temperature. The third table describes the results of missile probability calculations for each wheel of low pressure turbine at various times since the last inspection. Based on these calculations, the risk of missile generation for each rotor and the -

unit can be estimated. l

4.0 CONCLUSION

S AND RECOMMENDATIONS The methodology employed in the GE report for threalculation of wheel rupture and turbine missile generation probabilities is a straight forward application of probabilistic concepts to variation in surface FATT and deep-seated FATT, everspeed due to load rejection and/or failure of the control system to function properly. The fracture mechanics calculations include the i

statistical variations in (1) the toughness of the wheel material, (2) the depth of the crack, (3) the variation in the likelihood of crack initiation, (4) the ability to detect and to size cracks during inservice inspections, and (5) the rate of crack growth during subsequent service. Surface FATT can be measured, but since the deep-seated FATT values must be calculated from an cxperimental correlation, the overall standard deviation is a larger value

15 than the staff would anticipate. In this way conservatism is introduced at each step. The population of experimental tests and the actual data from the retired wheels is Nill small, and this results in a large standard deviation, thereby giving a conservative estimate. The staff finds that GE crack growth equatio'n gives a somewhat lower growth rate than that by another vendor; .

however, the allowable crack length is only one-half the critical crack length for the determination of an inspection interval.

In order to arrive at the final probability of a missile generation under normal and abnormal operating conditions, a series of numerical integrations are required and this may introduce some uncertainty. However, the missile penetration formula used is conservative (Ref. 10) in that disc fragments as heavy as 4600 lb at velocities as great as 300 mph penetrated less than half the thickness of walls at impact velocities that would have produced complete perforation according to the formulas.

The staff concludes that various safety factors or margins used in arriving at the final inspection interval are adequate and the report describes an acceptable method to determine such inspection intervals.

Therefore, the staff concludes that the present report may be used in determining the inspection interval for turbine wheels in operating and new

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raactor pla'nts. The inspection interval will vary from plant to plant based on -

the type of turbine in service and the previous inspection results. Applicants or licensees who wish to reference this report should commit to the turbine

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inspection intervals determined by GE and should submit a brief summary of how the GE method is used for their specific turbines. The latter should include graphs or tables of missile probability versus inspection interval.

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5.0 REFERENCES

1. S. H. Bush, " Probability of damage to Nuclear Components, " Nuclear Safety,14,3,'(May - June) 1973, p.187; and S. H. Bush, "A Reassess-ment of Turbine-Generator Failure Probability," Nuclear Safety, 19, 9 6, (Nov - Dec.) 1978, p. 681..'

2. NUREG/CR-1884, " observations and Comments on the Turbine Failurd at Yankee Atomic Electric ~ Company, Rowe, Massachusetts," March 1981.

3. Preliminary Notification of Event or Unusual Occurrence -- PNO - III -

81 - 104 - " Circle in the hub of the eleventh stage wheel in the main turbine" at Monticello Nuclear Power Station, Nov. 24, 1981.

4. Licensee Eveht Report No. 82.-132, Docket No. 50-361 - " failure of turbine stop valve 2VV-2200E to close fully" at San Onofre Nuclear Generating Station, Unit 2, Nov. 19, 1982.

5. . J. J. Burns, Jr. , " Reliability of Nuclear Power Plant Steam Turbine Overspeed Control Systems," 1977 ASME " Failure Prevention and Reli-ability Conference," Chicago, Illinois (Sept.) 1977, p. 27.

6. D. Kalderon, " Steam Turbine Failure at Hinkley Point A," Proc. Instn.

Mech. Engrs., 186, 31/72, 1972, p. 341. ,

7. W. G. Clark, Jr. , B. B. Seth, and D. H. Shaf fer. " Procedures for Estimating the Probability of Steam Turbine Disc Rupture from Stress Corrosion Cracking," ASME/IEEE Power Generation Conference Oct. 4-8, 1981, St. Louis, Missouri.

8. Gonyea, D. C., "An Analysis of the Energy of Hypothetical Wheel Missiles Escaping from Turbine Casings," General Electric Company - Turbine Department Report, February 1973. *

9. S. McHugh, L. Seaman and Y. Gupta, " Scale Modeling of Turbine Missile 4 Impact into Concrete," EPRI Report NP-2746, February 1983. I

10. R. L. Woodfin, " Full-scale Turbine Missile jconcrete Impact Experiments," i EPRI report NP-2745, February 1983. ,

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Mr. H. B. Tucker Duke Power Company Catawba Nuclear Station

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A.V. Carr, Esq. North Carolina Electric Membership j Duke Power Company Corp. .

3 422 South Church Street 3400 Sumner Boulevard Charlotte, North Carolina 28242 P.O. Box 27306 Raleigh, North Carolina 27611 J. Michael McGarry, III, Esq.

Bishop. Liberman, Cook, Purcell Saluda River Electric Cooperative, 1 and Reynolds Inc.

1200 Seventeenth Street, N.W. P.O. Box 929 Washington, D. C. 20036 Laurens, South Carolina 29360 florth Carolina MPA-1 Senior Resident Inspector

. Suite 600 Route 2 Box 179N r

3100 Smoketree Ct. York, South Carolina 29745 P.O. Box 29513 Raleigh, North Carolina 27626-0513 Regional Administrator, Region II

{

U.S. Nuclear Regulatory Commission, L.L. Williams 101 Marietta Street, NW, Suite 2900 l Area Manager, Mid-South Area Atlanta, Georgia -30323 i ESSD Projects

Westinghouse Electric Corp.

l P9fC West Tower - Bay 239 P.O. Box 355 Pittsburgh, Pennsylvania 15230 Mr. Heyward G. Shealy, Chief Bureau of Radiological Health South Carolina Department of Health and Environmental Control 2600 Bull Street ..

Columbia, South Carolina 29201 s

County Manager of York County York County Courthouse Karen E. Long York South Carolina 29745 Assistant Attorney General N.C. Department of Justice Richrd P. Wilson, Esq. P.O. Box 629 Assistant Attorney General Raleigh, North Carolina 27602 S.C. Attorney General's Office i P.O. Box 11549 Spence Perry, Esquire l Columbia, South Carolina 29211 General Counsel  !

Federal Emergency Management Agency 1 Piedmont Municipal Power Agency Room 840 100 Memorial Drive 500 C Street <

Greer, South Carolina 29651 Washington, D. C. 20472

! Mr. Michael Hirsch l Federal Emergency Management Agency Office of the General Counsel Room 840 500 C Street, S.W.

Washington, D. C. 20472 Brian P. Cassidy, Regional Counsel Federal Emergency Management Agency, Region I J. W. McCormach POCH

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