MFN-09-484 Enclosure 2 GE-ST ESBWR Steam Turbine - Low ... · Fracture Toughness -ASTM toughness tests of NiCrMoV material is the basis for the critical stress intensity versus excess

MFN-09-484

Enclosure 2

GE-ST "ESBWR Steam Turbine - Low Pressure

Rotor Missile Generation Probability Anlaysis"

ST-56834/N-P, Revision 1

Public Version

I NON-PROPRIETARY VERSION I

Economically Simplified Boiling Water ReactorSteam Turbine - Low Pressure Rotor

Missile Generation Probability Analysis

ST-56834/N-P, Revision 1Dated - June 17, 2009

General Electric Company. @ General Electric, 2009

ESBWR Missile Probability Analysis Pa~ge 2 of 56

Non-Proprietary Version - Information denoted by a R[- -.. A)]],is information considered proprietary to General Electric andhas been deleted from this report.


ESBWR Missile Probability Analysis Page 3 of 56

TABLE OF CONTENTS

1. SU M M A RY A N D O V ERV IEW ..................................................................... 6

1.1 RESULTS ......................................................................................................... 6

1.2 METHODOLOGY ................................................................................................. 7

1.2.1 Probability of Turbine Overspeed ............................................................................. 7

1.2.2 Rotor Burst Probability ............................................................................................... 7

1.2.3 Probability of Casing Penetration ............................................................................. 8

1.2.4 Key Param eters of Num erical Evaluation ................................................................ 8

2. TURBINE-GENERATOR DESCRIPTION ............................................... 9

2.1 GENERAL DESCRIPTION ...................................................................................... 9

2.1.1 Main Stop and Control Valves ................................................................................... 9

2.1.2 Com bined Interm ediate Stop and Intercept Valves ............................................. 10

2.1.3 Extraction Non-return Valves .................................................................................. 10

2.2 TURBINE OVERSPEED PROTECTION SYSTEM ..................................................... 10

2.3 TURBINE PROTECTION SYSTEM ......................................................................... 12

2.4 TESTING ........................................................................................................... 13

3. TU R B IN E INTEG R ITY ................................................................................ 15

3.1 MATERIALS SELECTION .................................................................................... 15

3.1.1 Fracture Toughness .................................................................................................. 15

3.1.2 High Tem perature Properties .................................................................................. 17

3.1.3 Pre-Service Inspection and Testing ........................................................................ 17

General Electric Company.

ESBWR.Missile Probability Analysis Page 4 of 56

4. TURBINE GENERATION MISSILE PROBABILITY ........................... 18

4.1 FRACTURE MECHANICS .................................................................................... 18

4.2 CYCLIC PROPAGATION OF UNDETECTED FORGING DEFECTS ............................ 19

4.2.1 Cyclic Crack Growth .................................................................................................. 19

4.2.2 Undetected Flaw Size ................................................................................................ 20

4.2.3 Cyclic Profile .............................................................................................................. 20

4.3 STRESS CORROSION CRACKING .............................................................. 21

4.4 DUCTILE ROTOR BURST .................................................................................... 23

5. O V ERSPEED PRO BA B ILITY .................................................................. 24

5.1 PROBABILITY CALCULATION ............................................................................. 24

5.1.1 Steam Valve Arrangem ent ....................................................................................... 25

5.1.2 Steam Valve Model .................................................................................................... 255 .1 .2 .1 V a lv e fa ilu re ra te s ....................................................................................................................... 2 7

5.1.3 Extraction system s ................................................................................................... 28

5.1.4 Hydraulic Model ........................................................................................................ 28

5.1.5 Emergency Trip System ............................................................................................ 30

5.1.6 Turbine Generator Control System (MARKTm Vie) .................................................. 31

6. CA SING PEN ETRA TIO N .......................................................................... 34

6.1 COMPONENTS .................................................................................................. 34

6.2 CASING PENETRATION CALCULATIONS ............................................................ 35

7. OVERALL PROBABILITY DETERMINATION .................................... 37

7.1 PROBABILITY OF BRITTLE FRACTURE (KI > Kic) .................................................. 37


ES13WR Missile Probability Analysis Page 5 of 56

7.2 PROBABILITY OF DUCTILE TENSILE FAILURE ..................................................... 42

7.3 NO RM AL O PERATIO N ....................................................................................... 42

7.4 ABNO RMAL O PERATION ................................................................................... 44

8. BORELESS (SOLID) ROTOR RESULTS SUMMARY ...................... 48

9. BORED ROTOR RESULTS SUMMARY ............................................... 50

10. IN-SERVICE INSPECTIO NS .................................................................. 52

10.1 IN-SERVICE MAINTENANCE AND INSPECTION OF TURBINE ROTORS52

10.1.1 Rotor Dovetail Inspections ..................................................................................... 52

10.1.2 In-service Inspection of Turbine Valves ............................................................... 52

R E F E R E N C E S ................................................................................................. 55



1. SUMMARY AND OVERVIEWThe evaluation of turbine missile effects is commonly characterized by the followingequation:

P4 = Pi x P2 x P3

Where P4 = annual probability of unacceptable damage resulting from aturbine missile

P1 = annual probability of turbine failure resulting in the ejection ofturbine rotor (or internal structure) fragments through the turbinecasing.

P2 = the probability that a turbine missile strikes a critical plant target,given generation.

P3 = the probability that the critical target is unacceptably damaged,given a missile strike.

The NRC licensing guidelines (Regulatory Guide 1.115 and NUREG-1048) use thisformulation to describe hypothetical turbine missiles and specifies that the probabilityof unacceptable damage from turbine missiles should be less than or equal to 1 in 10million per year (i.e., P4 should be


limiting missile generating mode becomes Stress Corrosion Cracking (SCC). Asdescribed in Section 8 and 9, SCC characteristics (initiation and growth) assumed in thisreport are based on historical field experience of older LP rotors featuring shrunk-onwheels that exhibited bore keyway SCC. The use of this data is considered to beconservative when applied to the ESBWR LP monoblock rotors as discussed later in thisreport.

General Electric has a long established practice of recommending periodic in-serviceinspection of turbine components as part of maintenance programs. The basis of themissile probability analysis assumes that all General Electric maintenancerecommendations and operating practices are satisfied and defines that the ESBWRturbine components should be inspected at frequency not to exceed 12 years. Theresults of the in-service inspections will be used to update unit specific missileprobability estimates and to adjust the re-inspection frequency to a shorter interval, ifconsidered necessary based on observed results.

1.2 METHODOLOGY

The methodology of missile generation probability analysis deals with one element ofthe overall missile issue, which is the probability (Pi) of generating a turbine missilefrom the LP turbine external to the LP inner casing and LP hood structure. The.methodology for determination of missile generation probability contains three majorcomponents:

* The probability of the turbine attaining speeds higher than those occurringduring normal operation (overspeed),

* The estimation of rotor burst probability as a function of speed, and• The probability of a rotor fragment penetrating the turbine casing and thus

generating an external missile.

1.2.1 Probability of Turbine Overspeed

The probability of a rotor burst and the probability that a fragment will penetrate theturbine casing are both dependent on the speed at which rotor burst is assumed tooccur. Under normal operating conditions, the turbine speed is close to the ratedspeed (1800 rpm). When an abnormal event occurs, such as a full load rejection andfailure of elements of the control system, turbine speeds significantly higher than therated speed may occur. A major component of the analysis is to estimate theprobability of attaining various overspeed levels.

1.2.2 Rotor Burst Probability

One rotor failure mode considered is brittle burst, specifically as the result of a cracklocated in the radial-axial plane growing to a critical size. Brittle burst scenariosaddressed are: 1) an undetected internal forging flaw that grows cyclically to criticalsize, and 2) a time dependent SCC that initiates on the outer body surface and grows toa critical size. A second failure mode due to tensile failure is also included in themethodology. This ductile failure mode contributes to the rotor burst probabilityparticularly during abnormally high overspeed occurrences.



1.2.3 Probability of Casing Penetration

The third major component of the missile probability analysis methodology deals withthe probability of a rotor burst fragment penetrating the turbine casing. Calculationsemploy the energy analysis method already developed by General Electric, describedin Reference 1, Section 5.

This method considers the kinetic energy of the assumed fragment at the instant ofburst as well as the energy absorbing capability of the stationary components of thelow-pressure turbine.

1.2.4 Key Parameters of Numerical Evaluation

Undetected Flaw Size - General Electric nuclear LP rotor forgings are manufactured bythe highest quality suppliers and are subjected to both volumetric and finish surfacenon-destructive inspection using methods and techniques that are well established inthe nuclear power industry. Assumptions about possible undetected flaws areconsistent with industry practice and the possible undetected flaws are conservativelyassumed to be cracks.

CVclic Crack Growth - Cyclic crack growth behavior was determined from GeneralElectric NiCrMoV specimen tests. For this analysis, a finite number of operationalcycles per year is assumed.

SCC Crack Initiation and Growth - Historical experience has shown that SCC can occurin NiCrMoV at locations featuring tensile stress and in contact with wet steam. Thismissile generation probability analysis considers the possibility that an SCC crackinitiates on the outer rotor surface, specifically at bucket attachment dovetail stressconcentrations, and grows radially inward to a critical size. Statistical assumptionsabout both SCC initiation time and SCC growth rate are based on General Electricmeasurements of stress corrosion cracks found in keyways of shrunk-on wheels(Reference 1, Appendices D and E).

Fracture Toughness -ASTM toughness tests of NiCrMoV material is the basis for thecritical stress intensity versus excess temperature correlation used in the analysis.These results are from tests conducted by General Electric and others.

Rotor Operating Conditions - Rotor stresses are derived from finite element analysis.Both mechanical (rotational) and thermal stresses are considered. Temperatures arederived from the overall unit heat balance.

Turbine Control System Reliability - The failure rates of various turbine controls systemcomponents are estimated from actual field experience, laboratory studies, andcommercial data.

Casing Penetration Behavior - The missile penetration probabilities used in the analysisare based on the published energy method developed by General Electric. Thismethod has been validated by full-scale tests, beginning in 1969 and the method isdiscussed in greater detail in Reference 1, Section 5.



2. TURBINE-GENERATOR DESCRIPTION

2.1 GENERAL DESCRIPTION

The turbine generator (TG) consists of an 1800-rpm turbine, external moistureseparator/reheaters, generator, static excitation system, controls, and associatedsubsystems. The turbine for the ESBWR standard plant consists of a double-flow, high-pressure turbine, and three double-flow low-pressure turbines in tandem, each with52-inch lost stage buckets.

The high-pressure turbine has extraction points for reheating steam and high-pressurefeedwater heating. Moisture separation and reheating of the high-pressure turbineexhaust steam is performed by external Moisture Separator Reheaters (MSRs). TwoMSRs are located on each side of the TG centerline. The steam then passes throughthe three low-pressure turbines, each with extraction points for the low pressurestages of feedwater heating, and exhausts into the main condensers. In addition tothe moisture separators in the external MSRs, the turbine steam path has provisionsfor removing some additional moisture and routing it to extraction lines.

The generator is a direct driven, three-phase, 60 Hz, 1800 rpm synchronous generatorwith a water-cooled armature winding and hydrogen-cooled rotor.

The TG uses a General Electric MARK TM IVe' Turbine Generator Control System (TGCS),which is a digital monitoring and control system. The TGCS, in coordination with theSteam Bypass and Pressure Control (SB&PC) system, controls the turbine speed, load,and flow for startup and normal operations. The control system operates the turbinestop valves, control valves, and combined intermediate stop and intercept valves. TGsupervisory instrumentation is provided for operational analysis and malfunctiondiagnosis.

TG accessories include the bearing lubrication oil system, turbine hydraulic system,turning gear, hydrogen gas control system, seal oil system, stator cooling watersystem, exhaust hood spray system, turbine gland seal system, excitation system, andturbine supervisory instrument system.

2.1.1 Main Stop and Control Valves

Four main stop and four control valves admit steam to the high-pressure turbine. Theprimary function of the main stop valves is to quickly shut off the steam flow to theturbine under trip conditions. The primary function of the control valves is to controlsteam flow to the turbine in response to the TGCS.

The main stop valves are hydraulically operated in an open-closed mode either by theturbine overspeed protection system in response to a turbine trip signal, or by a testsolenoid valve and a fast acting solenoid valve for periodic testing. The disks areunbalanced and cannot open against full differential pressure. A bypass is provided topressurize the below seat areas of the four valves and supply steam for turbine casingand steam chest warming. Springs in the valves are designed to improve the closing

The MARK turbine controls system is a trademark of General Electric.



time response of the main stop valve under the abnormal conditions. An equalizingheader is provided between the stop valves, upstream of the control valves. Eachmain stop valve is designed to accept a steam strainer to limit foreign material fromentering the control valves and turbine.

The control valves are designed to provide steam shutoff adequate for turbine speedcontrol. The valves are of sufficient size, relative to their cracking pressure, to require apartial balancing. Each control valve is hydraulically operated by high-pressurefire-resistant fluid supplied through a servo valve.

2.1.2 Combined Intermediate Stop and Intercept Valves

Hydraulically operated combined intermediate stop and intercept valves are providedin each hot reheat line just upstream of the inlet to each of the LP turbine inlets (sixtotal). The combined intermediate and intercept valves control steam flow to each ofthe LP turbines in response to the TGCS. Each combined valve includes twoindependently operated valve discs in series for open-closed operation.

2.1.3 Extraction Non-return Valves

Upon loss of load, the steam contained downstream of the turbine extractions canflow back into the turbine, across the remaining turbine stages, and into thecondenser. Associated condensate can flash to steam under this condition andcontribute to the backflow of steam or can be entrained with the steam flow anddamage the turbines. Non-return valves are employed in selected extraction lines tominimize potential for these conditions to contribute to the turbine overspeed.

2.2 TURBINE OVERSPEED PROTECTION SYSTEM

The normal speed control system comprises the first line of defense against turbineoverspeed. This system includes the main control valves, intercept valves, andfast-acting valve-closing functions within the TGCS. The normal speed control unitutilizes three speed signals. Loss of any two of these speed signals initiates a turbinetrip via the Emergency Trip System (ETS). Under normal operation, an increase inspeed above the speed setpoint initiates the closing of the control and intercept valvesin proportion to the speed increase. Rapid turbine acceleration resulting from asudden loss of load at higher power levels initiates the fast-acting solenoids via thenormal speed control system. The fast-acting solenoids rapidly close the main controland intercept valves irrespective of the current turbine speed.

The normal speed control system is designed to limit peak overspeed resulting from aloss of full load, to at least 1% below the overspeed trip setpoint. Typically, this peakspeed is in a range of 106-109% of rated speed, and the overspeed trip setpoint isapproximately 110% of rated speed. All turbine steam stop, control and interceptvalves are fully testable during normal operation. The fast closing feature, provided byaction of the fast-acting solenoids, is testable during normal operation.

Normal speed control is supplemented by the power load unbalance function. Thepower load unbalance function can protect the turbine from an overspeed trip



condition in the event of full load rejection. The power load unbalance function looksfor an unbalance between mechanical power and electrical load. Under specific loadrejection conditions, the power load unbalance will initiate main control valve andintercept valve fast closing functions to prevent rapid acceleration and a subsequentturbine trip.

If the normal speed control and power load unbalance function should fail, theoverspeed trip devices close the main stop and control valves, and the combinedintercept and intermediate stop valves. This turbine overspeed protection systemcomprises the second line of defense against turbine overspeed. It is both redundantand diverse.

Redundancy comes from the use of multiple speed probes, multiple controllers, andmultiple trip solenoid valves. The turbine hydraulic trip solenoid valve hydraulic circuitsare arranged in a dual, "two-out-of-three", de-energize to trip configuration. Anypower interruption to either set of the two-out-of-three trip solenoid valves in theEmergency Trip Device (ETD) results in a turbine trip.

Diversity is provided by separate sets of physically isolated primary and emergencyoverspeed protection controllers. The primary overspeed trip and emergencyoverspeed trip controllers are independent and diverse by providing unique hardwareand logic design and implementation. Power to the trip solenoids is interrupted byeither the primary overspeed protection controllers or by the emergency overspeedprotection controllers. An overspeed trip results if either set of redundant controllersdetermines an overspeed condition exists. Power interruption to the turbine controlcabinet (which also supplies power to the trip solenoids) results in a "fail-safe" turbinetrip. The trip solenoid valve and associated controller are fully testable during normaloperation.

For an actual overspeed trip condition, the primary overspeed controllers exchangeand vote their individual speed inputs so each controller executes its protectivealgorithm on the consensus speed value. Each primary overspeed controllerde-energizes trip solenoid valves in a two-out-of-three logic arrangement. Thetwo-out-of-three logic precludes a single failure in any of the three controllers fromblocking trip initiation.

A different implementation and operation takes place in the three completely separateand individual emergency overspeed trip controllers. Each of the three emergencycontrollers has a dedicated power supply and operates completely separate from eachof the other emergency overspeed trip controllers. The three emergency controllersoperate independently from the primary overspeed trip controllers. In the event of anoverspeed condition, the emergency controllers individually detect and determinespeed, and de-energize trip solenoid valves in a two-out-of-three logic arrangement.

The overspeed protection system is designed to ensure that failure of the normalspeed control system does not result in turbine speed exceeding -120% of ratedspeed. The components and circuits comprising the turbine overspeed protectionsystem are testable when the turbine is in operation.



The overspeed sensing devices are located in the turbine front bearing standard, andare therefore protected from the effects of missiles or pipe breakage. The hydrauliclines are fail-safe; if one is broken, loss of hydraulic pressure results in a turbine trip.The ETDs are also fail-safe. Each trip solenoid transfers to the trip state on a loss ofcontrol power, resulting in a turbine trip. These features provide inherent protectionagainst failure of the overspeed protection system caused by low trajectory missiles orpostulated piping failures.

Each turbine extraction line is reviewed for potential energy and contribution tooverspeed. The number and type of extraction non-return valves required for eachextraction line are specified based on the enthalpy and mass of steam and water inthe extraction line and feedwater heater. Higher energy lines are provided withpower-assisted open, spring-assisted closed non-return valves, controlled by air relaydump valves, which in turn, are activated by the ETS. The air relay dump valves,actuated on a turbine trip, dump air from the extraction non-return valve actuators toprovide rapid closing via actuator spring force. The closing time of the extractionnon-return valves is sufficient to minimize extraction steam contribution to the turbineoverspeed event.

The following component redundancies are employed to guard against excessiveoverspeed:

(1) Main stop valves/Control valves,

(2) Intermediate stop valves/Intercept valves,

(3) Normal speed control/Primary overspeed trip/Power Load unbalance/Emergencyoverspeed trip,

(4) Fast-acting solenoid valves/Emergency trip fluid system (ETS),

(5) Extraction non-return valves (as needed).

The main stop valves and control valves provide full redundancy in that these valvesare in series and have independent control signals and operating mechanisms.Closure of all four-stop valves or all four-control valves shuts off all main steam flow tothe high-pressure turbine. The intermediate stop and intercept valves are also in seriesand have independent control signals and operating mechanisms. Closure of eithervalve or both valves in each of the six sets of intermediate stop and intercept valveseffectively shuts off intermediate steam flow to the three low pressure turbines. Thisarrangement is such that failure of a single valve to close does not result in turbinespeed exceeding -120% of rated speed.

2.3 TURBINE PROTECTION SYSTEM

In addition to the overspeed trip signals discussed, the ETS closes the main stop andcontrol valves and the intermediate stop and intercept valves to shut down the turbineon the following signals.

* Emergency trip in control room,

* Moisture Separator high level,


ESBWR Missile Probability Analysis Pa]ge 13 of 56

* High condenser pressure,

* Low lube oil pressure,

* Low pressure turbine exhaust hood high temperature,

* High reactor water level,

* Thrust bearing wear,

* Emergency trip at front standard,

* Loss of stator coolant (if runback fails),

* Low hydraulic fluid pressure,

* Selected generator trips,

* Loss of TGCS electrical power,

* Excessive turbine shaft vibration,

* Loss of two speed signals - either two Normal Speed Control or two Emergency,

* Loss of two or more SB&PC System channels,

* Closure of Main Steam Isolation Valves,

* Differential and/or Rotor Expansion.

When the ETS is activated, it overrides all operating signals and trips (closes) the mainstop and control valves, and combined intermediate stop and intercept valves.

2.4 TESTING

The primary and emergency overspeed trip circuits and devices are tested remotely ator above rated speed by means of controls in the Main Control Room (MCR). Operationof the overspeed protection devices under controlled speed conditions is checked atinitial turbine startup and after each refueling or major maintenance outage,consistent with General Electric recommendations to perform off-line (actual)overspeed testing every 6 to 24 months. In some cases, operation of the overspeedprotection devices can be tested just prior to shutdown. This eliminates the need totest overspeed protection devices during the subsequent startup if no maintenance isperformed that affects the overspeed trip circuits and devices.

Main stop, main control, intermediate stop, and intercept valves are exercised atintervals as required by the turbine missile probability analysis (120-days) by closingeach valve and observing the remote valve position indicator for fully closedposition status. This test also verifies operation of the fast close function of each mainstop and main control valve during the last few percent of valve stem travel. Fastclosure of the intermediate stop and intercept valves is tested in a similar way if theyare required to have a fast close function that is different from the test exercise.

Access to required areas outside of the turbine shielding is provided on the turbinefloor under operating conditions.



Provisions are included for testing each of the following devices while the unit isoperating:

M Main stop valves and main control valves,

* Low pressure turbine intermediate stop and intercept valves,

* Testable Turbine Extraction non-return valves important to overspeedprotection,

* Lubricating oil pumps,

* Hydraulic fluid pumps,

* Emergency Trip Device,

* Power-Load Unbalance circuits,

* Other test/inspections identified by the periodic operational test summary thatimpact reliability of the overspeed protection system.



3. TURBINE INTEGRITY

3.1 MATERIALS SELECTION

LP turbine rotors forgings are NiCrMoV alloy material in accordance with GeneralElectric specification B50A373B8. The material properties of the rotor forgings areoptimized to ensure excellent fracture toughness. Highly refined manufacturingprocesses are designed to minimize melt related defects and impurities. Undesirableelements, such as sulfur and phosphorus, are controlled to the lowest practicalconcentrations consistent with good melting practice. Rotors have the lowest FractureAppearance Transition Temperatures (FATT) and highest Charpy V-notch energiesobtainable, on a consistent basis from material at the sizes and strength levels used.The FATT temperature, as obtained from Charpy tests performed in accordance withASTM A-370, is no higher than -1.10C (300 F) for a large integral forging. The CharpyV-notch energy at the minimum operating temperature is at least 6.23 kg-m (45 ft-lbf)for a large integral rotor forging.

3.1.1 Fracture Toughness

The fracture toughness of a material is measured by the critical stress intensity factorKic. The toughness of NiCrMoV material has been found to vary with temperature. Atlower temperatures, lower toughness occurs and at high temperatures, an increase intoughness occurs until a maximum is reached. The maximum value is commonlyreferred to as the upper shelf. The fracture appearance varies from a brittleappearance at low temperature to a ductile appearance at high temperature. FATT isthe ductile-to-brittle transition temperature at which the fracture surface appearanceis 50% plastically deformed indicating ductile failure and 50% cleaved indicating brittlefailure.

Fracture toughness has been correlated to the measured excess temperature (metaltemperature minus the FATT). As a result, General Electric does not require fracturetoughness testing on production rotors. Deep seated FATT testing, however, will beperformed for each production rotor forging during routine material acceptancetesting and the fracture toughness at the forging centerline will be derived based onhistorical correlations. For missile generation probability calculations, a normallydistributed FATT featuring a -30°F mean and a 30°F standard deviation is assumed.

The correlation between critical stress intensity Kic, and FATT is shown in Figure 3-1.The fracture toughness is approximated by two independent mathematicalrepresentations, one for the low excess temperature region and another representingthe upper shelf of fracture toughness in the high excess temperature region. For thelow excess temperature region, the best fit through the center of the data is the semi-log expression:

[[ .(eq. 3-1)



where:

X =excess temperature

And excess temperature is measured in degrees Fahrenheit

The standard deviation is calculated according to the equation:

[[ ----- {A)ll (eq.3-2)

where:

In= natural logarithm

a, = standard deviation

X = excess temperature

For the upper shelf region, a Rolfe -Novak relationship (Reference 2) has been appliedto estimate fracture toughness. As shown in Figure 3-1, a log mean value of -[-------{A}Jlsquare root inch and log standard deviation = [[ .... {AII are used.

For missile analysis calculations, the two regions (upper and lower shelf) are treated asstatistically independent. The probability of burst is calculated for each region and thetwo probabilities are combined.

II

Figure 3-1 NiCrMoV Toughness Curve



3.1.2 High Temperature Properties

The operating temperature range of the low-pressure rotors is below the stress rupturetemperature range of the materials used. Therefore, creep-rupture is not consideredto be a failure mechanism for these components.

3.1.3 Pre-Service Inspection and Testing

The pre-service inspection procedures and acceptance criteria specific to the rotatingparts of the steam turbine are as follows:

" Rotor forgings undergo 100% volumetric (ultrasonic) inspection subject toestablished inspection methods and acceptance criteria that are equivalent toor more restrictive than those specified for Class 1 components in ASME CodeSections III and V. Subsurface sonic indications are not accepted if found tocompromise the integrity of the unit.

" The entire finish machined outer rotor periphery including the bore surface (ifpresent) is subjected to magnetic particle test or liquid penetrant examination.Surface indications are evaluated and removed if found to compromise theintegrity of the unit during its service life.

" Each fully bladed turbine rotor assembly is factory spin-tested at 120% of ratedspeed.



4. TURBINE GENERATION MISSILE PROBABILITY

The determination of the probability of the generation of a missile is based onprobability of overspeed (PA), the probability of rotor burst (PB), and probability of casingpenetration (Pc):

PI = PAX PB X Pc (eq. 4-1)where:

PA = Probability of achieving a speed of concern or a particular speed {Brittle Burstcould happen at any speed. For cases of rated speed or below PA = 1.0} (Section 5)

PB = Probability of rotor burst (Section 4)Pc = Probability of casing penetration (Section 6)

The methodology described in this report is for determination of P1, the probability ofturbine missile generation external to the turbine casing. Earlier work (Reference 1)focused on shrunk-on-wheels as the critical source of missiles. Independentcalculations of Pi for each wheel were combined to obtain an overall value of P1 foreach LP rotor. In this analysis report, the integral ESBWR rotor is also assumed to bemade-up of independent disks and an overall rotor missile probability obtained bycombining individual disk probabilities. The NRC annual missile probability limit of1x10-5 is assumed to apply to the overall unit.Rotor burst is assumed to occur when a crack oriented in the radial-axial planereaches a critical final size. Missiles occur when rotor burst fragments (of sufficientmass and energy) penetrate both the inner casing and outer hood. Three rotor bodyburst scenarios are considered:

1. Brittle fracture resulting from cyclic fatigue propagation of undetectedinternal forging defects

2. Brittle fracture resulting from Stress Corrosion Cracking (SCC); specificallyexternal axial entry dovetail surface cracks that initiate and grow radiallyinward towards the center of the rotor

3. Tensile rupture attributed to gross overspeed

4.1 FRACTURE MECHANICS

From Linear Elastic Fracture Mechanics (LEFM), the crack tip stress intensity factor K1 isdefined as follows:

K1 = Cx a-x ffx-a (eq. 4-2)

where:

C = crack shape factor

or = applied tangential stress

a = characteristic crack depth



Brittle Fracture occurs when crack tip stress intensity exceeds the material fracturetoughness Kc. Stated mathematically, burst occurs when: K, _> K1 c.

4.2 CYCLIC PROPAGATION OF UNDETECTED FORGING DEFECTS

The analysis focuses on the rotor bore. Bore 2 tangential stress is due to bothmechanical (rotation) and thermal loading. Radial thermal gradients encounteredduring a start up result in tensile thermal stresses having greatest magnitude at thebore surface. The maximum value of mechanical tangential stress (due to rotation) isalso found at the bore.

Bored rotor tangential stress, derived using Finite Element Analysis (FEA), issummarized in Table 4-1. Mechanical stresses are evaluated at 1800 rpm (ratedspeed). Thermal stresses are dependent on the time rate of change in steamconditions leaving the Moisture Separator Reheater (MSR) during a start-up. Valuesshown represent a worst-case MSR temperature ramp rate developed for the starting& loading of the ESBWR steam turbine that will be used in the development of thecontrols logic. For probabilistic calculations, values shown in Table 4-1 are assumed tobe log normally distributed with a 0.05 log normal standard deviation.

Table 4-1 ESBWR Bore Stress1800 RPM Bore Maximum Bore

Mechanical ThermalStage Stress (ksi) Stress (ksi)

1 II ---.-..

2 ---__ _---

3 ---.-.-

- 4 --- ---__ __ __

5 ---__ _---

6 ---_ _ _ ---_ _ _ _7 .... - -{[ }11

*Bore stress for a solid (boreless) rotor is 50% of the above values

4.2.1 Cyclic Crack Growth

Cyclic crack growth testing ofNiCrMoV by General Electric is represented using thetraditional Paris form:

da/dN = Bx(AK)n (eq. 4-3)

where:

2 Both bored and solid rotors can be treated with this analysis. Bored rotors incur larger stress and have less

uncertainty regarding FATT and crack detection. Solid rotors are discussed in greater detail in Section 9.



AK = cyclic range in applied stress intensity

Variability in growth rate is captured by variation in the coefficient B& A log normallydistributed coefficient (B) featuring an average value of [[ --- A-]land standarddeviation of [[ ----- A}]J along with an exponent (n) value of [[ .----- ( A}], bestrepresents the test data.

4.2.2 Undetected Flaw Size

Bore surfaces are subjected to both magnetic particle and ultrasonic inspection.Undetected bore surface crack size is a function of both measurement sensitivity andrepeatability. For analysis of the bored rotor, an average undetected bore surfacecrack is assumed to be semi-circular in shape with an average size "a" (i.e., radius) = [[ -

----]] and crack tip stress intensity shape factor C = 0.73 (Reference 3). For asolid rotor, an average undetected embedded crack is assumed to be semi-circular inshape with an average size "a" = [[ -------- {A] and crack intensity shape factor of C= 0.64 (Reference 4). Undetected cracks are assumed to be log normally distributedwith a log normal standard deviation of [[ .------- W} for size (a). The crack shape factoris. also assumed to be log normally distributed with a log normal standard deviation of

-..{A]

4.2.3 Cyclic Profile

Because missile probability is reported on an annual basis, crack growth must becalculated on an annual basis. The annual cyclic loading profile assumed for thisanalysis is shown in Table 4-2.

During each start/stop cycle, the maximum value of tensile stress (peak value of thecyclic applied stress range) at the bore is assumed equal to thesum of the mechanicaland thermal stresses shown in Table 4-1. By comparison, load swings are limited to a50% drop in power during which speed is unchanged and MSR exit temperature ismaintained very close to rated condition. Finite element simulation of the load swingrevealed


4.3 STRESS CORROSION CRACKING

SCC requires the combination of a corrosive environment, a susceptible material, andtensile stress. The SCC susceptibility of NiCrMoV shrunk on wheel rotor designs in awet steam environment is well documented (Reference 1). SCC burst scenarios for theESBWR rotor consider outer surface tangential stress concentrations as potential SCCcrack initiation sites. Burst is assumed to occur when a crack growing radially inward(towards the center of the rotor) reaches a critical size.

As shown in Figure 4-1, ESBWR LP rotors feature both tangential entry dovetails(Stages 1 through 4) and axial entry dovetails (Stages 5 through 7).

if_ !

Stages 1-4Stages 5, 6 and 7

Figure 4-1 LP Rotor Dovetail Configurations

Prior 1960s and 1970s era BWR rotors also featured tangential entry dovetails. SCCcracks found in the wheel dovetail hook fillet radii of prior tangential entry designshave been confined to the radial circumferential and axial circumferential planes. Assuch, tangential entry dovetail SCC cracks are considered to be a maintenance issuerather than a rotor burst risk.

Cracks forming in the slot bottoms of axial entry dovetails (stages 5, 6, and 7) andoriented in the radial axial plane are considered a rotor burst risk. Shrunk-on-wheelkeyway SCC crack statistical behavior including time to initiation and growth rate(summarized below) is assumed. Because ESBWR axial entry dovetail slot bottomsfeature dramatically lower tangential stress (vs. shrunk-on-wheel keyways), use ofshrunk-on-wheel initiation and growth characteristics is considered conservative.

A summary of the key statistical distributions derived from analysis of cracked nuclearshrunk-on-wheel keyways (Reference 1) and featured in this study include:


ESBWR Missile Probability Analysis

Dovetail Crack Initiation Behavior:

Page 22 of 56

F1 (t) = 1 - e-(l/)k (Weibull distribution)

where:F, (t) = cumulative probability by time t

t = time in years

A = Weibull scale parameter for initiation defined as:

(eq. 4-4)

[[ ---------------------------------------------------------- (A)]]

exp(o = base of the natural logarithm raised to power of the value in brackets

T= dovetail temperature during normal operation measured in 'F .

Dovetail Crack Growth:

FG (G) = 1 - e-(GAG)k (Weibull Distribution)

where:

G = growth rate in inches/year

AG = Weibull scale parameter defined as:

(eq. 4-5)

exp( o = base of the natural logarithm raised to power of the value in brackets

--------------------- - ------------- (All]

A dovetail slot bottom SCC crack is conservatively assumed to extend the full axiallength of the dovetail. Stress intensity, or K, calculations at the dovetail slot bottomassume a crack shape factor, or C, equal to 1.12 (Reference 5). The shape factor isassumed to be log normally distributed with standard deviation of 0.02. Tangentialstresses input to the dovetail SCC stress intensity calculation were obtained from finite


ESBWR Missile Probability Analysis Page 2:3 of 56

element analysis. Dovetail tangential stress is assumed to be log normally distributedwith a standard deviation of 0.02.

4.4 DUCTILE ROTOR BURST

The probability of ductile burst is determined from the Average Tangential Stress (ATS)of each rotor stage and the Ultimate Tensile Strength (UTS). Failure is assumed tooccur when ATS equals or exceeds a fraction of the UTS, as described below.

The ATS values for the ESBWR rotor shown in Table 4-3 are assumed to be log normallydistributed with 0.02 log normal standard deviation.

The dependence of UTS on temperature is assumed to be log normally distributed with0.02 log normal standard deviation and best fit:

UTS = UTSRT - (.032 x (T - 70))

where:

UTSRT= room temperature UTS (ksi)

T= metal temperature IF

Tensile failure is assumed to occur when:

(eq. 4-6)

ATS > R x UTS (eq. 4-7)

The ratio R is assumed to be log normally distributed with 0.85 mean value and 0.05log normal standard deviation.

Table 4-3 ESBWR Average Tangential Stress (1800 rpm)

Stage Bored Rotor ATS (ksi) Boreless Rotor ATS (ksi)1 [[ ----...-- .1.... __ _ .___ _ _ _ _ _ _2 --- _ -__-

3 ----

4 ---- --

5 ---- --

6 --7....___.



5. OVERSPEED PROBABILITY

The overspeed probability calculation methodology described in Reference 1 has beenused for the current study to compute overspeed probability. The overspeed analysisconsiders the characteristics of the turbine control system, the steam turbine unitconfiguration, and test requirements for the steam valves and other overspeedprotection devices/systems.

The operating modes of the TG have been separated, into two event categories: normaland abnormal. The normal events consist of expected operating conditions at orbelow rated speed and the actual overspeed trip test. All normal events are assumedto occur on a regular or planned basis. Any event that is a result of a load rejectionand failure of the overspeed protection system is classified as an abnormal conditionand will be considered within this overspeed analysis. Abnormal events are unplannedand have a small probability of occurrence. Due to the incorporation of fail-safedesigns and testing procedures, speed levels greater than the peak overspeed (seeSubsection 2.2) can be achieved only through a combination of multiple failurescenarios as well as a load rejection. Overspeed may be divided into two groups,

1.) Controlled overspeed whereby the normal control system, emergency trip orthe back-up trip act to close off direct steam paths, and

2.) Runaway overspeed in which a direct steam path from supply to the turbinesection exists due to Main Steam Stop and Control Valves and/or CombinedIntermediate Stop and Intercept Valve remaining open

The probability of turbine overspeed depends on the probability that the main steamvalves and/or intermediate steam valves fail to close when required. There are twoprimary factors that affect this probability. The first is the basic design of theoverspeed protection system, including characteristics of the steam valves andassociated electronic and hydraulic systems. The second factor is the knowledge ofwhether the system's ability to perform has been compromised. The probabilitycalculation assumes that the turbine operator follows General Electricrecommendations; especially those concerned with steam valve and trip system testintervals and hydraulic fluid sampling and maintenance requirements. The analysisassumes that appropriate action is taken when the results of these tests so indicate.

5.1 PROBABILITY CALCULATION

The overspeed analysis presented herein for ESBWR considers a steam turbine systemdesign consistent with the Reference 1 typical General Electric 3-hood steam turbinedesign comprised of an Electro Hydraulic Control System (EHC) with stainless steel tripvalves, and titanium cooler. This design is closest to the ESBWR configuration and thefailure rates used in the analysis are deemed conservative when compared to thecurrent design upgrades employed on ESBWR (i.e. MARK TM IVe control system, modifiedtrip system, new hydraulic fluid conditioning equipment etc.). The resulting assessmentof conservatism will be explained in detail for the valves, hydraulics and controlsarrangements discussed in the following sections.



5.1.1 Steam Valve Arrangement

Turbine control systems manufactured by General Electric provide two independentvalve groups for defense against overspeed in each admission line to the turbine asshown below.

LegendMSV Main Stop ValveMSVs CVs rk] C'nr-+- \Inl-

MainSteam

HP HP TurbineLP LP TurbineMS/R Moisture Separator ReheoterIV Intercept ValveISV Intermediate Stop Valve

ISV

IV

VL- - - - - - --- - - - - - - - - --- -- - - - --- -- - - - - - - t... . .

Feedwater Heaters

Figure 3 Typical Arrangement

The normal overspeed control system closes the main control valves and interceptvalves in proportion to the increase in speed above the overspeed set point, and theETS closes both valve groups via fast acting solenoids upon a rapid turbineacceleration, irrespective of the current turbine speed. Steam from the nuclear steamsupply is admitted through the main stop valves then, enters a manifold, continuingthrough the main control valves to the high-pressure turbine. The manifold ahead ofthe control valves permits in-service testing of the stop valves with little effect on load.The control valves can be individually tested in service by reducing load down toapproximately 85% rated load.

5.1.2 Steam Valve Model

The steam valve model calculates the probability of the steam valves failing to closewhen signaled closed. The model considers two steam paths, the main steam pathand the reheat steam path. The main steam path represents the steam line from thenuclear steam supply, through the main stop and control valves, to the high-pressureturbine section. The reheat steam path represents the steam line from the MSR,through the combined valves, to the low-pressure turbines. The probability ofoverspeed and the speed level reached is a function of the combinations of failuremodes for stop, control, and combined intermediate stop and intercept valves.



The valves controlling the steam flow can be in one of four possible valve responses:fast. closed, closed by servos, tripped closed, or open. The combination of valveresponses results in ten events, as outlined below:

Table 5-1 Steam Path Events for Hydraulic Steam Valve Model

Event Main Steam Path Reheat Steam PathNumber Valving Valving

1 Open Open or Closed by servos2 Closed by servos Open

3 Closed by servos Closed by servos4 Closed by servos Tripped or fast closed5 Tripped Tripped6 Tripped Fast closed

7 Tripped or fast closed Open8 Tripped or fast closed Closed by servos9 Fast closed Tripped

10 Fast closed Fast closed

Probability equations were derived for each event: only the significantshown below, the symbol definitions have also been given for clarity:

events are

Table 5-2 Symbol Definition EHC Hydraulic and Steam Valve Models

Symbol Definition

SV Probability of stuck stop valve (steam side)CV Probability of stuck control valve (steam side)IV Probability of stuck intercept valve (steam side)RV Probability of stuck intermediate stop valve (steam side)

MV Probability of stuck mechanical trip valveEV Probability of stuck electrical trip valveFC Probability of stuck fast acting valve Ion control valve)FI Probability of stuck fast acting valve ion intercept valve)SC Probability of stuck servo valve (on control valve)Sl Probability of stuck servo valve Ion intercept valve)

NC Total number of control valvesNS Total number of stop valvesNP Total number of pairs of combined valvesnc Number of control valves failed to closenp Number of pairs of combined valves failed to close

PNI Probability of no indication of a problem (Hydraulic model only)



Table 5-3 Probability Equations Steam Valve ModelEvent

Number Probability of Overspeed[--- ----- ---

The steam side model is concerned only with incidents where the steam valves fail toclose (e.g., events 1, 5, 6, 7, 9 and 10). The hydraulic and electronic systems areassumed to function properly. The steam valves are assumed to have a constantfailure rate and are returned to a "like-new" condition after testing.

Fundamentally, the ESBWR valves are equivalent to those found on the current fleet.The valve bodies and internals are all of the same geometry and design. The onlyvisible difference is the use of direct actuators on the main steam control and theintercept valves. However, these actuators have the same design features andfunctionality of nuclear actuators of prior designs counterparts (spring fail close,hydraulic piston open, servo valve, solenoid valves, dump valves for fast closure). Thevalve designs for ESBWR and associated overspeed probability are therefore,considered equivalent to the current steam valve model equations used in Reference 1.

5.1.2.1 Valve failure rates

The steam valve failure rates have been calculated using the same methodologyfollowed in the Reference 1 and later General Electric assessments. Steam valvefailure rates are 50% confidence values, based on a chi-square test.

The original failure data used in the Reference 1 report, the failure data collected forthe 1993 Valve Test Interval Extension Supplement and failure data collected for avalve test interval assessment performed in 2008 have been combined to determineupdated steam valve failure rates. The data included actual steam valve test intervalsutilized in the field, valve test profiles and procedures, and valve inspection period andprocedures. The steam valve failure rates have been calculated using the sameapproach followed in the 1984 and 1993 reports and recent assessments completed in2008 and have been incorporated into the steam valve model. The 2008 valve failurerate reflects [[----------------------------- A]] of additional operation sincethe 1993 data and identified that no additional valve failures were experienced.

The overspeed probability resulting from valve failure rates was assessed for valve testintervals of 90-days and 120-days. General Electric recommendation for valve testintervals for existing nuclear units is currently quarterly (90-days). Additional failurerate data at the greater valve test frequency would need to be collected over several



years to improve the confidence in the valve reliability estimates for intervals greaterthan 90-days. However, approximately the same level of missile probability risk isrealized for a valve test frequency of 120-days (with the updated failure rates) versus a90-day test interval with the older failure rates. The values presented in this report arebased on a 120-day valve test interval and are considered acceptable based on theconservatism of the model and the additional valve failure rate data obtained since the1993 report, while maintaining a similar level of risk as compared with a 90-day valvetest interval.

5.1.3 Extraction systems

Extraction systems differ in the amounts of energy available that could contribute tothe overspeed of the TG. Depending on the amount of energy in the extraction line,various check valve schemes are employed, varying from no check valve in the line totwo power actuated check valves.

Each turbine extraction line is reviewed for potential energy and contribution tooverspeed. The number and type of extraction non-return valves required for eachextraction line are specified based on the enthalpy and mass of steam and water inthe extraction line and feedwater heater. Higher energy lines are provided withpower-assisted open, spring-assisted closed non-return valves, controlled by air relaydump valves, which in turn, are activated by the ETS. The closing time of the extractionnon-return valves is sufficient to minimize extraction steam contribution to the turbineoverspeed event.

The model is concerned with turbine overspeed caused by load loss with valves stucksuch that the trip system cannot function. An overspeed event caused by a stuckcheck valve, allowing steam to feed back from extraction processes is excluded fromthe probability model. Although the entrained energy associated with a stuckextraction line check valve is not inconsequential; it is not included in this assessmentas the entrained energy associated with load loss and stuck MS/CV and CIVs isconsidered more limiting for overspeed calculations.

5.1.4 Hydraulic Model

The hydraulic model is concerned with hydraulic component failures due to thecommon failure mode, hydraulic fluid contamination. The steam valves and electronicsystem are assumed to function properly.

All components of the hydraulic system are assumed to have a constant failure rateafter an initial delay time, during which, the probability of failure is zero. Testing ofhydraulic components is not assumed to restore the component to "like new"condition, but is assumed to verify that the system is functional with no noticeabledegradation. It is assumed that the components are restored to "like new" conditionduring each major planned outage. The component failure rates used in the hydraulicmodel are consistent with the values in the Reference 1 report.

Mathematical modeling is required for the hydraulic system to define the probability ofoccurrence of certain combinations of failures (given that hydraulic fluid



contamination is present), to produce an overspeed during a load loss. In addition, themethodology assumes that there may be no indication of such failures (e.g. throughtesting), or no response by the customer, if there is an indication. When the probabilityof multiple failure combinations and the probability of no indication are combined theresult gives the probability that multiple failure combinations of the hydraulic systemcan occur at any time (given that hydraulic fluid contamination is present) and that

* there is no indication of such failures or the indication goes unheeded by the customer.Combining these two probabilities may be accomplished by multiplication since theprobability of no indication is a conditional probability with respect to the multiplefailure combinations.

Table 5-4 Probability Equations EHC Hydraulic Fluid Model

EventNumber Probability of Overspeed

[ --- -------------------

As indicated above, the resultant probability is a conditional probability in that itassumes that hydraulic fluid contamination is present. The contamination is assumedto originate from two sources.

* hydraulic pump failure and corresponding filter failure causes silt to enterhydraulic valves, or

* hydraulic fluid cooler failure allows water into the system which caneventually lead to rust in the hydraulic valves.

It is important to note that the rust is assumed to develop at the valves and the filterdoes not affect the results of the hydraulic cooler model.

An overspeed cannot occur in the hydraulic models without a load loss. In order tohave an overspeed without a load loss a failure has to occur which drives the steam



valves open. No such hydraulic failures are postulated in the models. In summary, foran overspeed to occur in the hydraulic model, there has to be: 1.) a pump failure and afilter failure, and a combination of hydraulic valve failures, and no indication of a failureor an indication with no customer response, and a load loss or; 2.) a cooler failure, anda combination of hydraulic valve failures, and no indication of a failure or an indicationwith no customer response, and a load loss.

The ESBWR hydraulic system will include titanium hydraulic fluid coolers. The effect oftitanium hydraulic fluid coolers is significant in the modeling. This cooler configurationmitigates rusting issues far better than older cooler designs, therefore having a lowerfailure rate.

ESBWR will also incorporate new hydraulic fluid conditioning equipment that replacesthe old "selexsorb and fullers earth" conditioning media with an ion exchanger.Oxidation and acid formation combine to breakdown the fluid and causes sticking andvarnishing of spool valves in the system. Ion exchanger media better controls acidproduction in the hydraulic fluid and does not release metal soaps into the fluid (likethe old conditioning media), which lead to increased air retention and oxidation in thefluid. A dry air blanket system has also been implemented, blowing over the top of thefluid in the reservoir to remove water from the system, as water directly impacts acidproduction in the hydraulic fluid. This equipment has been validated in the field andwill mitigate fluid contamination, resulting in the assessment that the current hydraulicsystem model is considered to be equivalent or conservative.

5.1.5 Emergency Trip System

Although sharing the same basic hydraulic model and common failure modes, theprevious nuclear steam turbine designs differ in the Emergency Trip System layout.Extensive analyses have been conducted for various nuclear turbine controls retrofitprojects, quantifying the replacement of the existing control system, EHC andmechanical trip systems, with the MARK TM IVe and duplex TIR Emergency Trip Devices(ETDs). The previous EHC overspeed protection system consisted of two hydraulic tripvalves; the mechanical trip valve (MTV) and the electrical trip valve (ETV) which havesince been upgraded to a duplex TMR electronic trip device block.

If the normal speed control and power load unbalance function should fail on anESBWR unit, the emergency trip system closes the main and intermediate stop valves.This turbine overspeed protection system comprises the second line of defense againstturbine overspeed. The turbine hydraulic trip solenoid valve hydraulic circuits arearranged in a dual, "two-out-of-three," de-energize to trip configuration. Any powerinterruption to either set of the two-out-of-three trip solenoid valves in the ETD resultsin a turbine trip. The ETD is also fail-safe. Each trip solenoid transfers to the trip stateon a loss of control power, resulting in a turbine trip.

The on-line test of the ETDs provides the capability of individually de-energizing eachof the redundant ETDs and verifying its correct operation without tripping the unit.Only one ETD can be tested at a time, the controls logic allows the remaining ETDs totrip the turbine if necessary.



5.1.6 Turbine Generator Control System (MARK TM Vie)

The ESBWR MARK TM Vie overspeed analysis is leveraged from the MARK TM VieReliability, Availability & Maintainability (RAM) analyses completed in support of controlsystem upgrades to existing nuclear fleet units. These prior analyses include majorsystems, subsystems, and components of the control system including the turbinecontroller, various Input/Output (1/0) devices (i.e. LVDTs, servo valves, pressuretransducers, speed pickups, and proximity probes), and a modified front standard tripsystem. The modified front standard trip system replaces the previous mechanicaloverspeed trip function with a simple, diverse electronic/hydraulic trip system utilizingindependent speed probes and an electric trip device block.

The multiple overspeed protection methods and resulting redundancies incorporatedinto the ESBWR control system are described in Subsection 2.2. Several of the majorESBWR controls improvements are highlighted below:

* Primary speed pickups are each connected to their own Primary TurbineProtection I/O boards.

* Emergency speed pickups are fanned to redundant sets of EmergencyProtection I/O boards (Emergency Protection Terminal Board, SPRO & TurbineEmergency Trip Terminal Board).

* Each of the TMR ETD trip system legs are connected to separate Primary TripTerminal Boards & Turbine Emergency Trip Terminal Boards

* Dual Power Load Unbalance boards, serving dual Relay Output TerminalBoard I/O relay outputs for fast-acting solenoid valves (FASV's, with crossoverpressure transmitters fanned to multiple Analog Input Terminal Boards.

* AC source selector, preserves power source to FASV's in event loss of one ACinput.

The model considers two failure modes; 1) the first is the failure of the control systemdriving the control valves open during either start-up or during overspeed testing, and2) the failure of the overspeed trip systems and the failure of the control system drivingthe control valves open just prior to a loss of load event. Only the latter conditionresults in an overspeed event greater than 120% speed.

The MARK TM Vie TGCS, ETD, and all associated I/0, excluding valves were evaluated forthis same failure mode. Special consideration was given to common mode failuressince this was determined to be the hidden driving cause of system failure. Valveswere not included since they are considered separately in the overspeed model.

Since common mode failures were considered the critical driver to system failure,special attention was given to determine their probability. Completion of a FailureModes and Effects Analysis (FMEA) identified that the types of common mode failuresdid not have a predictable nature and would require a unique computation method.IEC-61508 part 6 has an internationally accepted methodology for computingcommon mode failures as a function of the predicted failure rates of each of the



components. This method was chosen because of the global acceptance of theprocess and the quantitative nature of the results.

The failure rates used in the probability model are based on several sources. TheMARKTM Vie TGCS electronics models are based on Bell Communications Research(Bellcore) parts count predictions. The Bellcore predictions were developed assuming a350C ambient temperature and a 50% applied stress level. The ETD system, I/Otransducers, and sensors were modeled utilizing existing GE Turbine fleet data. Allfailure rates were considered constant over time, which is represented by anexponential failure distribution.

The probability models include redundancy at the system level and, in a few cases, atthe circuit level. The model considers the three simultaneous conditions necessary tooccur to result in an overspeed event:

1) Loss of Load,

2) Failure of the Primary and Emergency Overspeed Trip system, and

3) Failure of the Control system, which results in driving the control valves open.

For this analysis, it was assumed that a load loss occurs once per year. The conditionsthat were critical in modeling the ETD trip system included in the model are:

* One of two parallel ETD sub-systems required to protect/trip the turbine,

" ETDs in each subsystem are arranged in a two-out-of-three configuration andremain energized/valve closed during normal operation,

* Weekly online functional testing of ETDs is performed,

" Functional testing is conducted on a two-year period during refuelingoutages, For ETD system components that are not tested via the weekly ETDtest,

The primary and emergency overspeed protection system is dominated byundetectable common mode failures. An assessment of the system in accordance.with IEC61508-6 results in an assumed 13-Factor of 1% for the controls and 2% for theI/O and ETD. This means that 1% (or 2%) of the sum of component failure rates for allparts in the protection system represent the common mode failure probability. Of thispercentage, 1% is not expected to be detectable. The Diagnostic coverage is assumedto be 99% for the logic and speed sensors and 60% for ETDs.

The TGCS failure probability resulting in driving a control valve open is also dominatedby undetectable common mode failures. An assessment of the system in accordancewith IEC61508-1 results in an assumed 13-Factor of 1% for the controls and 2% for theI/O and ETD. Here again 1% (or 2%) of all component failure rates in the redundantcontrol system represent the probability for the system failure, driving the valve open.Of this percentage, 1% is not expected to be detectable. The Diagnostic Coverage isassumed to be 99% for the logic and 60% for ETDs.


ESBWR Missile Probability Analysis Page :33 of 56

Ultimately, the results of the analyses for the MARK TM Vie TGCS and associated tripsystem upgrade yield a lower overspeed probability when compared to the existingfleet control and mechanical trip systems. However, for conservatism, the probabilityof an overspeed event due to ESBWR MARK TM Vie control system retains the probabilityidentified in Reference (1).



6. CASING PENETRATION

The probability of casing penetration values used, in the overall missile probability'analysis are based on the missile energy analysis 'previously developed by GeneralElectric and verified by full scale casing penetration tests sponsored by the ElectricPower Research Institute (Reference 1). The casing escape probability is a function ofburst speed and includes uncertainty of both fragment behavior and stationarystructure energy absorption capability.

6.1 COMPONENTS

A schematic section of the turbine components included in the casing penetrationmodel is shown in Figure 6-1. Individual wheel burst fragments are assumed to impactonly those stationary components that lie directly in the path of the exiting fragment.The missile is assumed to be a 120-clegree fragment of the wheel. During collision,kinetic energy is dissipated during fracture of the diaphragm and penetration of theinner casing and outer exhaust hood.

Exhaust Hood

Inner Casing

DiaphragmRing

Diaphragm WheelWeb

Fragment

---------- -----------

Shaft---------- ----------------------Centerline

Figure 6-1 Components Included in Penetration Stage Model



6.2 CASING PENETRATION CALCULATIONS

The initial translational and angular velocity of the wheel fragment is determinedbased on conservation of momentum. The linear velocity of the fragment equals theradius of the fragment multiplied by the wheel angular velocity (prior to burst). Theangular velocity of the fragment is assumed equal to the angular velocity of the wheelprior to burst. Translational kinetic energy equals one half the product of mass timestranslational velocity squared. Likewise, rotational kinetic energy is assumed equal toone half the product of polar moment of inertia and angular velocity squared:

KE, =I x M1 x V, 2 (eq. 6-1)21 2

KE,2 = x X f X J.2 (eq. 6-2

As summarized in Reference 1, diaphragm web and ring fragments are assumed to becreated and accelerated as a result of collision. The retained wheel fragment energyis:

2MI f 2

KEf = KE M1 (eq. 6-3)L-M +

where:

KE.T = energy of wheel fragment after collision

KEfi = energy of wheel fragment before collision

Ms = mass of wheel fragment

M,,, = mass of diaphragm fragments that are directly in the path of the wheel

fragments

The energy lost by the wheel fragment during penetration of the inner casing wrapperand exhaust hood is calculated from the empirical relation developed by Moore(Reference 6) commonly referred to as the "Stanford Formula". The Stanford formulaapplies to missiles having a right circular solid shape that impact a flat plate with theaxis of the cylinder normal to the plate. The energy loss during penetration (ft-lbs) is afunction of both the plate ultimate tensile strength (UTS) and thickness (T) as follows:

KEloss = UTS x DIA x T, x (.344T. + .O08DIA) (eq. 6-4)



where:

UTS = casing material ultimate tensile strength (psi)

DIA = equivalent circular diameter of fragment (in)

Tc= casing thickness (in)

KE._ = Energy Lost by Fragment (ft-lbs)

Based on the preceding discussion and further consideration of fragment size,orientation during penetration, impact fragment direction, energy at fracture andenergy absorption, a statistical spread in escaping or external energy is derived. Theprobability of casing penetration is taken to be the probability that the external energyis greater than zero.



7. OVERALL PROBABILITY DETERMINATION

7.1 PROBABILITY OF BRITTLE FRACTURE (K1 >Kjc)

The probability that the rotor will burst by time t, is the probability that K, is greater

than or equal to Kic by time t,. For each turbine stage, multiplying the following threeconditional probabilities:

* Crack depth existing

* Excess temperature occurring

" Ký > Kc given the crack depth and excess temperature

and integrating over the whole range of possible rotor material excess temperaturesgives the probability that the given crack depth will cause a burst by time t2. A furtherintegration over all possible crack sizes will give the probability of a rotor burst by timet2 for any depth of crack. Time is a factor since the crack model assumes that thecrack size increases with time. Stated mathematically:

Pb (t21ou, T) = J J JI(LogK, - LogK,,)fe(y)ff (X)fa (a)dydXda (eq. 7-1)0

where:

P, (t2 o,T) = cumulative burst probability at time = t 2 ,given stress (() and

temperature (T)

I(LogK, >_ LogK,,) = 1 if y, X, and a are such that Log Ki > Log Kic.=0 otherwise

fe(y) =the probability density of the error function

f- (X) = the probability density of excess temperature

fa(a) = the probability density of crack depth

-,T = Operational stress at the desired speed and temperature for the stage inquestion. (Each turbine stage has a unique combination of stress andtemperature.)



11- -- --

Page :38 of 56

This page is deleted based on proprietary information.

(Al]]



[[ ........

Pa]ge 39 of 56


-- -------- D(A]




__-- --------- A-]]



------------ WD

7.2 PROBABILITY OF DUCTILE TENSILE FAILURE

The probability of wheel burst in the ductile mode is determined from the AverageTangential Stress (ATS) and the Ultimate Tensile Strength (UTS). Failure occurs whenthe stress equals a set ratio of strength as described in subsection 4.4. As noted insubsection 4.4, UTS, ATS, and the ratio R are all assumed to be log normally distributed.

The ultimate tensile strength of the wheel material varies with the temperature of thematerial and the stress in the wheel varies with the speed of rotation. Therefore, theprobability of wheel burst in the ductile mode is conditional on the temperature of thewheel and on the speed of rotation.

7.3 NORMAL OPERATION

The probability of a rotor burst is a function of time, start/stop cycles, temperature, andspeed. Time and cycle count determine the overall crack size distribution. Largercracks lead to higher probability of brittle failure. Temperature is important in thedetermination of the rotor material toughness. Lower temperature will reduce thetoughness and increase the probability of brittle burst. Speed is important indetermining the stress intensity factor since a higher speed results in larger wheelstresses and larger stress intensity factors and therefore leads to a higher' burstprobability. Larger stresses also increase the ductile failure probability. Thus, theparameters of time, temperature, and speed must be properly accounted for in theoverall probability determination.

During operation, low-pressure rotor disks are subjected to various speeds andtemperatures. Each start/stop operating cycle shown in the Table 4-3 is assumed toconsist of the elements shown in Table 7-1. The occurrence of these conditions isconsidered normal and is assigned a probability of one.



Table 7-1 Conditions in a Normal Operating CycleDescription Speed/Rated Speed Min. Temp (OF)

Shutdown 0 50Turning Gear 0.001 50

Start (Unsync'd) 1 75Loading (including Thermal Stress) 1 75

Part Load 1 120Full Load 1 225Unloading 1 120

O/S Trip Test 1.11 120Coastdown 1 120

Cumulative burst probability (equation 7-1) is then calculated for each element of thestart/stop cycle shown in Table 7-1 and the maximum value determined as a functionof time for each stage. Figure 7-1 illustrates the maximum predicted cumulative burstprobability for a typical single LP stage derived from equation 7-1.

This cumulative burst probability (example Figure 7-1) defines a one-time burstprobability given no prior stressing of the material. The normal annual burstprobability is the annual failure rate given that no failure has occurred previously. Thisvalue is determined from the instantaneous slope of the cumulative burst probabilitycurve (example Figure 7-1) at the time of interest for the stage in question.

FAtio

Figure 7-1 Maximimum Cumulative Burst Probability of Typical Stage (Normal Operation)

Stated mathematically, annual normal burst probability is derived from the cumulativeburst probability as follows:



PBN'W) d[PMAX Wt>PBMAX(0)I (eq.7-17)where:

PBN (t) = annual normal burst probability at time = t

t = time of interest, in years

PBMAX(t) = maximum individual value of rotor burst cumulative probability attime = t (for the 9 normal "events")

PBMAX(O) = highest value of initial rotor burst cumulative probability at time t = 0(for the 9 normal "events")

During numerical evaluation, the annual normal cumulative burst probability iscalculated separately for the dovetail, body,, and tensile mechanisms and thencombined for each stage and time of interest.

7.4 ABNORMAL OPERATION

An abnormal event is defined as an occurrence of a control system failure and a fullload rejection by the turbine generator that causes operation above the normaloperating speed. Each abnormal event has a maximum speed and an annualprobability of occurrence as shown in Table 7-2. The values shown in Table 7-2 reflectfour month (120-days) valve testing frequency.

Table 7-2 Typical Data for Abnormal Events

Event Number 3 Speed/Rated Speed Abnormal Event Annual Probability4

E----- -- -

--. . .. . . .. . . .

---. .. .-------

3 Reference Table 1.5-14 Values are based on the results of the abnormal probability model discussed in Section 1.5.



When the abnormal event occurs, speed increases from normal running speed andcontinues until it reaches the maximum speed for the particular event. Burst will notoccur until the cumulative burst probability exceeds the level attained during normaloperation. The probability that the stage will burst at this temperature, given theoccurrence of the abnormal event, is simply the difference between the conditionalburst probability at the maximum speed and that at the worst cumulative burstprobability encountered during normal operation. A wheel stage will only burst due toan abnormal event if it has not burst during prior normal operation (conditionalprobability). Therefore, this probability difference must be divided by the probabilitythat no burst has occurred during normal operation:

PBA W Pit )-~j I P. (1) eq.7-18)PA~X~I )-PBMA (t

where:

P'A (t) = annual abnormal burst probability at time =t

PAi= annual probability of occurrence of abnormal event i. (ref. Table 7-2)

PBAi (t a-;T)= annual cumulative burst probability at time = t, given that theabnormal event U) occurs

PBMAx (t) = cumulative burst probability for the worst normal operatingcondition

oi = stress corresponding to the maximum speed of the event (reference Table

4-1 and noting that mechanical stresses increase with the square of speed)

T= minimum temperature during normal operation

During numerical evaluation, the annual abnormal cumulative burst probability.PBA,(tiu,,T), is calculated separately for the dovetail, body and tensile mechanisms

and then combined prior to the event summation indicated by equation 7-18. Inaddition, each stage is evaluated separately.

The probability of burst during abnormal operation is conditional on the stage notbursting during normal operation by the time in question. When the normal andabnormal individual burst probabilitites are combined, the probability of burst duringabnormal operation must be multiplied by the probability of no stage burst duringnormal operation. The total single stage annual burst probability is then determined asfollows:



PB (t) = PN (1)+ [1 - P.N (t)]x P"A (t)where:

P, (t) = annual stage burst probability at time = t

(eq. 7-19)

Figure 7-2 illustrates the total annual burst probability predicted for the same stageshown in Figure 7-1 using the above procedures and the abnormal event probabilitiesshown in Table 7-2.

Figure 7-2 Annual Burst Proability for Typical Stage (Normal and Abnormal Operation)

Missile probability calculation requires the addition of casing escape probability toEquations 7-17 and 7-18 as follows:

PMN (t)= PEPBmax X PBN(t) (eq. 7-20)

where:

PMN(t)= annual stage missile probability at time = t associated with normal

operation

PEJ'max = stage casing escape probability given that burst has occurred duringthe worst normal operating condition



PMA W)=PEX PAix -PBAi(ji5T)-PBMA4A(t) (eq. 7-211

where:

P= conditional casing escape probability given that a burst has occurred for

the stage in question during abnormal event i (reference Table 7-21

PMA (t) = annual stage missile probability at time =t associated with the abnormalevent i

Similar to Equation 7-19, the probability of generating a turbine missile external to thecasing during abnormal operation is conditional on the stage not bursting duringnormal operation by the time in question. When the normal and abnormal missileprobabilities are combined, the probability for an external missile during abnormaloperation must again be multiplied by the probability of no stage burst during normaloperation:

PM (t) = PMN (t)+ - PN (t)]x MA (t) = P (eq. 7-221

where:

PM (t) = annual stage missile probability at time = t

As noted, the rotor is assumed to be made up of a series of independent wheel stagesand individual stage burst probabilities are combined to obtain the probability that atleast one wheel stage on the rotor will burst. Missile probabilities are similarlycombined.



8. BORELESSI(SOLID) ROTOR RESULTS SUMMARY

The calculated annual missile probability for an individual ESBWR turbine containingthree statistically independent solid LP rotors (42 stages in total) and valve testing in120-day (4 month) intervals is shown in Figure 8-1. The annual missile probability (P11at the 12-year inspection frequency is 1.2x10-7 and remains less than 1x10-5 forgreater than 50 years of turbine operation, with the assumption that no correctiveactions are taken to address any observed crack indications in the dovetail slotbottom.[[i

WAl]

Figure 8-1 Unit Featuring Solid Rotors: Annual Missile Probability

The missile probability assessment described in Sections 6 and 7 applies to both a solid(boreless) and bored rotor. The statistical distribution of the crack size is dependent onthe bore configuration. For the solid rotor, the rotor body undetected flaw is assumedto be a half penny shaped embedded crack that is oriented in the radial axial plane. Bycomparison, the bored rotor undetected flaw is assumed to be a half penny shapedsurface crack. A smaller undetected crack can be claimed because a bored rotor isinspected from both the outer periphery and the inner periphery (e.g., bore surface). Incontrast, the solid rotor is only inspected from the outer periphery. These concepts aresummarized in subsection 4.2.2. Section 9 provides a comparison of results between abored and solid rotor. The bored rotor stresses shown in Tables 4-1 and 4-3 arereflected in Figure 8-1 as is the undetected flaw assumptions

The flat region of Figure 8-1 (up to [[ --- (AWl years) is dominated by valve failure (event#1 in Table 7-2). The probability shown in this region is associated with gross tensile



failure during the subsequent overspeeding of the turbine. The curve is flat in thisregion because 1) rotor tensile strength is time invariant and 2) the probability of thevalve failure is dependent only on the in-service valve-testing interval.

The gradual increase in missile probability beginning after[[ --- }]] years reflects theincreased probability with time of a dovetail SCC crack reaching critical size. Theconservative SCC crack growth model described in subsection 4.3 results in a gradualincrease in predicted crack depth with time. The gradual increase in missile probabilityis a reflection of the expected increase in crack size with time.

The SCC crack initiation and growth distributions shown in Equations 4-4 and 4-5 andused in the calculations summarized in Figure 8-1 are based on historical fieldexperience with shrunk-on wheel bore keyway cracks. The use of this data isconsidered to be conservative when applied to the ESBWR LP monoblock rotors. Thejustification for this assessment is based on: 1.) The ESBWR rotors feature greatlyreduced concentrated tensile stress magnitudes in critical locations (vs. earlier shrunk-on wheel keyways), and 2.) critical locations will be shot-peened and thus feature abeneficial layer of compressive residual stress at or below the outer surface (SCCcharacteristics utilized in this report are from non shot-peened shrunk-on

MFN-09-484 Enclosure 2 GE-ST ESBWR Steam Turbine - Low ... · Fracture Toughness -ASTM toughness tests of NiCrMoV material is the basis for the critical stress intensity versus excess

Documents