ML15022A661

Stress Re-Evaluation of Nine Mile Point Unit 2 Steam Dryer at 115% CLTP, CDI Report No. 14-08NP, Revision 0, Non-proprietary Version

ML15022A661
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Site:	Nine Mile Point
Issue date:	07/31/2014
From:	Boschitsch A H Continuum Dynamics
To:	Office of Nuclear Reactor Regulation
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ATTACHMENT 2 NONPROPRIETARY VERSION OF CDI REPORT NO.14-08P This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information CDI Report No. 14-08NP Stress Re-Evaluation of Nine Mile Point Unit 2 Steam Dryer at 115% CLTP Revision 0 Prepared by Continuum Dynamics, Inc.34 Lexington Avenue Ewing, NJ 08618 Prepared under Purchase Order No. 7736902 for Constellation Energy Group Nine Mile Point Nuclear Station, LLC P.O. Box 63 Lycoming, NY 13093 Prepared by Alexander".

Boschitsch Approved by Alan J. Bilanin July 2014 This report complies with Continuum Dynamics, Inc. Nuclear Quality Assurance Program currently in effect.

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Executive Summary The stresses resulting from acoustic loads at the 115% CLTP operating condition (also referred to herein as the extended power uprate or EPU condition) are re-evaluated for the Nine Mile Point Unit 2 (NMP2) steam dryer using a finite element model and frequency-based analysis methodology.

The re-evaluation is performed following identification

[1] and correction

[2] of an inconsistency in the acoustics loads predictions that resulted in a new set of loads requiring a stress analysis.

The re-evaluation produced a new stress state that required a more detailed examination of the stress state at a limited number of locations and, for one location, necessitated installation of a modification to limit vibration of the inner side plate connecting the inner vane banks. In all other respects, the finite element model of the steam dryer (including the modifications

[3-6] to achieve an alternating stress ratio at EPU operation of SR-a>2) are essentially identical to those previously described in [5]. Similarly the stress evaluation is consistent with those carried out in the U.S. for prior dryer qualification to EPU conditions.

The resulting stresses are assessed for compliance with the ASME B&PV Code 2007 [7],Section III, subsection NG, for the load combination corresponding to normal operation (the Level A Service Condition).

The acoustic loads are prepared using the acoustic circuit model (ACM) version 4.1 R [2, 8, 9]. This version represents the analysis that has been corrected to address an inconsistency

[1] in the representation of the acoustic solution in the gap between the above-water skirt and reactor vessel wall, and also a minor deviation associated with the use of single rather than double precision in the Helmholtz solver used to procure acoustic loads. Three load conditions are considered:

Baseline:

This is the normal operating load associated with EPU steam dryer operation.

Prior to power ascension this was the only relevant load at 115% CLTP power.Drain Trap Out-of-Service:

This off-normal, but not infrequent operating condition was identified during power ascension and is associated with a noticeable 92.5 Hz signal occurring when the drain trap in the reactor core isolation cooling (RCIC) system is isolated.RCIC Valve Closed: This off-normal operating load occurs when the RCIC valve is intentionally closed and is characterized by an 89.3 Hz signal.The present analysis evaluates the complete dryer using the Baseline load processed using ACM 4.1 R. The identified limiting locations are then used to construct a node list that in turn is provided as input to real time stress evaluations at the other off-normal conditions.

In a real time stress evaluation, stresses are only calculated at the nodes in this list rather than for the entire dryer. As described in Section 5.7, the real time evaluations use conservative MSL entrance signals developed by retaining the higher of the total uncertainty identified with the ACM 4.1 and ACM 4.1 R loads models over each frequency interval.It is found that application of the baseline load to the steam dryer results in several locations requiring additional refined modeling to accurately define the alternating stress ratio. These locations have been addressed through a combination of high resolution modeling, repairing poor quality mesh elements and improved modeling of connections to reproduce the as-built configuration.

For one location involving the common junction between the inner side plate, top i This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information and side plates of the inner vane banks and tie bar, it was shown that the recommended margin of 2.0 could be achieved using a combination of embedded modeling and stress interpolation along the weld lines. However, uncertainty regarding the applicability of the interpolation procedure at this complex geometry junction involving multiple welds motivated consideration of alternate concepts for meeting margin. The long-term resolution ultimately adopted was to install the U-stiffener modification.

This modification-based approach was considered appropriate as the high stress estimate was attributed to a vibration mode of the end plate and was further increased by the frequency shifting required by the analysis.With this channel in place the real time stress evaluation shows that all nodes have a peak stress ratio, SR-P, of 1.3 or higher at all load combinations thus meeting the required margin for this stress type. With regard to alternating stresses, all of the nodes on the steam dryer have an alternating stress ratio of 2.0 or higher under the baseline and drain trap out-of service loads so that the dryer qualifies for these conditions.

For the RCIC valve closed, certain locations have alternating stress ratios below 2.0 with the minimum value being SR-a=I.5.

However, since this condition occurs infrequently, it is appropriate to assess fatigue using cycle counting.

The cumulative usage factors (CUFs) for these locations are calculated in [10] and shown to all lie well below 1.0. Taken in their entirety, these results show that the dryer qualifies for level A service operation with the U-section stiffener installed.

In producing these results refined estimates of the linearized stresses at selected high stress locations were obtained using (3)ii This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table of Contents Section Page Executive Sum m ary .........................................................................................................................

i Table of Contents ...........................................................................................................................

iii Nom enclature

.................................................................................................................................

iv 1. Introduction and Purpose ............................................................................................................

1 1.1 Summ ary of Overall Evaluation M ethodology

.................................................................

3 2. M ethodology

& Evaluation Procedures

.................................................................................

5 2.1 Overview ...............................................................................................................................

5 2 .2 [[ ......................................................

(3)]] ..............................................................

7 2.3 Computational Considerations

..........................................................................................

8 3. Finite Element M odel Description

........................................................................................

11 3.1 Steam Dryer Geometry ...................................................................................................

11 3.2 M aterial Properties

..............................................................................................................

16 3.3 M odel Simplifications

......................................................................................................

16 3.4 Perforated Plate M odel ...................................................................................................

17 3.5 Vane Bank M odel ...............................................................................................................

19 3.6 W ater Inertia Effect on Submerged Panels ......................................................................

20 3.7 Structural Damping ........................................................................................................

20 3.8 M esh Details and Element Types ...................................................................................

20 3.9 Connections between Structural Components

..............................................................

21 3.10 Pressure Loading ...............................................................................................................

33 3.11 [[ ....................................................( ].......................

36 4. Structural Analysis ....................................................................................................................

37 4.1 Static Analysis ....................................................................................................................

37 4.2 Harm onic Analysis ..............................................................................................................

37 4.3 Post-Processing

...................................................................................................................

43 4.4 Computation of Stress Ratios for Structural Assessment

...............................................

43[[4.5 .............................................

(3)]] .................................................................

46 5 .R e su lts .......................................................................................................................................

5 3 5.1. Preliminary Stress Assessment of Normal EPU Load + Base Dryer Model .................

54 5.2 Construction of Real Tim e Node List .............................................................................

65 5.3 Examination of Low Stress Ratio Locations

.................................................................

69 5.4 General Stress Distribution and High Stress Locations

.................................................

81 5.5 Load Combinations and Allowable Stress Intensities

...................................................

90 5.6 Frequency Content and Filtering of the Stress Signals .....................................................

108 5.7 Real Time Analysis W ithout U-Stiffener

........................................................................

117 5.8 Real Time Analysis Adjusted with U-Section Stiffener Included ...................................

121 6. Conclusions

.............................................................................................................................

123 7. References

...............................................................................................................................

125[[ .................................................

]]..... ............................

128 Appendix B. U-Section Stiffener

...............................................................................................

132 iii This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Nomenclature ACM acoustic circuit model CDI Continuum Dynamics, Inc.CLTP current licensed thermal power CUF cumulative usage factor DOF degree of freedom DRF design record file FFT fast Fourier transform FSI fluid-structure interaction EPU extended power uprate FEA finite element analysis IGSCC inter-granular stress corrosion cracking MSL main steam line NRC Nuclear Regulatory Commission ooS out-of-service Pm membrane stress intensity Pb bending stress intensity PT pressure transducer QC Quad Cities RCIC reactor core isolation cooling RFO refueling outage RPS reduced point set (described in Section 5)RPV reactor pressure vessel Sa service limit for alternating stress intensity Salt alternating stress intensity SER safety evaluation report Sm service limit for membrane stress intensity SR-a alternating stress ratio SR-P peak stress ratio SRF stress reduction factor SS stainless steel NMP2 Nine Mile Point Unit 2 USR upper support ring WEC Westinghouse Electric Corporation WF weld factor iv This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

1. Introduction and Purpose Current licensing procedures to qualify the Nine Mile Point Unit 2 (NMP2) nuclear plant for operation at Extended Power Uprate (EPU) operating condition require a stress assessment of the steam dryer to ensure adequate stress margins under the increased loads. The steam dryer loads due to acoustic pressure fluctuations in the main steam lines (MSLs) are potentially damaging and the cyclic stresses from these loads can produce fatigue cracking if loads are sufficiently high. The industry has addressed this problem with physical modifications to the dryers, as well as a program to define steam dryer loads and their resulting stresses.

The EPU qualification process requires a stress evaluation for the NMP2 steam dryer using acoustic loads acquired at EPU operation.

The present report documents this stress evaluation by calculating the maximum and alternating stresses generated using strain gage MSL pressure measurements acquired at EPU operation.

It updates a previous stress evaluation carried out in [5] to account for identification

[1] and correction

[2] of an inconsistency in the acoustics loads predictions which resulted in a new set of loads requiring a stress analysis.The load combination considered here corresponds to normal operation (the Level A Service Condition) and includes fluctuating pressure loads developed from NMP2 main steam line data, and weight. The fluctuating pressure loads, induced by the flowing steam, are predicted using a separate acoustic circuit analysis of the steam dome and main steam lines [13]. Level B service conditions, which include seismic loads, are not included in this evaluation.

Stress ratios are obtained by comparing these stresses (appropriately adjusted at welds) against allowable values and used to ensure compliance with the ASME Code (ASME B&PV Code,Section III, subsection NG).The stress analysis is carried out in the frequency domain, which confers a number of useful computational advantages over a time-accurate transient analysis including the ability to assess the effects of frequency scaling in the loads without the need for additional finite element calculations.

The analysis develops a series of unit stress solutions corresponding to the application of a unit pressure at a MSL at specified frequency, f. Each unit solution is obtained by first calculating the associated acoustic pressure field using a separate analysis that solves the Helmholtz equation within the steam dryer [14]. This pressure field is then applied to a finite element structural model of the steam dryer and the harmonic stress response at frequency, f, is calculated using the commercial ANSYS 10.0 finite element analysis software.

This stress response constitutes the unit solution and is stored as a file for subsequent processing.

Once all unit solutions have been computed, the stress response for any combination of MSL pressure spectrums (obtained by Fast Fourier Transform of the pressure histories in the MSLs) is determined by a simple matrix multiplication of these spectrums with the unit solutions.

Details of the frequency-based stress evaluation methodology are contained in Section 2.0.EPU Load Conditions The stress evaluation is performed for three different EPU load conditions described in [2].In addition to the normal baseline EPU load, two other loads were identified during power ascension.

The first occurs when the drain trap is out of service (ooS) and primarily differs from the baseline load by the presence of a distinct 92.5 Hz signal. This condition occurs sufficiently often that it is treated as an alternative

'base' EPU load. It mostly, but not always, produces I This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information slightly higher stresses and so is considered along with (rather then replacing) the baseline EPU load. The third condition occurs when the reactor core isolation cooling (RCIC) valve is closed and is characterized by an 89.3 Hz signal. This signal occurs infrequently and therefore is amenable to fatigue assessment based on cycle counting methods to calculate the cumulative usage factor.The signal files identified with these three EPU loads are documented in [2] as: 20120721142636 (baseline EPU); 20130524092644 (drain trap out-of service);

and 20120904092600 (RCIC valve closed).Acoustic Loads Estimation The current stress evaluation of the NMP2 steam dryer is performed using acoustic loads generated using a revised Acoustic Circuit Model (ACM) Rev. 4.1 R [2, 8, 9]. The development of this revision was motivated primarily by a requirement for consistent usage of noise filtering strategies during both model calibration against available data and application of the model to plants. Other than the removal of known non-acoustic discrete frequencies (e.g., electrical noise at multiples of 60 Hz) and the application of coherence filtering (which was also invoked when processing the Quad Cities data) no other filtering methods are used. In particular, no noise subtraction using low power data is performed.

Further details of the ACM Rev. 4.1 R calibration activity and its application to obtain NMP2 steam dryer acoustic loads are detailed in[2, 8, 9]. As described in [9] re-benchmarking the ACM against available Quad Cities data produced updated estimates of the acoustic speed and damping in the acoustics description and also revised biases and uncertainties due to changes in the model, coherence-based noise filtering and comparison method. The biases and uncertainties used for the present load estimates are based on the comparison with QC data at 790MWe using 16 sensors. For each frequency interval, the biases and uncertainties obtained in ACM 4.1 R are similar to those in ACM 4.1.While it is technically consistent to use the ACM 4. 1R values in the evaluations, the higher of the ACM 4.1 and ACM 4.1 R total uncertainties (bias + uncertainty) over each frequency interval is also defined to be consistent with the NMP2 SER. The real time stress assessments are carried out using these conservative signals consistent with the NMP2 SER[ 15].Stress Processing The frequency-based harmonic stress evaluation methodology, finite element model and post-processing procedures

((3)) are fully identical to those described in the previous NMP2 steam dryer stress evaluation at EPU conditions

[5] and are described in Sections 2-4. In order to qualify the NMP2 steam dryer for normal EPU operation it is required that the limiting alternating stress ratio be above a target level of 2.0. To meet margin several modifications to the dryer were implemented prior to power ascension (in 2012). The modifications made to the dryer are described in Section 5 of [5] and also detailed in [6]. These modifications are fully accounted for in the current stress evaluation including the stiffened closure plate, the masses added to the inner and middle hoods, and the 1/8" thick reinforcement plate placed over the middle hood section outboard of the closure plate.Based on the present stress evaluation an additional modification consisting of a U-section stiffener bolted onto the inner side plate spanning the inner vane banks, was added. This modification is described in Appendix B.2 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]1.1 Summary of Overall Evaluation Methodology The stress evaluation involves:* 3 different collected signals -baseline, drain trap ooS, and RCIC valve closed.* 2 bias and uncertainty combinations

-one pertaining to ACM 4.1R and the other corresponding to the higher of the total uncertainty in ACM 4.1 and ACM 4.1R in each frequency interval.* 4 different steam dryer models -(i) the baseline model for which the full calculation was performed over the full 0-250 Hz frequency range; (ii) a revised model where the connections between the lifting rod and upper lifting rod brace are changed to the as-built configuration; (iii) an improved baseline model where a poor quality grid near a lifting rod brace was remeshed to improve element quality; and (iv) a modified model containing the U channel stiffener on the inner side plate. The latter models were only developed over reduced frequency ranges.Given the large number of analysis permutations, an overall strategy was developed that conducted a full dryer analysis using the baseline load and available full frequency range unit solution, and then executed a series of real time evaluations to analyze the other permutations.

This strategy proceeded as follows (sections where the calculations are performed are given in parentheses):

1. Analyze complete dryer using the baseline load and baseline structural model (Section 5.1).2. Identify the limiting locations and assemble a list of nodes for real time processing (Section 5.2).3. For locations with alternating stress ratios, SR-a<2, exercise analytical options to obtain more accurate stress estimates.

For the current analysis these options included: a. Improving mesh quality about high stress locations.

b. Developing and/or utilizing embedded models at high stress sites.c. Changing the connection between the lifting rod and top brace from welded (indicated in drawings) to non-welded (as-built).

d. Interpolating along the welds to distinct hotspots -this is only done at the tops of the welds connecting the closure plates and curved (inner and middle)hoods.(Section 5.3)3 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

4. Using the results from step 3, repeat step 1. Also generate stress ratio tables and PSDs (Sections 5.4-5.6).5. Perform real time stress evaluations for the other loads and total uncertainties using the node list generated in step 2 (Section 5.7).6. Consider modifications for locations with alternating stress ratios that remain below 2.0 in step 5. For the current stress evaluation, the resulted in the attachment of a U-shaped cross-section stiffener to the inner side plate.7. Conduct real time analyses using a FEA model that implements the modifications in step 6.8. For locations that still have alternating stress ratios, SR-a<2.0 either: (i) consider additional modifications and repeat real time analysis (this option proved unnecessary in the current evaluation), or (ii) for stress ratios below 2.0 occurring at the RCIC valve closed EPU condition conduct cycle counting to confirm cumulative fatigue usage below 1.0 for this infrequent load condition.

The results from the execution of this overall strategy are presented in the following sections: Section 5 presents a full stress assessment of the baseline structural model subjected to the baseline load (steps I & 2). The results from this analysis motivate Section 6 which summarizes the measures carried out to address the locations with alternating stress ratios below 2.0. These include both analytical approaches (step 4) and the addition of physical modification (the U-section stiffener

-step 6). Section 7 presents a reanalysis of the full steam dryer under the baseline load, but with the analytical corrections in step 4 now applied (part of step 5). It also includes a series of real time analyses at the other load conditions both without (other part of step 5) and with (step 7) the U-section stiffener.

This stress evaluation reports that the limiting alternating stress ratio on the dryer at the baseline EPU load is SR-a=2.0 and at the drain trap ooS condition, SR-a=2.0.

These results are based on the most conservative combinations of the bias and uncertainties from ACM 4.1 and ACM 4.1 R, and assume that the U-section stiffeners are installed.

For the RCIC valve closed condition, the limiting alternating stress ratio, SR-a=1.52.

This value meets the ASME stress margin, but not the NRC staff recommended value of 2.0. However, because this load condition is infrequent it is addressed using cycle counting in [10]. The limiting peak stress ratio for any of the three load conditions due to maximum membrane and bending stresses including static contributions is SR-P=I.3.

These values show that the present modified steam dryer meets the recommended stress margin at EPU operation.

Flaw evaluations for flaws identified in the original dryer are summarized in [16] and in Section 2.5 of [5]; evaluations of flaws identified during 2014 refueling outage are documented in [17].4 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

2. Methodology

& Evaluation Procedures

2.1 Overview

Based on previous analysis undertaken at Quad Cities Units 1 and 2, the steam dryer can experience strong acoustic loads due to the fluctuating pressures in the MSLs connected to the steam dome containing the dryer. C.D.I. has developed an acoustic circuit model (ACM) that, given a collection of strain gage measurements of the fluctuating pressures in the MSLs, predicts the acoustic pressure field anywhere inside the steam dome and on the steam dryer [8, 9, 13, 14, 18]. The ACM is formulated in frequency space and contains two major components that are directly relevant to the ensuing stress analysis of concern here. (3)5 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]6 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]2.2 [[Continuum Dynamics, Inc. Proprietary Information (3)11 2.3 Computational Considerations Focusing on the structural computational aspects of the overall approach, there are a number of numerical and computational considerations requiring attention.

The first concerns the transfer of the acoustic forces onto the structure, particularly the spatial and frequency resolutions.

The ANSYS finite element program inputs general distributed pressure differences using a table format. This consists of regular 3D rectangular (i.e., block) nxxnyxnz mesh where n(, is the number of mesh points in the i-th Cartesian direction and the pressure difference is provided at each mesh point (see Section 3.10). These tables are generated separately using a program that reads the loads provided from the ACM software, distributes these loads onto the finite element mesh using a combination of interpolation procedures on the surface and simple diffusion schemes off the surface (off-surface loads are required by ANSYS to ensure proper interpolation of forces), and written to ASCII files for input to ANSYS. A separate load file is written at each frequency for the real and imaginary component of the complex force.The acoustic field is stored at 5 Hz intervals from 0 to 250 Hz. While a 5 Hz resolution is sufficient to capture frequency dependence of the acoustic field (i.e., the pressure at a point varies gradually with frequency), it is too coarse for representing the structural response especially at low frequencies.

For 1% critical structural damping, one can show that the frequency spacing needed to resolve a damped resonant peak at natural frequency, fn, to within 5% accuracy is Af=0.0064xfn.

Thus for fn=1 0 Hz where the lowest structural response modes occur, a frequency interval of 0.064 Hz or less is required.

In our calculations we require that 5% maximum error be maintained over the range from fn=5 Hz to 250 Hz resulting in a finest frequency interval of 0.0321 Hz at the low frequency end (this adequately resolves all structural modes up to 250 Hz). Since there are no structural modes between 0 to 5 Hz, a 0.5 Hz spacing is used over this range with minimal (less than 5%) error. The unit load, fn(O),R), at any frequency, Ok, is obtained by linear interpolation of the acoustic solutions at the two nearest frequencies, coi and coi+1, spaced 5 Hz apart. Linear interpolation is sufficient since the pressure load varies slowly over the 5 Hz range (linear interpolation of the structural response would not be acceptable over this range since it varies much more rapidly over the same interval).

Details regarding the frequency resolution have been provided in [20].Solution Management

(3)8 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]Structural Damping In harmonic analysis one has a broader selection of damping models than in transient simulations.

A damping factor, z, of 1% critical damping is used in the structural analysis.

In transient simulations, this damping can only be enforced exactly at two frequencies (where the damping model is "pinned").

Between these two frequencies the damping factor can by considerably smaller, for example 0.5% or less depending on the pinning frequencies.

Outside the pinning frequencies, damping is higher. With harmonic analysis it is straightforward to enforce very close to 1% damping over the entire frequency range. In this damping model, the damping matrix, D, is set to D=2zK (7)()where K is the stiffness matrix and co the forcing frequency.

When comparing the response obtained with this model against that for a constant damping ratio, the maximum difference at any frequency is less than 0.5%, which is far smaller than the 100% or higher response variation obtained when using the pinned model required in transient simulation.

Load Frequency Rescaling One way to evaluate the sensitivity of the stress results to approximations in the structural modeling and applied loads is to rescale the frequency content of the applied loads. In this procedure the nominal frequencies, cok, are shifted to (l+X)cok, where the frequency shift, X, ranges between +10%, and the response recomputed for the shifted loads. The objective of the frequency shifting can be explained by way of example. Suppose that in the actual dryer a strong structural-acoustic coupling exists at a particular frequency, o*. This means that the following conditionshold simultaneously: (i) the acoustic signal contains a significant signal at co*; (ii) the structural model contains a resonant mode of natural frequency, con, that is near co*'; and (iii) the associated structural mode shape is strongly coupled to the acoustic load (i.e., integrating the product of the mode shape and the surface pressure over the steam dryer surface produces a significant modal force). Suppose now that because of discretization errors and modeling idealizations that the predicted resonance frequency differs from co* by a small amount (e.g., 1.5%). Then condition (ii) will be violated and the response amplitude therefore significantly diminished.

By shifting the load frequencies one re-establishes condition (ii) when (1+ X)"o* is 9 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information near (On- The other two requirements also hold and a strong structural acoustic interaction is restored.[[I (3)]Evaluation of Maximum and Alternating Stress Intensities Once the unit solutions have been obtained, the most intensive computational steps in the generation of stress intensities are: (i) the FFTs to evaluate stress time histories from (5); and (ii) the calculation of alternating stress intensities.

(3)The high computational penalty incurred in calculating the alternating stress intensities is due to the fact that this calculation involves comparing the stress tensors at every pair of points in the stress history. This comparison is necessary since in general the principal stress directions can vary during the response, thus for N samples in the stress history, there will be (N-1)N/2 such pairs or, for N=64K (the number required to accurately resolve the spectrum up to 250 Hz in 0.01 Hz intervals), 2.1 x 10 9 calculations per node each requiring the determination of the roots to a cubic polynomial.

(3)10 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

3. Finite Element Model Description A description of the ANSYS model of the Nine Mile Point Unit 2 steam dryer follows.3.1 Steam Dryer Geometry A geometric representation of the Nine Mile Point Unit 2 steam dryer was developed from available drawings (provided by Constellation Energy Group and included in the design record file, DRF-C-279C) within the Workbench module of ANSYS. The completed model is shown in Figure 1. This model includes on-site modifications to the Nine Mile Point Unit 2 steam dryer.These are as follows.On-Site Modifications (Before 2012 and EPU operation)(i) The top tie rods are replaced with thicker ones.(ii) Inner side plates are replaced with thicker ones.(iii) Middle hoods are reinforced with additional strips.(iv) Lifting rods are reinforced with additional gussets.(v) Per FDDR KG 1-0265 the support conditions are adjusted to ensure that the dryer is supported 100% on the seismic blocks.These additional modifications have been incorporated into the NMP2 steam dryer model and are reflected in the results presented in this report. The affected areas are shown in Figure 2.Modifications Implemented for EPU Operation In [23] several modifications were proposed to meet target EPU stress margins using a previous acoustic loads model (ACM Rev. 4.0) without noise subtraction.

These modifications are now superseded here by the ones below and detailed in Section 5 that are obtained by on the basis of acoustic loads processed using the ACM Rev. 4.1R analysis.

These planned modifications include: (vi) Reinforcement strips are added to the closure plates.(vii) Increase the attachment weld size of the lower-most lifting rod brace from 1/4" to1/22.(viii) Reinforcements to the upper-most and middle lifting rod braces are made in the form of additional strengthening plates.(ix) A 1/8th in curved plate is placed over the middle hood section lying outboard of the closure plate.(x) Four 15 lb masses are added to the central inner hood panels.(xi) Stress relief cut-outs are added to the outer hood/hood support/base plate junctions to alleviate local stresses.(xii) A wrap-around weld is added to the bottom of the drain channel/skirt weld.(xiii) Four 10 lb masses are added to the central middle hood panels.All of the modifications summarized here and detailed in [4], [6] and Section 5 of [24] are implemented in the full steam dryer evaluation produced in Section 5.11 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Modification Implemented During 2014 Refiueling Outage Following the revision to the acoustic loads analysis from ACM 4.1 to ACM 4.IR and the identification of additional signals during power ascension and subsequent collection of data [2], an additional modification was made to the dryer to suppress vibration of the inner side plate spanning the two inner vane banks. The modification consists of a: (xiv) Stiffener with a U-shaped cross-section that is bolted to the inner side of the inner side plate spanning the two inner vane banks. The channel whose upper edge is located 11" below the upper edge of the inner side plate, is described in Appendix B. The evaluation of the bolt stresses is described in [25].A summary of these modifications, references for additional details and how they are implemented in the finite element analysis is given in Table 1.Reference Frame The spatial coordinates used herein to describe the geometry and identify limiting stress locations are expressed in a reference frame whose origin is located at the intersection of the steam dryer centerline and the plane containing the base plates (this plane also contains the top of the upper support ring and the bottom edges of the hoods). The y-axis is parallel to the hoods, the x-axis is normal to the hoods pointing from MSL C/D to MSL A/B, and the z-axis is vertical, positive up.12 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 1. Summary of modifications made to the NMP2 steam dryer.Reinforcement

/ Modification Details FEA Implementation

1. Add reinforcement ribs to all (8) closure plates. Section 5.5 in [5] Closure plates are thickened to obtain dynamically equivalent structure a described in Section 3.5 2. Increase weld of the lowest lifting rod brace/vertical Section 5.4 in [5]; [[plate welds to 0.5" Section 3.5 in [12] (3)]]3. Reinforce middle and upper lifting rod braces to Section 5.1 in [5] Reduce stresses by 0.18 at this location based on eliminate stress concentration on weld to vertical plate. FEA reductions shown for Concept 2 in Table 11 of[26])4. Add 1/8" thick plate over the middle hood section Section 5.2 in [5] Thicken the existing plate by 1/8".lying between the closure plate and existing reinforcement strip.5. Add total of four 15 lb masses to the central sections Section 5.3 in [5] Place 15 lb point masses on the inner hoods at the of the inner hoods. mass centers.6. Add stress relief cut-out at the bottom edge of the Section 5.4 in [5]; [[outer hood supports.

Section 3.4 in [12] (3)]]7. Reinforce the bottom of the drain channel/skirt weld Section 5.4 in [5]; [[with thickened wrap-around weld. Section 3.1 in [12] (3)]]8. Add total of four 10 lb masses to the central sections Section 5.4 in [5] Place 10 lb point masses on the middle hoods at the of the middle hoods. mass centers.9. Add U-section stiffener to inner side plate Appendix B Add shell element model of stiffening using beam connecting inner vane banks. elements to represent attachment bolts.13 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information ANSYS1ln%FBENCH/a zY 50.00 Figure 1. Overall geometry of the Nine Mile Point Unit 2 steam dryer model.14 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 2. Existing on-site pre-EPU modifications accounted for in the model and associated geometrical details.15 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

3.2 Material

Properties The steam dryer is constructed from Type 304 stainless steel and has an operating temperature of 550'F. Properties used in the analysis are summarized below in Table 2.Table 2. Material properties.

Young's Modulus Density Poisson (106 psi) (Ibm/in 3) Ratio stainless steel 25.55 0.284 0.3 structural steel with added water 25.55 0.856 0.3 inertia effect The structural steel modulus is taken from Appendix A of the ASME Code for Type 304 Stainless Steel at an operating temperature 550'F. The effective properties of perforated plates and submerged parts are discussed in Sections 3.4 and 3.6. Note that the increased effective density for submerged components is only used in the harmonic analysis.

When calculating the stress distribution due to the static dead weight load, the unmodified density of steel (0.284 lbmr/in 3) is used throughout.

Inspections of the NMP Unit 2 dryer have revealed IGSCC cracks in the upper support ring (USR) and skirt. A separate analysis of these cracks [27] has been performed to determine whether: (i) they will propagate further into the structure and (ii) their influence upon structural response frequencies and modes must be explicitly accounted for. To establish condition (i) the stress calculated in the global stress analysis is used in conjunction with the crack geometry to calculate the stress intensity factor which is then compared to the threshold stress intensity.

For the USR and skirt cracks the highest stress intensity factors are 1.47 ksi-in 0 5 and 2.75 ksi-in°5 respectively; both values are below the threshold value (3 ksi-in 0 5) implying that fatigue crack growth will not occur.To determine (ii) the change in modal response frequencies due to the presence of a flaw is predicted by analytical means (in the case of the USR) or using finite element analysis (for the skirt). In each case, the flaw size used in these calculations is increased to ensure conservative estimates (for example, in the case of the skirt flaws extending up to 1/22 the panel width are considered).

For the USR, the change in modal frequencies due to the presence of the cracks is less than 0.5%. For the skirt, using a conservative estimate for the crack to panel width of 0.3 (the measured value is less than 0.17) the change in modal frequency is also less than 0.5%. In both cases such small changes in modal frequencies are considered negligible and are readily accounted for when performing frequency shifting.3.3 Model Simplifications The following simplifications were made to achieve reasonable model size while maintaining good modeling fidelity for key structural properties:

• Perforated plates were approximated as continuous plates using modified elastic properties designed to match the static and modal behaviors of the perforated plates. The 16 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information perforated plate structural modeling is summarized in Section 3.4 and Appendix C of[28]." The drying vanes were replaced by point masses attached to the corresponding trough bottom plates and vane bank top covers (Figure 4). The bounding perforated plates, vane bank end plates, and vane bank top covers were explicitly modeled (see Section 3.5)." The added mass properties of the lower part of the skirt below the reactor water level were obtained using a separate hydrodynamic analysis (see Section 3.6).(3)1]" Four steam dryer support brackets that are located on the reactor vessel and spaced at 900 intervals were explicitly modeled (see Section 3.9)." Most welds were replaced by node-to-node connections; interconnected parts share common nodes along the welds. In other locations the constraint equations between nodal degrees of freedom were introduced as described in Section 3.9.3.4 Perforated Plate Model The perforated plates were modeled as solid plates with adjusted elastic and dynamic properties.

Properties of the perforated plates were assigned according to the type and size of perforation.

Based on [29], for an equilateral square pattern with given hole size and spacing, the effective moduli of elasticity were found.The adjusted properties for the perforated plates are shown in Table 3 as ratios to material properties of structural steel, provided in Table 2. Locations of perforated plates are classified by steam entry / exit vane bank side and vertical position.Tests were carried out to verify that this representation of perforated plates by continuous ones with modified elastic properties preserves the modal properties of the structure.

These tests are summarized in Appendix C of [28] and compare the predicted first modal frequency for a cantilevered perforated plate against an experimentally measured value. The prediction was obtained for 40% and 13% open area plates (these are representative of the largest and lowest open area ratios of the perforated plates at NMP2, as seen in Table 3) using the analytical formula for a cantilevered plate and the modified Young's modulus and Poisson's ratio given by O'Donnell

[29]. The measured and predicted frequencies are in close agreement, differing by less than 3%.(3)17 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

[1 (3)]][1 (3)]]Figure 3. I (3)18 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 3. Material properties of perforated plates.I (3)3.5 Vane Bank Model The vane bank assemblies consist of many vertical angled plates that are computationally expensive to model explicitly, since a prohibitive number of elements would be required.

These parts have significant weight, which is transmitted through the surrounding structure, so it is important to capture their gross inertial properties.

Here the vane banks are modeled as a collection of point masses located at the center of mass for each vane bank section (Figure 4).The following masses were used for the vane bank sections, based on data found on provided drawings: inner banks, 1618 Ibm, 4 sections per bank;middle banks, 1485 Ibm, total 4 sections per bank; and outer banks, 1550 Ibm, 3 sections per bank.These masses were applied to the base plates and vane top covers using the standard ANSYS point mass modeling option, element MASS2 1. ANSYS automatically distributes the point mass inertial loads to the nodes of the selected structure.

The distribution algorithm minimizes the sum of the squares of the nodal inertial forces, while ensuring that the net forces and moments are conserved.

Vane banks are not exposed to main steam lines directly, but rather shielded by the hoods.The collective stiffness of the vane banks is expected to be small compared to the surrounding support structure and is neglected in the model. In the static case it is reasonable to expect that this constitutes a conservative approach, since neglecting the stiffness of the vane banks implies that the entire weight is transmitted through the adjacent vane bank walls and supports.

In the dynamic case the vane banks exhibit only a weak response since (i) they have large inertia so that the characteristic acoustically-induced forces divided by the vane masses and inertias yield small amplitude motions, velocities and accelerations; and (ii) they are shielded from acoustic loads by the hoods, which transfer dynamic loads to the rest of the structure.

Thus, compared to the hoods, less motion is anticipated on the vane banks so that 19 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information approximating their inertial properties with equivalent point masses is justified.

Nevertheless, the bounding parts, such as perforated plates, side panels, and top covers, are retained in the model. Errors associated with the point mass representation of the vane banks are compensated for by frequency shifting of the applied loads.3.6 Water Inertia Effect on Submerged Panels Water inertia was modeled by an increase in density of the submerged structure to account for the added hydrodynamic mass. This added mass was found by a separate hydrodynamic analysis (included in DRF-C-279C supporting this report) to be 0.143 Ibm/in 2 on the submerged skirt area. This is modeled by effectively increasing the material density for the submerged portions of the skirt. Since the skirt is 0.25 inches thick, the added mass is equivalent to a density increase by 0.572 lbm/in 3.This added water mass was included in the ANSYS model by appropriately modifying the density of the submerged structural elements when computing harmonic response.

For the static stresses, the unmodified density of steel is used throughout.

3.7 Structural

Damping Structural damping was defined as 1% of critical damping for all frequencies.

This damping is consistent with guidance given on pg. 10 of NRC RG-1.20 [33].3.8 Mesh Details and Element Types Shell elements were employed to model the skirt, hoods, perforated plates, side and end plates, trough bottom plates, reinforcements, base plates and cover plates. Specifically, the four-node, Shell Element SHELL63, was selected to model these structural components.

This element models bending and membrane stresses, but omits transverse shear. The use of shell elements is appropriate for most of the structure where the characteristic thickness is small compared to the other plate dimensions.

For thicker structures, such as the upper and lower support rings, solid brick elements were used to provide the full 3D stress. The elements SURF154 are used to assure proper application of pressure loading to the structure.

Mesh details and element types are shown Table 4 and Table 5.The mesh is generated automatically by ANSYS with refinement near edges. The maximum allowable mesh spacing is specified by the user. Here a 2.5 inch maximum allowable spacing is specified with refinement up to 1.5 inch in the following areas: drain pipes, tie rods, the curved portions of the drain channels and the hoods. Details of the finite element mesh for the baseline model (i.e., without the U-section stiffeners) are shown in Figure 5. Numerical experiments carried out using the ANSYS code applied to simple analytically tractable plate structures with dimensions and mesh spacings similar to the ones used for the steam dryer, confirm that the natural frequencies are accurately recovered (less than I % errors for the first modes). These errors are compensated for by the use of frequency shifting.The baseline analysis is carried out without the U-section stiffener.

A supplemental model is also developed with the U-section stiffener represented using shell elements and connected to the inner side plate as shown in Figure 25 using beam elements to represent the bolts. This model is described further in Appendix B.20 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

3.9 Connections

between Structural Components Most connections between parts are modeled as node-to-node connections.

This is the correct manner (i.e., within the finite element framework) of joining elements away from discontinuities.

At joints between shells, this approach omits the additional stiffness provided by the extra weld material.

Also, locally 3D effects are more pronounced.

The latter effect is accounted for using weld factors. The deviation in stiffness due to weld material is negligible, since weld dimensions are on the order of the shell thickness.

The consequences upon modal frequencies and amplitude are, to first order, proportional to t/L where t is the thickness and L a characteristic shell length. The errors committed by ignoring additional weld stiffness are thus small and readily compensated for by performing frequency shifts.When joining shell and solid elements, however, the problem arises of properly constraining the rotations, since shell element nodes contain both displacement and rotational degrees of freedom at every node whereas solid elements model only the translations.

A node-to-node connection would effectively appear to the shell element as a simply supported, rather than (the correct) cantilevered restraint and significantly alter the dynamic response of the shell structure.

To address this problem, constraint equations are used to properly connect adjacent shell- and solid-element modeled structures.

Basically, all such constraints express the deflection (and rotation for shell elements) of a node, R 1 , on one structural component in terms of the deflections/rotations of the corresponding point, P 2 , on the other connected component.

Specifically, the element containing P 2 is identified and the deformations at P 2 determined by interpolation between the element nodes. The following types of shell-solid element connections are used in the steam dryer model including the following:

1. Connections of shell faces to solid faces (Figure 6a). While only displacement degrees of freedom are explicitly constrained, this approach also implicitly constrains the rotational degrees of freedom when multiple shell nodes on a sufficiently dense grid are connected to the same solid face.2. Connections of shell edges to solids (e.g., connection of the bottom of closure plates with the upper ring). Since solid elements do not have rotational degrees of freedom, the coupling approach consisted of having the shell penetrate into the solid by one shell thickness and then constraining both the embedded shell element nodes (inside the solid)and the ones located on the surface of the solid structure (see Figure 6b). Numerical tests involving simple structures showed that this approach and penetration depth reproduce both the deflections and stresses of the same structure modeled using only solid elements or ANSYS' bonded contact technology.

Continuity of rotations and displacements is achieved.The use of constraint conditions rather than the bonded contacts advocated by ANSYS for connecting independently meshed structural components confers better accuracy and useful numerical advantages to the structural analysis of the steam dryer including better conditioned and smaller matrices.

The smaller size results from the fact that equations and degrees of freedom are eliminated rather than augmented (in Lagrange multiplier-based methods) by additional degrees of freedom. Also, the implementation of contact elements relies on the use of very high stiffness elements (in penalty function-based implementations) or results in indefinite 21 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information matrices (Lagrange multiplier implementations) with poorer convergence behavior compared to positive definite matrices.The steam dryer rests on four support blocks which resist vertical and lateral displacement.

The support blocks contact the seismic blocks welded to the USR so that 100% of the dryer weight is transmitted through the seismic blocks per the FDDR KG1 -265. Because the contact region between the blocks and steam dryer is small, the seismic blocks are considered free to rotate about the radial axis. Specifically nodal constraints (zero relative displacement) are imposed over the contact area between the seismic blocks and the support blocks. Two nodes on each support block are fixed as indicated in Figure 7. One node is at the center of the support block surface facing the vessel and the other node is 0.5" offset inside the block towards the steam dryer, half way to the nearest upper support ring node. This arrangement approximates the nonlinear contact condition where the ring can tip about the block.For the finite element model with the U-section stiffeners on the inner side plates, 5 short beam elements per beam are added to represent the bolts that attach it to the side plate. In addition constraints to prevent separation and relative rotation between the side plate and U-section stiffener surfaces that are in common contact, are imposed. The U-channel itself is represented using shell elements whose nodal positions and element connectivities are 'lofted'(copied and translated) from the underlying side plate nodes to ensure a direct one-to-one correspondence between the nodes in the channel and side plate. This facilitates the imposition of aforementioned constraints between the channel and side plate shell elements.22 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information CE Masses are connected to top and bottom supports Point masses Gussets to lifting'rods connections Skirt to support-, rings connections Simply supported restraints A ,/Figure 4. Point masses representing the vanes. The pink shading represents where constraint equations between nodes are applied (generally between solid and shell elements, point masses and nodes and (3)).23 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 4. FE Model Summary.Description Description Quantity Total Nodes I baseline model 159,793 with U-section stiffener 189,862 Total Elements baseline model 124,496 with U-section stiffener 155,680 1. Not including additional damper nodes and elements.Table 5. Listing of Element Types.Generic Element Type Name Element Name 20-Node Quadratic Hexahedron SOLID 186 10-Node Quadratic Tetrahedron SOLID187 4-Node Elastic Shell SHELL63 2-Node Beam element (only used for U-section stiffener BEAM188 bolts)Mass Element MASS21 Pressure Surface Definition SURF154 Damper element COMBIN14 24 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5a. Mesh overview.25 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5b. Close up of mesh showing on-site modifications.

26 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5c. Close up of mesh showing drain pipes and hood supports.27 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5d. Close up of mesh showing node-to-node connections between various plates.28 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5e. Close up of mesh showing node-to-node connections between the skirt and drain channels; hood supports and hoods; and other parts.29 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5f. Close up view of tie bars.30 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Shell nodes DOF are related to solid element shape functions Surface of solid element Figure 6a. Face-to-face shell to solid connection.

Shell nodes DOF are related to solid element shape functions Surface of solid element Figure 6b. Shell edge-to-solid face connection.

31 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 7. Boundary conditions.

Inside node is half way between outer surface of support block and upper support ring.32 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 3.10 Pressure Loading The harmonic loads are produced by the pressures acting on the exposed surfaces of the steam dryer. At every frequency and for each MSL, the pressure distribution corresponding to a unit pressure at the MSL inlet is represented on a three-inch grid lattice grid (i.e., a mesh whose lines are aligned with the x-, y- and z-directions) that is superimposed over the steam dryer surface. This grid is compatible with the 'Table' format used by ANSYS to 'paint' general pressure distributions upon structural surfaces.

The pressures are obtained from the Helmholtz solver routine in the acoustic analysis [14].In general, the lattice nodes do not lie on the surface, so that to obtain the pressure differences at the surface it is necessary to interpolate the pressure differences stored at the lattice nodes. This is done using simple linear interpolation between the 8 forming nodes of the lattice cell containing the surface point of interest.

Inspection of the resulting pressures at selected nodes shows that these pressures vary in a well-behaved manner between the nodes with prescribed pressures.

Graphical depictions of the resulting pressures and comparisons between the peak pressures in the original nodal histories and those in the final surface load distributions produced in ANSYS, all confirm that the load data are interpolated accurately and transferred correctly to ANSYS.The harmonic pressure loads are only applied to surfaces above the water level, as indicated in Figure 8. In addition to the pressure load, the static loading induced by the weight of the steam dryer is analyzed separately.

The resulting static and harmonic stresses are linearly combined to obtain total values which are then processed to calculate maximum and altemating stress intensities for assessment in Section 5.I (3) This is useful since revisions in the loads model do not necessitate recalculation of the unit stresses.33 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information NODE S NF WW PRE S-NORK.349671 .749872-.212228 .612429 .887315 Figure 8a. Real part of unit pressure loading MSL A (in psid) on the steam dryer at 50.1 Hz. No loading is applied to the submerged surface and lifting rods.34 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information ODZ 6 UF U W PRE S-NORM-. 490178 28 .831766 Figure 8b. Real part of unit pressure loading MSL A (in psid) on the steam dryer at 200.45 Hz.No loading is applied to the submerged surface and lifting rods.35 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 3.11 [[Continuum Dynamics, Inc. Proprietary Information

4. Structural Analysis The solution is decomposed into static and harmonic parts, where the static solution produces the stress field induced by the supported structure subjected to its own weight and the harmonic solution accounts for the harmonic stress field due to the unit pressure of given frequency in one of the main steam lines. All solutions are linearly combined, with amplitudes provided by signal measurements in each steam line, to obtain the final displacement and stress time histories.

This decomposition facilitates the prescription of the added mass model accounting for hydrodynamic interaction and allows one to compare the stress contributions arising from static and harmonic loads separately.

Proper evaluation of the maximum membrane and membrane+bending stresses requires that the static loads due to weight be accounted for. Hence both static and harmonic analyses are carried out.4.1 Static Analysis The results of the static analysis are shown in Figure 9. The locations with highest stress include the inner vane bank connection to inner base plate near support brackets with stress intensity 9,598 psi. There are four locations with artificial stress singularity, which are excluded from the analysis.

The static stresses one node away are used at these locations as more realistic estimate of local stress. These locations are at the connections of the inner end plate to the inner base plate at the ends of the cut-out, as shown in Figure 9c.4.2 Harmonic Analysis The harmonic pressure loads were applied to the structural model at all surface nodes described in Section 3.10. Typical stress intensity distributions over the structure are shown in Figure 10. Stresses were calculated for each frequency, and results from static and harmonic calculations were combined.To evaluate maximum stresses, the stress harmonics including the static component are transformed into a time history using FFT, and the maximum and alternating stress intensities for the response, evaluated.

According to ASME B&PV Code,Section III, Subsection NG-3216.2 the following procedure was established to calculate alternating, stresses.

For every node, the stress difference tensors, Fmnm = (5n -Um, are considered for all possible pairs of the stresses ca and cym at different time levels, t,, and tmn. Note that all possible pairs require consideration since there are no "obvious" extrema in the stress responses.

However, in order to contain computational cost, extensive screening of the pairs takes place (see Section 2.3) so that pairs known to produce alternating stress intensities less than 250 psi are rejected.

For each remaining stress difference tensor, the principal stresses S 1 , S 2 , S3 are computed and the maximum absolute value among principal stress differences, Snm=max{ISI-S 2 1, SI-S 3 1,1S 2-S31}, obtained.

The alternating stress at the node is then one-half the maximum value of Snm taken over all combinations (n,m), i.e., salt = max IS. }. This alternating stress is compared against allowable n,m values, depending on the node location with respect to welds.37 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information NODAL SOLUTION STEP=1 SUB =1 TIME=1 USUM (AVG)RSYS=O DMX =.068847 SMN =.505E-03 SMX =.068847 SO50 E-03 .0 8 9 0 15 jll ý6067 .0 36 061254 .0 8 4.008099*056.687 Figure 9a. Overview of static calculations showing displacements (in inches). Maximum displacement (DMX) is 0.069". Note that displacements are amplified for visualization.

38 This Document Does Not Contain IContain uu Dynamics Inc. Proprietary Informnation.Lmwm6 1 AN Figure 9b. O verview ofm smtC l lto inten sty ) is 9,598 psi. N o0Ss] ote tha t disp Sho~we ing stress intensities (in psi). Stres Placement are amplified for Psli), um stress Sualization 39 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 7 Figure 9c. Stress singularities.

Model is shown in wireframe mode for clarity. NSHotSpots represents the node selection of the stress singularities.

40 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information NODAL SOLUTION STEPw377 SUBD -1 FREQ-50.097 REAL ONLY SINT (VCG)DUX -.083772 SWN -.299472 SX =)11002'ii 3 '4444 555. 556 )3889 5000 Figure 10a. Overview of harmonic calculations showing real part of stress intensities (in psi)along with displacements.

Unit loading MSL A at 50.1 Hz (oriented to show high stress locations at the hoods).41 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information AN NODAL SOLUTION I STEP-337 SUB =1 FREQ-200.446 REAL ONLY SINT (AVG)DEC =.014L39 SN u.219769 SAX =13006 3000 Figure 10b. Overview of harmonic calculations showing real part of stress intensities (in psi)along with displacements.

Unit loading MSL A at 200.5 Hz.42 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 4.3 Post-Processing The static and transient stresses computed at every node with ANSYS were exported into files for subsequent post-processing.

These files were then read into separate customized software to compute the maximum and alternating stresses at every node. The maximum stress was defined for each node as the largest stress intensity occurring during the time history.Alternating stresses were calculated according to the ASME standard described above. For shell elements the maximum stresses were calculated separately at the mid-plane, where only membrane stress is present, and at top/bottom of the shell, where bending stresses are also present.For nodes that are shared between several structural components or lie on junctions, the maximum and alternating stress intensities are calculated as follows. First, the nodal stress tensor is computed separately for each individual component by averaging over all finite elements meeting at the node and belonging to the same structural component.

The time histories of these stress tensors are then processed to deduce the maximum and alternating stress intensities for each structural component.

Finally for nodes shared across multiple components the highest of the component-wise maximum and alternating stresses is recorded as the "nodal" stress. This approach prevents averaging of stresses across components and thus yields conservative estimates for nodal stresses at the weld locations where several components are joined together.The maximum stresses are compared against allowable values which depend upon the stress type (membrane, membrane+bending, alternating

-Pm, Pm+Pb, Salt) and location (at a weld or away from welds). These allowables are specified in the following section. For solid elements the most conservative allowable for membrane stress, Pm, is used, although bending stresses are nearly always present also. The structure is then assessed in terms of stress ratios formed by dividing allowables by the computed stresses at every node. Stress ratios less than unity imply that the associated maximum and/or alternating stress intensities exceed the allowable levels.Post-processing tools calculate the stress ratios, identifying the nodes with low stress ratios and generating files formatted for input to the 3D graphics program, TecPlot, which provides more general and sophisticated plotting options than currently available in ANSYS.4.4 Computation of Stress Ratios for Structural Assessment The ASME B&PV Code,Section III, subsection NG provides different allowable stresses for different load combinations and plant conditions.

The stress levels of interest in this analysis are for the normal operating condition, which is the Level A service condition.

The load combination for this condition is: Normal Operating Load Combination

= Weight + Pressure + Thermal The weight and fluctuating pressure contributions have been calculated in this analysis and are included in the stress results. The static pressure differences and thermal expansion stresses are small, since the entire steam dryer is suspended inside the reactor vessel and all surfaces are exposed to the same conditions.

Seismic loads only occur in Level B and C cases, and are not considered in this analysis.43 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Allowable Stress Intensities The ASME B&PV Code,Section III, subsection NG shows the following (Table 6) for the maximum allowable stress intensity (Sm) and alternating stress intensity (Sa) for the Level A service condition.

The allowable stress intensity values for type 304 stainless steel at operating temperature 550'F are taken from Table 1-1.2 and Fig. 1-9.2.2 of Appendix I of Section III, in the ASME B&PV Code. The calculation for different stress categories is performed in accordance with Fig. NG-3221-1 of Division I, Section I1, subsection NG. The allowable value for alternating stress is taken from curve C of Fig. 1-9.2.2 in Appendix I in Section III of the ASME B&PV Code.Table 6. Maximum Allowable Stress Intensity and Alternating Stress Intensity for all areas other than welds. The notation Pm represents membrane stress; Pb represents stress due to bending; Q represents secondary stresses (from thermal effects and gross structural discontinuities, for example);

and F represents additional stress increments (due to local structural discontinuities, for example).Type Notation Service Limit Allowable Value (ksi)Maximum Stress Allowables:

General Membrane Pm Sm 16.9 Membrane + Bending Pm + Pb 1.5 Sm 25.35 Primary + Secondary Pm + Pb + Q 3.0 Sm 50.7 Alternating Stress Allowable:

Peak = Primary + Secondary

+ F Salt Sa 13.6 When evaluating welds, either the calculated or allowable stress was adjusted, to account for stress concentration factor and weld quality. Specifically: " For maximum allowable stress intensity, the allowable value is decreased by multiplying its value in Table 6 by 0.55.* For alternating stress intensity, the calculated weld stress intensity is multiplied by a weld stress intensity (fatigue) factor of 1.8 for a fillet weld and 1.4 for a full penetration weld, before comparison to the Sa value given above.The weld factors of 0.55 and 1.4 (full penetration weld) or 1.8 (fillet weld) were selected based on the observable quality of the shop welds and liquid penetrant NDE testing of all welds (excluding tack and intermittent welds, which were subject to 5X visual inspection) during fabrication.

These factors are consistent with fatigue strength reduction factors recommended by the Welding Research Council, [34], and stress concentration factors at welds, provided in [35]and [36]. In addition, critical welds are subject to periodical visual inspections in accordance with the requirements of GE SIL 644 SIL and BWR VIP-139 [37]. Therefore, for weld stress intensities, the allowable values are shown in Table 7. These factors (0.55 and 1.4 or 1.8) also conservatively presume that the structure is joined using fillet welds unless specified otherwise.

Since fillet welds correspond to larger stress concentration factors than other types of welds, this assumption is a conservative one.44 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 7. Weld Stress Intensities.

Type Notation Service Limit Allowable Value (ksi)Maximum Stress Allowables:

General Membrane Pm 0.55 Sm 9.30 Membrane + Bending Pm + Pb 0.825 Sm 13.94 Primary + Secondary Pm + Pb + Q 1.65 Sm 27.89 Alternating Stress Allowables:

Peak = Primary + Secondary

+ F Salt Sa 13.6 Comparison of Calculated and Allowable Stress Intensities The classification of stresses into general membrane or membrane + bending types was made according to the exact location, where the stress intensity was calculated; namely, general membrane, Pm, for middle surface of shell element, and membrane + bending, Pm + Pb, for other locations.

For solid elements the most conservative, general membrane, Pm, allowable is used.The structural assessment is carried out by computing stress ratios between the computed maximum and alternating stress intensities, and the allowable levels. Locations where any of the stresses exceed allowable levels will have stress ratios less than unity. Since computation of stress ratios and related quantities within ANSYS is time-consuming and awkward, a separate FORTRAN code was developed to compute the necessary maximum and alternating stress intensities, Pm, Pm+Pb, and Salt, and then compare it to allowables.

Specifically, the following quantities were computed at every node: 1. The maximum membrane stress intensity, Pm (evaluated at the mid-thickness location for shells), 2. The maximum membrane+bending stress intensity, Pm+Pb, (taken as the largest of the maximum stress intensity values at the bottom, top, and mid thickness locations, for shells), 3. The alternating stress, Salt, (the maximum value over the three thickness locations is taken).4. The stress ratio due to a maximum stress intensity assuming the node lies at a non-weld location (note that this is the minimum ratio obtained considering both membrane stresses and membrane+bending stresses):

SR-P(nw) = min{ Sm/Pm, 1.5

• Sm/(Pm+Pb)

}.5. The alternating stress ratio assuming the node lies at a non-weld location, SR-a(nw) = Sa/(I.1

• Salt), 6. The same as 4, but assuming the node lies on a weld, SR-P(w)=SR-P(nw)

• 0.55 7. The same as 5, but assuming the node lies on a weld, SR-a(w)=SR-a(nw)

/ fsw.45 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Note that in steps 4 and 6, the minimum of the stress ratios based on Pm and Pm+Pb, is taken.The allowables listed in Table 7, Sm=16,900 psi and Sa=13,600 psi. The factors, 0.55 and fsw, are the weld factors discussed above with fsw=1.8 being appropriate for a fillet weld and fsw=1.4 for a full penetration weld. The factor of 1.1 accounts for the differences in Young's moduli for the steel used in the steam dryer and the values assumed in alternating stress allowable.

According to NG-3222.4 in subsection NG of Section III of the ASME Code [7], the effect of elastic modulus upon alternating stresses is taken into account by multiplying alternating stress Salt at all locations by the ratio, E/Emodel=

1.1, where: E = 28.3 106 psi, as shown on Fig. 1-9.2.2. ASME BP&V Code Emodel = 25.55 106 psi (Table 2)The appropriate maximum and alternating stress ratios, SR-P and SR-a, are thus determined and a final listing of nodes having the smallest stress ratios is generated.

The nodes with stress ratios lower than 4 are plotted in TecPlot (a 3D graphics plotting program widely used in engineering communities

[38]). These nodes are tabulated and depicted in the following Results Section.(3)]](*3)j 46 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]47 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]48 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]49 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]50 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

(3)51 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]52 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

5. Results The stress intensities and associated stress ratios resulting from the Rev. 4.1 R acoustic/hydrodynamic loads [2, 8, 9] with associated biases and uncertainties factored in, are presented below. The acoustic loads applied to the steam dryer are obtained using the most recent and complete strain gage signals [2] and processed using the ACM Rev. 4. 1R analysis with associated biases and uncertainties updated to reflect the new revision as described in [2, 8, 9]. For the FEM structural model there are three main contributors to the bias and uncertainty.

The first is an uncertainty (21.5%) that accounts for modeling idealizations (e.g., vane bank mass model), geometrical approximations and other discrepancies between the modeled and actual dryer such as neglecting of weld mass and stiffness in the FEA. The second contributor is a bias of 9.53% accounting for discretization errors associated with using a finite size mesh, upon computed stresses.

The third contributor is also a bias and compensates for the use of a finite discretization schedule in the construction of the unit solutions.

The frequencies are spaced such that at 1% damping the maximum (worst case) error in a resonance peak is 5%. The average error for this frequency schedule is 1.72%.Results are presented with frequency shifting included.

Unless specified otherwise, frequency shifts are generally performed at 2.5% increments.

The tabulated stresses and stress ratios are obtained using a 'blanking' procedure that is designed to prevent reporting a large number of high stress nodes from essentially the same location on the structure.

In the case of stress intensities this procedure is as follows. The relevant stress intensities are first computed at every node and then nodes sorted according to stress level.. The highest stress node is noted and all neighboring nodes within 10 inches of the highest stress node and its symmetric images (i.e., reflections across the x=O and y=O planes) are "blanked" (i.e., excluded from the search for subsequent high stress locations).

Of the remaining nodes, the next highest stress node is identified and its neighbors (closer than 10 inches) blanked. The third highest stress node is similarly located and the search continued in this fashion until all nodes are either blanked or have stresses less than half the highest value on the structure.

For stress ratios, an analogous blanking procedure is applied. Thus the lowest stress ratio of a particular type in a 10" neighborhood and its symmetric images is identified and all other nodes in these regions excluded from listing in the table. Of the remaining nodes, the one with the lowest stress ratio is reported and its neighboring points similarly excluded, and so on until all nodes are either blanked or have a stress ratio higher than 5. The set of points thus obtained is referred to as a reduced point set (RPS).As described in Section 1.1, the large number of combinations of loads, bias and uncertainties selections, and dryer models necessitates a tailored analysis combining full and real time stress analyses.

The following sections are structured and ordered as follows. Section 5.1 lists the limiting stress ratios obtained from a full steam dryer evaluation using the baseline EPU load (i.e., using the bias and uncertainty values pertaining to ACM 4.1R and without considering the drain trap out of service or RCIC valve closed conditions) applied to the baseline dryer geometry (i.e., without the numerical corrections developed in Section 5.3). This calculation is used to develop a list of nodes for real time analysis in Section 5.2 and to motivate analytical refinements for reanalysis in Section 5.3. The full stress evaluation is repeated using the same baseline EPU load and the refined structural model; the results from this evaluation are given in 53 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Sections 5.4 to 5.6. The highest maximum and alternating stress intensities indicating which points on the dryer experience significant stress concentration and/or modal response, are presented in Section 5.4. The lowest stress ratios obtained by comparing the stresses against allowable values, accounting for stress type (maximum and alternating) and location (on or away from a weld), are reported Section 5.5. The frequency dependence of the stresses at nodes experiencing the lowest stress ratios is depicted in the form of accumulative PSDs in Section 5.6.Real time assessments of the other loads and total uncertainty combinations using the list developed in Section 5.2 are summarized in Section 5.7. These assessments are carried out without the U-section stiffener which allows reuse of existing unit solutions and post-processing utilities (since the mesh indexing did not change). While a finite element model of the dryer with the U-section stiffener added was developed, schedule requirements did not permit generating the associated unit solutions over the entire 0-250 Hz frequency range or developing the mesh specific post-processing tools (e.g., assignment of SRFs, handling the added beam elements representing the U-section stiffener attachment bolts, etc.). The stress evaluation with the U-section stiffener included is therefore carried out using real time analysis as described in Section 5.8.5.1. Preliminary Stress Assessment of Normal EPU Load + Base Dryer Model The first stress evaluation performed serves as a preliminary assessment and screening of low stress ratio locations and processed all steam dryer nodes using the baseline or normal EPU load (i.e., drain trap out-of-service and RCIC valve closed loads were not considered) adjusted in accordance with the ACM 4.1 R bias and uncertainty values [2]. The steam dryer model does not include the U-section stiffener described in Appendix B nor the reanalysis options in Section 5.3.It also employs a stress reduction factor of 0.77 at locations 2, 8 and 14 in Table 9c. This SRF is adopted from a similar location on the outer vane bank. In the subsequent analysis in Section 5.3 and also the real time evaluations it is replaced by the SRFs of 0.66 and 0.80 developed in Section 4.5 under embedded model 6 which pertains to this exact junction location.

Note too that junctions involving the lifting rods and upper lifting rod braces were excluded from the evaluation in this Section 5.1 while awaiting resolution as to whether or not these junctions are welded. It ultimately became clear that while the original drawings indicated a weld, the as-built configuration does not, so that no weld factor is applicable at this particular junction.

Again, in the detailed evaluation of Sections 5.4 to 5.6 and the real time assessments, these locations are retained.

The purpose of this evaluation is mainly to identify locations that have alternating stress ratios, SR-a<2 or are likely to drop below 2.0 at other load conditions.

Locations with low stress ratios are used to assemble the list of nodes in Section 5.2 for use in subsequent real time processing in Sections 5.7 and 5.8. Additionally scrutiny of the lowest stress ratio sites led to the analytical analysis methods in Section 5.3, the construction of the embedded model 6 and associated SRFs in Section 4.5 and Table 8, and the U-section stiffener in Appendix B.At zero frequency shift the limiting peak and alternating stress intensities are SR-P=I. 18 and SR-a=l.87.

The effects of frequency shifts are conservatively accounted for by identifying the minimum stress ratio at every node, where the minimum is taken over all the frequency shifts considered (including the nominal or 0% shift case). The stress ratios computed for EPU with frequency shifting included are listed in Table 9. The stress ratios are grouped according to type 54 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (SR-P for maximum membrane and membrane+bending stress, SR-a for alternating stress) and location (away from welds or on a weld). The tabulated nodes with the smallest alternating stress ratios in Table 9c are also depicted in Figure 13 together with all nodes with SR-a<5 (for brevity, comparable plots of the other stress ratio types are not listed here as they essentially reappear in the full analysis summarized in Section 5.5).With frequency shifting included, the lowest stress ratio, SR-P=1.14, occurs at the same location as in the nominal case and is only slightly lower. This value is dominated by the static contribution and so is not very sensitive to frequency shifting.

Three locations in Table 9c are found to have stress ratios below 2.0. Close up views of these locations are depicted in Figure 14. The lowest alternating stress ratio occurs at the top of the closure plate/middle hood weld (see location 1 in Figure 13a, and Figure 14a) and assumes a value of SR-a=l.68.

Based on the largest Fourier coefficient, the dominant frequency contributing to this stress is 86.8 Hz. The second location involves the common junction of the inner vane bank top and side plates, the inner side plate spanning the inner vane banks, and the inner vane bank perforated plate (see Figure 14b). Its stress ratio is 1.75 after application of an SRF of 0.77. This stress ratio is revised in the subsequent sections using the methods in Section 5.3. While the recommended margin can be achieved using a combination of the methods summarized in Sections 5.3.1 and 5.3.2, uncertainty over the applicability of the stress interpolation procedure along the weld lines at this complex geometry junction involving multiple welds motivated consideration of alternate concepts for meeting margin. This resulted in the implementation of the U-section stiffener to directly suppress the resonant response of the inner side plate. The third location with SR-a less than 2 occurs near the end of the upper lifting rod brace on the outer vane bank side plate/closure plate junction and has SR-a=I.80 (see Figure 14c). Comparison with the other three upper braces at the reflected locations shows that their alternating stress ratios are all well above 2.0 and also that the grid is more regular. This motivated a re-evaluation of the local stress using a better quality mesh as described in Section 5.3. Virtually all of the limiting stress ratios occur at either the +7.5% or +10% shifts.55 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 9a. Limiting non-weld locations with at EPU conditions with frequency shifts. Stress ratios are grouped according to stress type (maximum -SR-P; or alternating

-SR-a).Stress Location Location (in.) node Stress Intensity (psi) Stress Ratio % Freq. Dom.Ratio x y z Pm Pm+Pb Salt SR-P SR-a Shift Freq. (Hz)SR-P 1. Inner Side Plate 3.1 119 0.5 37229 6685 8548 2281 2.53 5.42 5 123.5 2. Support/Seismic Block 10.2 123.8 -9.5 113286 6210 6210 4314 2.72 2.87 10 13.7 3. Thin Vane Bank Plate -15.6 -118.4 0.6 2558 4311 4727 833 3.92 14.84 10 16.0 SR-a 1. Support/Seismic Block 10.2 123.8 -9.5 113286 6210 6210 4314 2.72 2.87 10 13.7 2. Inner Side Plate 14.4 -119 88 37592 764 5055 4235 5.02 2.92 7.5 78.7 3. Side Plate -79.4 -85.2 76.8 10819 436 3463 3122 7.32 3.96 10 17.9 4. Hood Support 89 -28.4 0 14474 5050 5210 3068 3.35 4.03 5 82.0 56 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 9b. Limiting peak stress ratios, SR-P, on welds at EPU conditions with frequency shifts. Bold text indicates minimum stress ratio on the structure.

Location SRF(a) Location (in.) node Stress Intensity (psi) Stress Ratio % Freq. Doam.x y z Pm Pm+Pb Salt SR-P SR-a Shift Freq. (Hz)1. USR/Support/Seismic Block -6.9 -122.3 -9.5 113554 8133 8133 3048 1.14 2.25 5 13.7 2. Middle Base Plate/Inner Backing Bar Out/Inner 39.9 108.6 0 85631 1478 10031 2750 1.39 2.5 10 82.6 Backing Bar/Inner Hood 3. Side Plate Ext/Inner Base Plate 16.3 119 0 94143 6396 9246 1525 1.45 4.51 10 39.6 4. Hood Support/Outer Base Plate/Middle

-71.3 0 0 95428 6159 6227 2634 1.51 2.61 7.5 14.4 Backing Bar 5. Inner Side Plate/Inner Base Plate -2.3 -119 0 99200 4026 8778 3167 1.59 2.17 7.5 123.5 6. Tie Bar(2) 0.77 -49.3 -108.1 88 143795 5644 5644 1773 1.65 3.87 10 82.6 7. Thin Vane Bank Plate/Hood Support/Inner 24.1 -59.5 0 85191 5245 5330 2030 1.77 3.38 10 13.3 Base Plate 8. Thin Vane Bank Plate/Hood Support/Middle 55.6 -54.3 0 98968 5241 5318 2232 1.77 3.08 10 82 Base Plate 9. Hood Support/Outer Cover Plate/Outer Hood(4) 0.8 -102.8 28.4 0 95267 5054 5081 2996 1.84 2.29 10 13.7 10. Hood Support/Middle Base Plate/Inner

-39.9 59.5 0 90468 5030 5165 1843 1.85 3.73 10 82.6 Backing Bar/Inner Hood(b)11. Upper Support Ring/Seismic Block/Support

-122.1 10.2 -9.5 113508 4769 4769 2060 1.95 3.33 10 16.1 12. Outer Cover Plate/Outer Hood 102.8 -58.1 0 94498 1252 6440 1422 2.17 4.83 7.5 82 13. Closure Plate/Middle Hood 60.2 -85.2 87 89317 1459 6286 4085 2.22 1.68 10 86.8 Notes: (a) , Inc. Proprietary Information Figure 13b. Locations of minimum alternating stress ratios, SR-a_<5, at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-a values at welds in Table 9c. View showing locations 5 and 15.60 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information x Figure 13c. Locations of minimum alternating stress ratios, SR-a.<5, at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-a values at welds in Table 9c. View around locations 9, 10, 12 and 15.61 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14a. Close up view of Entry 1 in 9c at the top of the closure plate/middle hood weld.62 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14b. Close up view of Entry 2 Table 9c involving the components near the top of the inner vane bank and tie bar.63 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14c. Close up view of Entry 3 Table 9c involving the perimeter of the reinforcement plate added to the lifting rod restraint bracket.64 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

5.2 Construction

of Real Time Node List As discussed in [2] other load conditions occurring when the drain trap is out of service and when the RCIC valve is closed, also require attention.

Moreover, it has been noted [2] that while using the bias and uncertainties developed for ACM 4. 1R are technically appropriate, it is also of value to consider a hybrid and conservative set of biases and uncertainties that retains the higher of the ACM 4.1 and ACM 4.IR total uncertainties over each frequency interval.

To expedite analysis of these nodes a subset of nodes with smallest stress ratios is selected and stresses re-evaluated at those locations only. The particular set of nodes is selected in a manner similar to that used for real time stress evaluation during power ascension

[24]. Specifically the list is comprised of the following nodes: " All nodes with an EPU alternating stress ratio, SR-a<2.0.

There are 12 such nodes on a weld.* The RPS sub-set of nodes with SR-a<3 at EPU. There are 30 such nodes on a weld and another 2 away from a weld." The RPS set of nodes on a weld with SR-P<2 at EPU. There are 16 such nodes.* The RPS set of nodes away from a weld with SR-P<3 at EPU. There are 2 such nodes.After eliminating redundant nodes (i.e., ones that appear in more than one of the above sets) the total number of nodes is 53. These are listed below in Table 10. Locations with stresses that are affected by the installation of the U-section stiffener are separated out in a separate group to facilitate subsequent presentation of results. In real time stress analysis, the stress intensities are calculated in the same manner as when analyzing the entire dryer and the limiting stresses with frequency shifting included are reported.

The resulting stress ratios are summarized in Table 17 together with the dominant frequencies.

65 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 10a. List of nodes used for real time evaluation.

Nodes 1-18.Index Description Node Coordinates

[in ] Reason for Other x y z SRF Selection (1)1 Inner Side Plate 37229 3.1 119.0 0.5 1 1 2 Support/Seismic Block 113286 10.2 123.8 -9.5 1 2 (2)3 Inner Side Plate 37592 14.4 -119.0 88.0 1 2 4 Side Plate Ext/Inner Base Plate 94143 16.3 119.0 0.0 1 3 5 Thin Vane Bank Plate/Hood Support/Inner Base Plate 85191 24.1 -59.5 0.0 1 3 6 Tie Bar 143795 -49.3 -108.1 88.0 0.77 3 7 Upper Support Ring/Seismic Block/Support 113508 -122.1 10.2 -9.5 1 3 8 Thin Vane Bank Plate/Hood Support/Middle Base Plate 98968 55.6 -54.3 0.0 1 3 9 Outer Cover Plate/Outer Hood 94498 102.8 -58.1 0.0 1 3 10 Hood Support/Middle Base Plate/Inner Backing Bar/Inner Hood 90468 -39.9 59.5 0.0 0.78 3 11 Side Plate/Top Plate 101600 -17.6 -119.0 88.0 0.66 3 12 Hood Support/Middle Base Plate/Inner Backing Bar/Inner Hood 88639 39.9 0.0 0.0 0.78 3 13 Thin Vane Bank Plate/Hood Support/Inner Base Plate 92995 24.1 0.0 0.0 1 3 14 Tie Bar 141237 25.0 108.1 88.0 1 3 15 Closure Plate/Middle Hood 89317 60.2 -85.2 87.0 1 4 (X)16 Side Plate/Closure Plate/Exit Top Perf/Exit Mid Top Perf 87784 -78.5 -85.2 74.5 1 4 (3)17 Outer End Plate/Outer Hood 94509 101.9 -63.3 24.6 1 4 18 Thin Vane Bank Plate/Inner Base Plate 99635 15.6 114.4 0.0 1 4 Notes: 1. Reasons for selection:

1 -small SR-P on no-weld node; 2 -small SR-a on non-weld node; 3 -SR-P<2; 4 -node on weld with SR-a<3 taken from RPS. 14 -remaining nodes on welds with SR-a<2.2. Stress corrected to account for 0.75 extension of support lug under USR (see Section 5.3.5).3. Compensates for poor mesh quality near top lifting rod brace (see Section 5.3.3).X. Extrapolation used to estimate stress at top of closure plate/curved hood weld (see Section 5.3.1).node on weld with 66 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 10b. List of nodes used for real time evaluation.

Nodes 19-36.Index Description Node Coordinates

[in ] Reason for Other x y z SRF Selection (1)19 Inner Side Plate/Inner Base Plate 99200 -2.3 -119.0 0.0 1 4 20 USR/Support/Seismic Block 113554 -6.9 -122.3 -9.5 1 4 (2)21 Side Plate/Top Plate 103080 49.6 -108.6 88.0 0.77 4 22 Seismic Block/Support 113400 -123.8 10.2 -9.5 1 4 23 Hood Support/Outer Cover Plate/Outer Hood 95267 -102.8 28.4 0.0 0.8 4 24 Thick Vane Bank Plate/Thin Vane Bank Plate/Side 90786 87.0 -85.2 11.6 1 4 Plate/Side Plate Ext/Outer End Plate 25 Tie Bar 137575 17.6 59.8 88.0 1 4 26 Closure Plate/Inner Hood 95172 28.8 -108.6 87.0 1 4 (X)27 Side Plate/Top Plate 91055 81.1 -85.2 88.0 0.77 4 28 Submerged Drain Channel/Submerged Skirt 93488 -91.0 -76.7 -100.0 1 4 29 Thin Vane Bank Plate/Hood Support/Outer Base Plate 98956 -87.0 28.4 0.0 0.58 4 30 Middle Base Plate/Inner Backing Bar Out/Inner 85631 39.9 108.6 0.0 1 4 Backing Bar/Inner Hood 31 Entry Bottom Perf/Side Plate/Outer End Plate 101818 -87.0 85.2 29.3 1 4 32 Thin Vane Bank Plate/Side Plate Ext/Outer Base Plate 98624 78.5 85.2 0.0 1 4 33 Outer End Plate/Outer Hood 94514 100.8 -64.9 36.8 1 4 34 Hood Support/Outer Base Plate/Middle Backing Bar 95428 -71.3 0.0 0.0 1 4 35 Outer Cover Plate/Outer End Plate/Outer Hood/Outer End Plate Ext 84090 102.8 -62.0 0.0 1 4 36 Thin Vane Bank Plate/Hood Support/Middle Base Plate 99451 55.6 54.3 0.0 1 4 Notes: 1. Reasons for selection:

1 -small SR-P on no-weld node; 2 -small SR-a on non-weld node; 3 SR-P<2; 4 -node on weld with SR-a<3 taken from RPS. 14 -remaining nodes on welds with SR-a<2.2. Stress corrected to account for 0.75 extension of support lug under USR (see Section 5.3.5).3. Compensates for poor mesh quality near top lifting rod brace (see Section 5.3.3).X. Extrapolation used to estimate stress at top of closure plate/curved hood weld (see Section 5.3.1).-node on weld with 67 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 10c. List of nodes used for real time evaluation.

Nodes 37-53.Index Description Node Coordinates

[in I Reason for Other x y z SRF Selection (1)37 Submerged Drain Channel/Skirt 93451 -11.5 -118.4 -101.5 0.56 4 38 Hood Support/Outer Base Plate/Middle Backing Bar 98172 -71.3 54.3 0.0 0.78 4 39 Submerged Drain Channel/Submerged Skirt 90924 91.0 76.7 -101.5 0.56 4 40 Side Plate/Exit Top Perf/Inner Side Plate 100989 15.6 -119.0 85.3 1 4 41 Submerged Drain Channel/Submerged Skirt 90926 11.5 118.4 -100.0 1 4 42 Skirt/Skirt USR overlap 99931 54.2 105.9 -9.5 1 4 43 Thick Vane Bank Plate/Thin Vane Bank Plate/Side 91091 24.1 119.0 11.6 1 4 Plate/Side Plate Ext/End Plate 44 Closure Plate/Middle Hood 88702 -60.2 85.2 87.0 1 14 (X)Nodes Addressed Using U-Section Stiffener 45 Top Thick Plate/Side Plate/Exit Top Perf/Inner Side Plate 101861 -15.6 -119.0 86.5 0.8 4 (M)46 Top Thick Plate/Side Plate/Exit Top Perf/Inner Side Plate 95197 15.6 -119.0 86.5 0.8 14 (M)47 Side Plate/Inner Side Plate/Top Plate 99407 -16.6 -119.0 88.0 0.66 14 (M)48 Top Thick Plate/Side Plate/Exit Top Perf/Inner Side Plate 98442 15.6 119.0 86.5 0.8 14 (M)49 Top Thick Plate/Side Plate/Inner Side Plate/Top Plate 98444 15.6 119.0 88.0 0.8 14 (M)50 Top Thick Plate/Side Plate/Exit Top Perf/Inner Side Plate 98451 -15.6 119.0 86.5 0.8 14 (M)51 Top Thick Plate/Side Plate/Inner Side Plate/Top Plate 98452 -15.6 119.0 88.0 0.8 14 (M)52 Top Thick Plate/Side Plate/Inner Side Plate/Top Plate 99408 -15.6 -119.0 88.0 0.8 14 (M)53 Top Thick Plate/Side Plate/Inner Side Plate/Top Plate 91240 15.6 -119.0 88.0 0.8 14 (M)Notes: 1. Reasons for selection:

1 -small SR-P on no-weld node; 2 -small SR-a on non-weld node; 3 -node on weld with SR-P<2; 4 -node on weld with SR-a<3 taken from RPS. 14 -remaining nodes on welds with SR-a<2.2. Stress corrected to account for 0.75 extension of support lug under USR (see Section 5.3.5).3. Compensates for poor mesh quality near top lifting rod brace (see Section 5.3.3).X. Extrapolation used to estimate stress at top of closure plate/curved hood weld (see Section 5.3.1).M. Nodes to be addressed in Section 5.8 using U-section stiffener.

68 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

5.3 Examination

of Low Stress Ratio Locations Section 5.1 identifies three distinct locations with alternating stress ratios below 2.0. These are addressed below.5.3.1 Analysis of Nodes Associated with Entry I in Table 9c The limiting alternating stress node corresponds to the top of the weld joining the closure plate to the curved surface of the middle hood. This location had previously been troublesome due to vibrations of the unmodified closure plate. Before modification it was noted that the origin of high stresses in the weld were caused by the presence of a 2-1 plate mode (2n-order mode in the vertical direction and 1 st order mode in the horizontal direction).

This led to high weld stresses both near the top of the weld and further down near where the second peak of the modal response occurs. Because the stress was high at interior weld nodes (i.e., not only at the end) it was deemed physically correct and was evaluated accordingly.

At first, an investigation was undertaken to increase weld size, but this was found insufficient for meeting the target stress ratios. Therefore, a more aggressive option was implemented of installing reinforcement ribs that completely suppress closure plate vibration at frequencies below 250 Hz.With the reinforced plate present, the top node on the closure plate/middle hood attachment weld has a computed alternating stress ratio of SR-a=l.68 which is below the target of 2.0. A second node, node 88702, at the 1800 rotated location has an alternating stress ratio of SR=1.98, which is also barely below the target level. In each case, unlike the weld stress distribution observed in the pre-modification dryer (i.e., before ribs were installed) the current stress peak is highly localized, confined to a single node and occurs at the end (rather than within the run of) a weld. The high stress occurs in the middle hood and is mainly due to bending of the shell element. Examination of the weld node immediately below the limiting node location reveals that its stress intensity is less than 20% of the limiting value -i.e., the stress grows by more than a factor of 5 over the 0.9" distance separating the limiting node and the one immediately below it. Such rapid rises in stress are characteristic of a structural discontinuity and are not realistically captured in finite element modeling.

Moreover, using the stresses at such discontinuities is not consistent with proper application of the ASME code, which relies on'nominal' stresses -i.e., the gradually varying stresses one would normally obtain in beam or plate analysis -and concentration factors to account for stress rise at structural discontinuities.

This complication is well known -see [39] for a layman's review -and is particularly problematic at the end of a weld since even the linearized stress can become singular at such locations (which also makes these locations more difficult to analyze using embedded modeling and explicit representation of the welds). Thus instead of utilizing a numerical code to estimate a stress that increases without bound as mesh resolution is reduced, it is preferred to refer to stresses that are converged and then account for the stress increase through the use of adjustment factors obtained through years of field and testing experience of welded structures.

Weld factors are a familiar example where rather than conducting a detailed geometry weld analysis, resulting in stresses that grow without bound as grid spacing is reduced, it is more accurate (in terms of agreement with actually observed weld failure) to multiply a converged nominal stress with a weld factor that accounts for stress concentration, weld quality and plastic deformations at toes and roots.69 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information It is sometimes the practice to simply dismiss such locations on the basis that they are clearly singularities and thus not compatible with ASME code stress definitions.

Instead, in the present case, a realistic estimate of the stress at the end of the weld is obtained by extrapolating the converged stresses from weld nodes located below the limiting point. In order to obtain a conservative estimate of the end-of-weld stress intensity suitable for ASME code evaluation, the following procedure is followed.

First the succession of nodes along the y-direction is identified.

The first adjacent node is denoted as node 'A', the second by 'B' and the third by 'C'. Next, various extrapolations of the stresses at these nodes to the limiting location are performed.

These include (for uniform spacing): Linear extrapolation through nodes A & B: GLEI =2*YA -GB Linear extrapolation through nodes A & C: GL2 = 1.5"cA -0.5yC Quadratic extrapolation through nodes A, B & C: crQ = 3*aA -3*aB + GC where 0 A, caB and crC are the stresses at nodes, A, B and C, respectively.

These formulae presume uniform spacing of the involved nodes along the edge. More general formulae are easily developed using Lagrangian interpolation functions so that: Caext = WA*GA + WB*aB + Wc*cyc where Lin. extrap. (A & B): wA_= d = d A dA dB dA dB dc -dA Lin. extrap. (A & C): wA dc , dw dA dC dA C d~dC dcdA dAdB Quadratic extrap.: w A = d , w B ==c ,WCd (dA dB)(dA dc) (dB dc)(dB dA)' WC (dc dA)(dc dB)and dA, dB and dC are the respective distances of nodes A, B and C from the singular location.The last step is to define the limiting stress as: ,7* = max{ UL1, YL2, aQ aA, A B, }which is simply the maximum stress obtained from any interpolation method or the three adjacent nodes along the respective edge.The results of this extrapolation are summarized in Table 11. In this case, the quadratic order fit through the alternating stress intensities of the three nodes immediately below the limiting node is found to be the most conservative of the six extrapolation options (1 quadratic, 2 linear, and 3 zero-th order fits). Even so, for the limiting node the stress at the weld end obtained using the quadratic interpolation is less than 25% of the value extracted from the finite element code, further confirming that this FEA stress estimate is fictitious and not representative of the stress actually existing in the steam dryer. If one uses the weld factor of 1.8 used at other fillet welds, the limiting alternating stress ratio obtained with quadratic extrapolation is SR-a=6.83 confirming that the target alternating stress ratio of 2.0 margin is amply met.In the subsequent analyses in Sections 5.4-5.8 this extrapolation is employed at the tops of the welds connecting the closure plates and curved hoods (inner and middle). This is a total of 8 welds each exhibiting the same kind of singular behavior near the top end of the weld.70 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 11. Alternating stress intensities distribution along the middle hood/closure plate weld near the limiting stress locations with FEA computed stress ratios, SR-a<2. Stress values correspond to limiting values with frequency shifting.

The top of the weld occurs at z=87 in.End of weld stress estimates obtained by quadratic extrapolation are also shown together with the quadratic fit parameters.

Node 89317 88702 (x,y) (in) (+60.2, -85.2) (-60.2,+85.2) z (in) Salt (psi)82.2258 (node C) 655.4 555.9 84.1692 (node B) 585.9 486.2 86.1098 (node A) 807.1 757.6 87 4183111 3459 Lin. extrap. (nodes A, B) 908.6 882.1 Lin. extrap. (nodes A, C) 841.9 803.9 Quad. extrap. (nodes A, B, C) 1006 996.2 Max {extrap, A, B, C 1 1006 996.2 Extrapolated end SR-a 6.83 6.89 Note: 1. The stress differs slightly from Table 9c since they correspond to results obtained with the lifting rod/upper brace junctions modified to reflect as-built constraints (i.e., no weld at these junctions) 71 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Stresses Near Node 89317! -I--,' .[..5000-1 J 4000 0.U)U)a)L.3000 2000 0 Nodal Stresses-Linear Extrap. (A,B).........

Linear Extrap. (A,C)-Quadratic Extrap.1000 0 0 0 1 2 3 4 5 6 7 s (= 87 -z) [in ]Stresses Near Node 88702 3500.-.I I--....-I o0 U)3000 2500 2000 1500 1000 500 0* Nodal Stresses-Linear Extrap. (A,B).........

Linear Extrap. (A,C)-Quadratic Extrap....................................

0 0 1 2 3 4 5 6 7 s (= 87 -z) [in ]Figure 15. Nodal FEA stresses and various extrapolations to the limiting stress location at nodes 89317 and 88702 located on the lower and upper edges respectively of the top thick plate. The variable s is measured from the limiting stress location (end of horizontal weld or edge).72 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

5.3.2 Analysis

of Nodes Associated with Entr, 2 in Table 9c This location involves a complex junction between multiple, orthogonally oriented elements and experiences high stresses that are readily identifiable with the modal response of the inner side plate spanning between the inner vane banks. Several options were considered for addressing the stress at this location including: (i) Developing an embedded model where the local welds are modeled explicitly and linearized stresses extracted from various paths. (ii)Utilizing the stress interpolation method of Section 5.3.1. (iii) Various concepts for suppressing vibration of the inner side plate, culminating in the U-section stiffener that has been installed during the 2014 outage.Development of the embedded model (option (i) above), which was carried out after the preliminary evaluation in Section 5.1, is summarized in Section 4.5. During the preliminary evaluation a stress reduction factor of SRF=0.77 was used which corresponds to the value calculated for a geometrically similar junction referenced as Embedded Model 3 in [12] and in Table 8. For the final evaluation carried out below an embedded model for the actual location was developed that more accurately represents the local joined components and characteristic local loading. Two weld-specific SRFs are calculated using the paths indicated in Figure 12 and comparing the linearized embedded model stresses against those in the baseline global model at the same locations over the 60-120 Hz frequency interval.

The first is associated with the weld connecting the top thick plate to the inner side plate for which the SRF=0.80, which is slightly higher than the 0.77 value previously imputed to this location in the preliminary screening.

The second stress reduction factor pertains to the weld connecting the tie bar to the top plate and is calculated to be SRF=0.66.

The revised stress reduction factors for these welds are used in the ensuing results below. By themselves, these values are still insufficient to qualify Entry 2 in Table 9c.Therefore the option (ii) of utilizing the stress interpolation method in Section 5.3.1 was considered for the present location also. It was determined that if stresses along the horizontal welds connecting the top thick plate to: (a) the top plate of the vane bank and (b) the perforated plate, were extrapolated to the high stress locations and the revised SRF applied, then the recommended margin could be achieved at all load conditions.

However, because the present location involves multiple parts and weld lines that connect in a complex manner, there was some uncertainty as to whether the stress extrapolation method could be applied. Also, for one of the load conditions involving closure of the RCIC valve the limiting alternating stress ratio at this location was slightly below the recommended value thus necessitating cycle counting to establish adequate fatigue margin for this occasional operating condition.

For these reasons it was decided that physical modification of the dryer at this location should be considered as an option to meet the recommended alternating stress ratio of SR-a>2.The dominant stress contribution at Entry 2 is attributed to vibration of the side plate so that suppressing this vibration reduces the stress. This can be accomplished by shifting the response frequency away from the dominant signal peaks, and/or strengthening the structure so that the response amplitude is reduced. The reinforcement strip accomplishes both goals. The added stiffness and mass are such that the fundamental frequency is shifted upward. This is beneficial in the present case since the limiting stresses are obtained at positive frequency shifts of +7.5 to 73 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

+10%. However, even if the modal frequency were unchanged, the reinforcement strip would reduce the response amplitude leading to a lower stress.As summarized in Appendix B the final design consists of a U-section stiffener bolted to the side plate. Bolts are preferred over welding to limit diver exposure to radiation and also to avoid heat-induced stresses in the materials.

The final design bolts the U-cross-section stiffening beam on the interior side (exposed to the steam path) of the side plate using 6 bolts. Further details of the final design and its FEA implementation are given in Appendix B. Unit solutions with the bolted stiffening channel attached to the dryer were generated over 60-130 Hz. Since the mesh for this dryer model differs from that used in previous global dryer evaluations it is not possible to simply 'mix-and-match' the associated unit solutions.

Instead, the dryer stresses with the U-section stiffener attached are only evaluated at the current Entry 2, at the ends of the stiffening channel (to ensure there are no local stress increases that may exceed margin) and at the middle of the channel for the purpose of bolt design [25]. Section 5.8 shows that with the stiffening channel in place all nodes at Entry 2 meet the recommended SR-a>2 margin for all load and total uncertainty combinations.

These results in Section 5.8 with the channel in place still make use the SRFs (0.80 and 0.66 as calculated under option (i) above). However, it has also been shown that even if these SRFs are not invoked the alternating stress ratios at these locations remain above 2.0 except for the occasional loads associated with RCIC valve closure. For the RCIC line closed cases, it is appropriate to use cycle counting and evaluate the cumulative usage factors (CUFs) to assess fatigue. Based on the CUFs calculated for other nodes, it is expected that the resulting CUFs at these locations for these load cases would be well below 1.5.3.3 Analysis of Nodes Associated with Entry 3 in Table 9c This location lies on the perimeter of the thickened plate used to reinforce the upper lifting rod bracket. When considering all other nodes in the neighborhood of the limiting node (87784)and its reflected images, it is found (see Table 9c and Figure 14c) that this location only involves a single node with an alternating stress ratio, SR-a=I.80.

It occurs at the common element junction of the outer vane bank side plate, closure plate and the perforated plates. The side plate experiences the highest stress even though it is also the thickest of the connected members. This is because the forces induced by rod motion are equilibrated by the support structure

-in this case, primarily the bracket itself and the vane bank side plate. The modifications prior to the EPU operation reinforced this location using reinforcement plates. However, the stress becomes high at the perimeter of the reinforcement where the effective thickness transitions back to the original 3/8" value.Examination of the stress ratios at the corresponding locations at the three other corresponding upper lifting rod braces reveals that these ratios are considerably higher than 2.0 (2.45 at node 87592, 2.18 at node 89598 and 2.79 at node 101870). Also, the mesh at the limiting location is highly skewed whereas at the other three locations, it is more isotropic.

The high stress at the limiting location is therefore attributed to numerical overestimation associated with low mesh quality. To address this location the unit solutions were recomputed over the 13-26 Hz frequency range with a locally improved and refined (using a 1" spacing rather than the 2" used before) mesh that also implements the revised upper brace/lifting rod constraint (see next).74 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information The respective grids are compared in Figure 16. The specified frequency range covers one decade (as was done in embedded models) and completely encompasses the dominant response peak. The unit solution stresses near the response peak frequency 19.3 Hz on the old and new grids are compared in Figure 17.A structural evaluation, with frequency shifting included, is then carried out over this limited 13-26 Hz using the original and updated unit solutions using the normal operation EPU entrance pressures (appropriately zeroed outside the 13-26 Hz). The limiting alternating stress ratios obtained in this manner on the original and revised grids are SR-a=2.16 and 3.53 respectively so that over this frequency range only one can reduce the computed stress by a factor of SRF=2.16/3.53=0.61.

A conservative estimate of the limiting stress can then be obtained from: a = a(0,13) + SRF*a(13,26)

+ a(26,250)where a(f 1 ,f 2) is the stress contribution over the frequency interval [f ,f2] on the original mesh.Using this formula, the limiting stress ratio is SR-a=2.51 for' the baseline EPU load.A summary of the stress ratios at this location for the other loads is given below.Load SR-a outN55 (base EPU operation):

2.510 outN59 (drain trap out of service -92.5Hz signal): 2.636 outN58 (RCIC line close -89.3 Hz signal): 2.090 outN61 (base EPU load with conservative bias+unc.):

2.482 outN62 (same as outN59, but with conservative bias+unc.):

2.599 outN63 (same as outN58, but with conservative bias+unc.):

2.072 75 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information~Node 87784, over stressed New nodes to analyze Figure 16. Original (top) and refined (bottom) meshes and the nodes accessed for stress comparison.

Solid elements representing the lifting rod are omitted here for better viewing of the shell element mesh.76 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information ANSYS M0G MA JAN 31 2014 09 :51:45 NODAL SOLUTION STEP-I SUB -1 FREQ=19. 334 REAL ONLY SINT (AVo)PoverOraphics EFACET-1 AVRES-Xat DUX -.258722 SUN -.627013 SUx -61033 0 3333 6667 10000 13333 16667 20000 23333 26667 30000 ANSYS 10. OA1 JAN 31 2014 09:52:41 NODAL SOLUTION STEP-13 SUB =1334 REAL ONLY SINT (AVG)PoverOraphics EFACET=1 AVRES-Mat DUX -.28S587 SUN -.738192 SUX -56831 0 3333 6667 10000 13333 16667 20000 23333 26667 30000 Figure 17. Nodal stresses obtained on the original (top) and refined (bottom) grids for the unit solution at 19.3 Hz. The stress at the limiting node is 17597 psi, whereas on the new mesh the limiting stress in the same neighborhood (see Figure 16) is 13818 psi, which corresponds to a 22% reduction in stress. The solid elements associated with the lifting rod are omitted here for better viewing of the high stress region.77 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

5.3.4 Evaluation

of Nodes on Upper Brace/Lifting Rod Junctions In previous evaluations, the upper lifting rod brace and the lifting rod were assumed connected by a weld. This reflected the best interpretation of the combined understanding of the drawings (particularly steam dryer modification drawing number 0016010001504) and as-built configuration, which are known to differ (for example, the drawings originally indicated the installation of four braces and the addition of straps to the upper braces). Several nodes located on the upper lifting rod brace were observed to have stress ratios below 2.0 under the assumption that the upper brace and lifting rod are welded together, thus requiring imposition of a weld factor of 1.8 on the alternating stress intensity.

However, careful review and photographic evidence have since confirmed that the upper most brace/lifting rod junctions are not welded (the lower two braces however, are). Application of the weld factor at the upper brace/lifting rod junction is therefore no longer warranted.

Also, lack of a welded connection between the upper brace and lifting rod means that the lifting rod only imparts lateral forces to the top brace; no transmission of vertical forces or any moments occurs between the rod and remaining steam dryer structure.

Therefore it is appropriate to model these connections as displacement constraints in the horizontal directions only (i.e., require that at the lines of connection the bracket and lifting rod experience identical deflections in the horizontal directions);

no constraints are imposed on displacements in the vertical direction or on any rotational degrees of freedom. Modifying the constraints in this manner does not alter the mesh structure or node indexing so that one retains the option of generating new unit solutions over the frequency ranges of concern and then overwriting existing unit solutions over these ranges. Despite this change in constraints conditions, one finds that the modal frequencies associated with the lifting rod do not change significantly.

This is because the structural reaction moments imposed by the support bracket upon the significantly stiffer and more massive lifting rod are small compared to the deformation stresses and inertial forces of the lifting rod. The main change is that the support structure no longer experiences moments and vertical forces imparted from the lifting rod leading to an overall reduction in local stresses.

The natural frequency of the fundamental mode changes from 19.62 Hz (fully welded) to 19.28 Hz (laterally constrained), which is amply covered by the +/-10% frequency shifting required in the stress evaluation.

In the real time stress evaluations performed below, the unit solutions are regenerated over the 17-22 Hz frequency interval with the boundary conditions at the lifting rod/upper brace connection modified as described above. All subsequent post-processing in Sections 5.4 to 5.8 is performed with these connections treated as non-welded.

Specifically, no weld factor is applied.This results in a limiting alternating stress ratio of SR-a=2.09 occurring on the contact edge between the upper brace and lifting rod. Several high stress locations also emerge on the brace that are on the junction with the vane side plate. These are located approximately 0.35" to 0.6" from the vertical face representing the outer vane bank wall. Since the reinforcement extends out to 1.63" from the outer vane bank wall (drawing 10082C94 in [6]), these nodes are actually within the reinforcement structure.

Stresses at these nodes are thus substantially reduced and the associated stress ratios well above 2.0 (SR-a=6.44 or higher).78 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

5.3.5 Correction

of Upper Support Ring/Support Interface Geometry The limiting peak stress occurs on the weld connecting the upper support ring to the earthquake block. This high stress was discovered to be due to an approximate and overly conservative representation of the USR/support interface contact area. Specifically in previous models the support lugs did not contact the USR so that all of the dryer weight is transferred through the weld attaching the earthquake block to the USR. However, based on drawing 197R624, I of 3 (transmitted by E-mail from Exelon on 03-03-2014), the support lugs actually extend under the support ring by 0.75". This means that a considerable portion of the steam dryer weight is now transmitted though this contact area, thus relieving the stresses in the earthquake block/USR weld. This relief was quantified by loading the USR with the deadweight of the steam dryer and comparing the limiting static stresses in the USR/earthquake block welds obtained with and without the 0.75" support lug projection underneath the USR. The two models are depicted in Figure 18 and are both meshed with the same settings, namely, 2" spacing everywhere, with 0.75" spacing near support lugs. The results from the two cases are compared in Table 12 and show that a conservative correction factor of 0.64 can be used to account for the actual 0.75" contact surface between the USR and support lug, and thus more accurately estimate the stresses in the USR/earthquake attachment weld.Table 12. Stresses in the USR/earthquake attachment welds at each of 4 support locations due to dryer deadweight load. Variations in stress intensities are due to mesh differences and dryer asymmetry.

Support Location Weld stress [ ksi] Weld stress [ ksi] Ratio Original support lug Extended (0.75") support lug 1 19.9 12.3 0.62 2 21.5 13.7 0.64 3 21.8 13.0 0.60 4 21.6 12.7 0.59 79 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information ANSYS'(311.

zýAWSJSQ LO WW ANY~ "~z)I-r X Figure 18. Geometry of the USR/support lug location involving the earthquake block (top, with 0.75" overlap included) and meshes with (left) and without (right) the 0.75" overlap.80 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

5.4 General

Stress Distribution and High Stress Locations The stress evaluation performed in Section 5.1 using the normal operating load is repeated here in greater detail and using the following analysis modifications listed in Section 5.3: (i) The extrapolation procedures described in Section 5.3.1 are applied to the tops of the welds connecting the closure plates to the curved hoods. (ii) The stress reduction factors developed in Section 4.5 for embedded model 6 and further summarized in Section 5.3.2 are appropriately applied instead of the 0.77 value used in the preliminary evaluation of Section 5.1. (iii) Real time results are used to manually adjust the stress ratio near the lifting rod brace as indicated in Section 5.3.3 to compensate for poor mesh quality. (iv) No weld factor is applied at the junction between the lifting rod and upper restraint brace as described in Section 5.3.4. (v) Nodes on the side plate found to under the brace reinforcement plate (see Section 5.3.4) are adjusted in the same manner as other nodes on the brace/side plate junction. (vi) The stresses in the USR/earthquake block attachment welds are adjusted as described in Section 5.3.5. As before the base EPU load with ACM 4. 1R bias and uncertainty values, is applied. The full steam dryer evaluations in Sections 5.4 to 5.6 do not account for the U-section stiffener and retain a full connection between the lifting rod/upper brace junction (i.e., all degrees of freedom are coupled).The maximum stress intensities obtained by post-processing the ANSYS stress histories for EPU at nominal frequency and with frequency shift operating conditions are listed in Table 13.Contour plots of the maximum stress intensities with all frequency shifts included are shown in Figure 19. The figures are oriented to emphasize the high stress regions. Note that these stress intensities do not account for weld factors but do include end-to-end bias and uncertainty.

Further, it should be noted that since the allowable stresses vary with location, stress intensities do not necessarily correspond to regions of primary structural concern. Instead, structural evaluation is more accurately made in terms of the stress ratios which compare the computed stresses to allowable levels with due account made for stress type and weld. Comparisons on the basis of stress ratios are made in Section 5.5 as well as the real time analysis Sections 5.7 -5.8.From Figure 19a and Table 13 the maximum stress intensities in most areas are low (less than 1000 psi). For the membrane stresses (Pm) the high stress regions tend to occur: (i) on the welds joining the seismic blocks and upper support ring (USR); (ii) the portion of the inner hood located outboard of the closure plate connecting the inner and middle vane banks; (iii) the bottom of the central vertical side plate that joins the innermost vane banks (stress concentrations occur where this plate is welded to the inner base plates resting on the upper support ring); (iv)the welds joining the tie bars to the top cover plates on the vane banks; and (v) the bottoms of the inner vane bank side plates where they connect to the USR.The membrane + bending stress (Pm+Pb) distributions evidence a more pronounced modal response especially on the inner and middle hood structures, and on the inner closure plates.High stress concentrations are recorded on the bottom edge of the inner hood outboard of the closure plate where it joins to the base plate and also near the dryer support locations.

Other areas with high Pm+Pb stress concentrations include: (i) the welded junctions between the tie bars and the top plates of the vane banks; (ii) tops of the closure plates where they are welded to a hood or vane bank end plates; (iii) the skirt/drain channel welds; and (iv) the outer hood side plates and their welded connection to each outer hood (see Figure 19b-c).81 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information The alternating stress, Salt, distributions are most pronounced on the hoods and their connected side plates. Though not exposed directly to the MSL acoustic sources, the inner and middle hoods are thinner than the outer ones and their responses are driven mainly by structural coupling rather than direct forcing. The highest stress intensity at any frequency shift occurs at the bottom of the inner hood where it meets the middle base plate. Significant response is also observed on: (i) the welds connecting the tie bars to the vane bank top plates; (ii) parts involving the inner side plate; (iii) the bottoms of drain channels and the junctions between the hoods, hood supports and base plates; (iv) the welds joining the closure plates to the hoods and vane banks;and (v) parts connecting to the lifting rods. These locations are characterized by localized stress concentrations as indicated in Figure 19e and have emerged as high stress locations in other steam-dryers also.82 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 13a. Locations with highest predicted stress intensities for EPU conditions with no frequency shift.Stress Location Weld SRF(a) Location (in) node Stress Intensities (psi) Dom.Category x y z Pm Pm+Pb Salt Freq. (Hz)Pm Inner Side Plate No 3.1 119 0.5 37229 6648 8168 1761 83.3 Side Plate Ext/Inner Base Plate Yes 16.3 119 0 94143 6208 9017 1275 19.8" USR/Support/Seismic Block(c) Yes -6.9 -122.3 -9.5 113554 5060 5060 1852 15.7" Thin Vane Bank Plate/Hood Support/Inner Base Plate Yes -24.1 59.5 0 99487 4882 4895 1677 15.7" Hood Support/Outer Cover Plate/Outer Hood(4) Yes 0.8 -102.8 28.4 0 95267 4774 4820 2602 15.7 Pm+Pb Middle Base Plate/Inner Backing Bar Out/Inner Backing Bar/Inner Hood Yes-39.9-108.6 0 84197 1293 9339 1799 15.7 Side Plate Ext/Inner Base Plate Yes 16.3 119 0 94143 6208 9017 1275 19.8 Inner Side Plate No 3.1 119 0.5 37229 6648 8168 1761 83.3 Side Plate/Top Plate(2) Yes 0.77 49.6 108.6 88 93256 1995 6963 1465 19.7 Outer Cover Plate/Outer Hood Yes 102.8 -58.1 0 94498 1061 6065 1014 15.7 Collar/Collar Contact No -79.2 -87.5 75.8 91651 929 5394 5192 19.8 Brace No -79.6 -85.5 53.5 37693 4069 4312 3403 19.8 Inner Side Plate No 14.4 -119 88 37592 657 3831 3126 82.6 Top Thick Plate/Side Plate/Exit Top Perf/Inner Yes 0.8 -15.6 -119 86.5 101861 535 3935 3107 82.6 Side Plate(6)Side Plate/Closure Plate/Exit Top Perf/Exit Yes -78.5 -85.2 74.5 87784 1171 2627 2624 19.8 Mid Top Perf(d) I I I Notes: (a) [I ((b) Full penetration weld so that weld factor, WF=1.4.(c) Corrected for support lug contact area per Section 5.3.5.(d) Compensated for mesh quality per Section 5.3.3.(1-6) Number referring to the 3) Entry is empty if no SRF is applied.(3)j]83 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 13b. Locations with highest predicted stress intensities taken over all frequency shifts at EPU conditions.

Stress Location Weld SRF(a) Location (in) node Stress Intensities (psi) % Freq. Dom.Category x y z Pm Pm+Pb Salt Shift Freq. (Hz)Pm Inner Side Plate No 3.1 119 0.5 37229 6685 8548 2281 5 123.5" Side Plate Ext/Inner Base Plate Yes 16.3 119 0 94143 6396 9246 1525 10 39.6 Tie Bar(2) Yes 0.77 -49.3 -108.1 88 143795 5644 5644 1773 10 82.6 Thin Vane Bank Plate/Hood Support/Inner Yes 24.1 -59.5 0 85191 5245 5330 2030 10 13.3 Base Plate USR/Support/Seismic Block(c) Yes -6.9 -122.3 -9.5 113554 5205 5205 1951 5 13.7 Pm+Pb Middle Base Plate/Inner Backing Bar Yes 39.9 108.6 0 85631 1478 10031 2750 10 82.6 Out/Inner Backing Bar/Inner Hood III" Side Plate Ext/Inner Base Plate Yes 16.3 119 0 94143 6396 9246 1525 10 39.6" Inner Side Plate/Inner Base Plate Yes -2.3 -119 0 99200 4026 8778 3167 7.5 123.5 Side Plate/Top Plate(2) Yes 0.77 49.6 108.6 88 93256 2154 7999 2563 10 82.6" Outer Cover Plate/Outer Hood Yes 102.8 -58.1 0 94498 1252 6440 1422 7.5 82.0 Salt Collar/Collar Contact No -79.2 -87.5 75.8 91651 968 6176 5922 5 18.8 Inner Side Plate No 14.4 -119 88 37592 764 5055 4235 7.5 78.7 Top Thick Plate/Side Plate/Exit Yes 0.8 -15.6 -119 86.5 101861 571 4859 4084 7.5 78.7 Top Perf/Inner Side Plate(6)Brace No -79.6 -85.5 53.5 37693 5100 5328 3913 7.5 18.7 Outer End Plate/Outer Hood Yes 101.9 -63.3 24.6 94509 795 4368 3233 10 82.0 Notes: (a) (3) Entry is empty if no SRF is applied.(b) Full penetration weld so that weld factor, WF= 1.4.(c) Corrected for support lug contact area per Section 5.3.5.(d) Compensated for mesh quality per Section 5.3.3.(1-6) Number referring to the (3)84 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Y Pm (psi]7000 6500 6000 5500 5000 4500 4000 3500 3000 2500 2000 1500 1000 500 0 Figure 19a. Contour plot of maximum membrane stress intensity, Pm, for EPU operation with frequency shifts. The recorded stress at a node is the maximum value taken over all frequency shifts. The maximum stress intensity is 6685 psi.85 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Z Y Pm+Pb [psi]10000 9000 8000 7000 6000 5000 4000 3000 2000 1000 0 Figure 19b. Contour plot of maximum membrane+bending stress intensity, Pm+Pb, for EPU operation with frequency shifts. The recorded stress at a node is the maximum value taken over all frequency shifts. The maximum stress intensity is 10031 psi.First view.86 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 19c. Contour plot of maximum membrane+bending stress intensity, Pm+Pb, for EPU operation with frequency shifts. Second view from beneath.87 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Z Sal t[psi)6000 5500 5000 4500 4000 3500 3000 2500 2000 1500 1000 500 0 Figure 19d. Contour plot of alternating stress intensity, Salt, for EPU operation with frequency shifts. The recorded stress at a node is the maximum value taken over all frequency shifts. The maximum alternating stress intensity is 5922 psi. First view.88 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

,, Y Salt [psi]6000 5500 5000 4500 4000 3500 3000 2500 2000 1500 1000 500 0 Figure 19e. Contour plot of alternating stress intensity, Salt, for EPU operation with frequency shifts. The recorded stress at a node is the maximum value taken over all frequency shifts. Second view from below.89 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 5.5 Load Combinations and Allowable Stress Intensities The stress ratios computed for EPU at nominal frequency and with frequency shifting are listed in Table 14 (without frequency shifting) and Table 15 (with frequency shifting).

The stress ratios are grouped according to type (SR-P for maximum membrane and membrane+bending stress, SR-a for alternating stress) and location (away from welds or on a weld). The tabulated nodes with frequency shifting in Table 15 are also depicted in Figure 20. The plots corresponding to maximum stress intensities depict all nodes with stress ratios less than 5 or less than 4 as indicated.

For EPU operation at nominal frequency (no frequency shift) the minimum stress ratio is identified as a maximum stress, SR-P=l.49, and is recorded at the bottom of the inner hood where it meets the middle base plate. In previous evaluations

[5] the weld joining the upper support ring to the earthquake block had the lowest stress ratio, but this location is no longer limiting following the correction for support lug extension under the USR as described in Section 5.3.5. The minimum alternating stress ratio at zero frequency shift, SR-a=2.21, occurs at the top of the inner vane bank where it meets the inner side plate and vane bank end plate (location 1 in Figure 20h).The effects of frequency shifts are conservatively accounted for by identifying the minimum stress ratio at every node, where the minimum is taken over all the frequency shifts considered (including the nominal or 0% shift case). The results are summarized in Table 15 and show that the lowest stress ratio, SR-P=l.39, occurs at the same location as in the nominal case (rotated by 1800) and is only slightly lower. The next three lowest SR-P locations in Table 15b are the same as in Table 14b or at a reflected image location (for the second entry). With frequency shifting the lowest alternating stress ratio also occurs at the same location (see location 1 in Figure 20h)and assumes a value of SR-a=l.68.

Based on the largest Fourier coefficient, the dominant frequency in the signal contributing to this stress is 78.7 Hz; with the +7.5% shift accounted for it induces a structural response at 84.6 Hz. More details of the corresponding stress response spectra are provided in the following section. Other than this one location, all other entries in Table 15c have alternating stress ratios above 2.0. The alternating stress ratio below 2.0 is addressed in Section 5.8 by adding a U-section stiffener.

When this channel is added to the side plate, all stress ratios about this location achieve stress ratios that are well above 2.0 for all of the load and total uncertainty combinations.

The next lowest alternating stress ratio is in Table 15a and occurs where the upper brace contacts the lifting rod with SR-a=2.09.

In the CLTP-based stress evaluation

[11] the limiting node location was node 95267, which appears here as the 6 th entry in Table 15c. Virtually all of the limiting stress ratios occur at either the +7.5% or +10%shifts.90 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 14a. Limiting non-weld locations with at EPU conditions with no frequency shift. Stress ratios are grouped according to stress type (maximum -SR-P; or alternating

-SR-a).Stress Location Location (in.) node Stress Intensity (psi) Stress Ratio Dom.Ratio x y z Pm Pm+Pb Salt SR-P SR-a Freq. (Hz)SR-P 1. Inner Side Plate 3.1 119 0.5 37229 6648 8168 1761 2.54 7.02 83.3 2. Hood Support 89 -28.4 0 14474 4491 4545 2562 3.76 4.83 15.7" 3. Thin Vane Bank Plate -15.6 -118.4 0.6 2558 4204 4617 565 4.02 21.89 19.7 SR-a 1. Collar/Collar Contact -79.2 -87.5 75.8 91651 929 5394 5192 4.70 2.38 19.8" 2. Brace -79.6 -85.5 53.5 37693 4069 4312 3403 4.15 3.63 19.8" 3. Inner Side Plate 14.4 -119 88 37592 657 3831 3126 6.62 3.96 82.6 91 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 14b. Limiting peak stress ratios, SR-P, on welds at EPU conditions with no frequency shift. Bold text indicates minimum stress ratio on the structure.

Location SRF(a) Location (in.) node Stress Intensity (psi) Stress Ratio Dom.X y z Pm Pm+Pb Salt SR-P SR-a Freq. (Hz)

Plate/Inner Backing Bar Out/Inner,--...

39.9 -108.6 0 84197 1293 9339-. 1799 .1.49 3,82 .1517..... ... ... ......

• H... =. .. .. .-. ...... .....*. ....... ......_ _ :. .._. ...... : ... ._ : /7 ..: o ..2. Side Plate Ext/Inner Base Plate 16.3 119 0 94143 6208 9017 1275 1.50 5.39 19.8 3. Hood Support/Outer Base Plate/Middle Backing Bar -71.3 0 0 95428 5691 5803 2216 1.63 3.10 15.7 4. Inner Side Plate/Inner Base Plate -2.3 -119 0 99200 3972 7741 2363 1.80 2.91 82.6 5. Upper Support Ring/Support/Seismic Block(c) -6.9 -122.3 -9.5 113554 5060 5060 1852 1.84 3.71 15.7 6. Hood Support/Middle Base Plate/Inner Backing -39.9 0 0 85723 5008 5242 1500 1.86 4.58 73.0 Bar/Inner Hood(b)7. Tie Bar(2) 0.77 49.3 108.1 88 141275 4911 4911 1042 1.89 6.59 19.7 8. Thin Vane Bank Plate/Hood Support/Inner Base Plate -24.1 59.5 0 99487 4882 4895 1677 1.90 4.10 15.7 9. Hood Support/Outer Cover Plate/Outer Hood(4) 0.8 -102.8 28.4 0 95267 4774 4820 2602 1.95 2.64 15.7 10. Thin Vane Bank Plate/Hood Support/Middle Base Plate 55.6 -54.3 0 98968 4542 4625 1716 2.05 4.00 60.4 11. Hood Support/Middle Base Plate/Inner Backing 39.9 -59.5 0 101435 4525 4918 1621 2.05 4.24 60.4 Bar/Inner Hood(b)12. Outer Cover Plate/Outer Hood 102.8 -58.1 0 94498 1061 6065 1014 2.30 6.77 15.7 Notes: (a) [[(b) Full penetration weld so that weld factor, WF=1.4.(c) Corrected for support lug contact area per Section 5.(d) Compensated for mesh quality per Section 5.3.3.(1-6) Number referring to the Continuum Dynamics, Inc. Proprietary Information Table 14c. Limiting alternating stress ratios, SR-a, on welds at EPU conditions with no frequency shift.Location SRF(a) Location (in.) node Stress Intensity (psi) Stress Ratio Dom.x y z Pm Pm+Pb Salt SR-P SR-a Freq. (Hz)1. Top Thick Plate/Side Plate/Exit Top Perf/Inner 0.8 -15.6 -119 86.5 101861 535 3935 3107 3.54 2.21 82.6 Side Plate(6)2. Side Plate/Closure Plate/Exit Top Perf/Exit Mid -78.5 -85.2 74.5 87784 1171 2627 2624 5.31 2.62 19.8 Top Perf(d)3. Hood Support/Outer Cover Plate/Outer Hood(4) 0.8 -102.8 28.4 0 95267 4774 4820 2602 1.95 2.64 15.7 4. Inner Base Plate 23.1 113.2 0 66988 862 6254 2473 2.23 2.78 19.7 5. Inner Side Plate/Inner Base Plate -2.3 -119 0 99200 3972 7741 2363 1.80 2.91 82.6 6. Side Plate/Brace(d) 79.7 -85.2 31.2 87633 3034 3175 2319 3.06 2.96 19.7 Notes: (a) [I (b) Full penetration weld so that weld factor, WF=1.4.(c) Corrected for support lug contact area per Section 5.(d) Compensated for mesh quality per Section 5.3.3.(1-6) Number referring to the Continuum Dynamics, Inc. Proprietary Information Table 15a. Limiting non-weld locations with at EPU conditions with frequency shifts. Stress ratios are grouped according to stress type (maximum -SR-P; or alternating

-SR-a). Locations are depicted in Figure 20.Stress Location Location (in.) node Stress Intensity (psi) Stress Ratio % Freq. Dom.Ratio x y z Pm Pm+Pb Salt SR-P SR-a Shift Freq. (Hz)SR-P 1. Inner Side Plate 3.1 119 0.5 37229 6685 8548 2281 2.53 5.42 5 123.5" 2. Hood Support 89 -28.4 0 14474 5050 5210 3068 3.35 4.03 10 82.0" 3. Inner Hood 26.8 108.2 88 72608 1780 6838 1743 3.71 7.09 10 82.6 4. Thin Vane Bank Plate -15.6 -118.4 0.6 2558 4311 4727 833 3.92 14.84 10 16.0 SR-a 1. Collar/Collar Contact -79.2 -87.5 75.8 91651 968 6176 5922 4.10 2.09 5 18.8 2. Inner Side Plate 14.4 -119 88 37592 764 5055 4235 5.02 2.92 7.5 78.7" 3. Brace -79.6 -85.5 53.5 37693 5100 5328 3913 3.31 3.16 7.5 18.7" 4. Side Plate/Brace(a)

-79.7 -85.2 75.8 90307 2363 3327 3232 7.15 3.83 2.5 19.5 Note: (a) Adjusted according to Table 11 of [26]94 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 15b. Limiting peak stress ratios, SR-P, on welds at EPU conditions with frequency shifts. Bold text indicates minimum stress ratio on the structure.

Locations are depicted in Figure 20.Location SRF(a) Location (in.) node Stress Intensity (psi) Stress Ratio % Freq. Dom.x y z Pm Pm+Pb Salt SR-P SR-a Shift Freq. (Hz)1. Middle Base Plate/Inner Backing Bar 39.9 108.6 0 85631 1478 10031 2750 1.39 2.50 10 82.6 Out/Inner Backing Bar/Inner Hood 2. Side Plate Ext/Inner Base Plate 16.3 119 0 94143 6396 9246 1525 1.45 4.51 10 39.6 3. Hood Support/Outer Base Plate/Middle Backing -71.3 0 0 95428 6159 6227 2634 1.51 2.61 7.5 14.4 Bar 4. Inner Side Plate/Inner Base Plate -2.3 -119 0 99200 4026 8778 3167 1.59 2.17 7.5 123.5 5. Tie Bar(2) 0.77 -49.3 -108.1 88 143795 5644 5644 1773 1.65 3.87 10 82.6 6. Thin Vane Bank Plate/Hood Support/Inner Base 24.1 -59.5 0 85191 5245 5330 2030 1.77 3.38 10 13.3 Plate 7. Thin Vane Bank Plate/Hood Support/Middle 55.6 -54.3 0 98968 5241 5318 2232 1.77 3.08 10 82.0 Base Plate 8. USR/Support/Seismic Block(c) -6.9 -122.3 -9.5 113554 5205 5205 1951 1.79 3.52 5 13.7 9. Hood Support/Middle Base Plate/Inner Backing 39.9 0 0 88639 5159 5558 1926 1.80 3.57 10 69.5 Bar/Inner Hood(b)10. Hood Support/Outer Cover Plate/Outer Hood(4) 0.8 -102.8 28.4 0 95267 5054 5081 2996 1.84 2.29 10 13.7 11. Hood Support/Middle Base Plate/Inner Backing -39.9 59.5 0 90468 5044 5180 1849 1.84 3.72 10 82.6 Bar/Inner Hood(b)Notes: (a) (3) Entry is empty if no SRF is applied.(b) Full penetration weld so that weld factor, WF=1.4.(c) Corrected for support lug contact area per Section 5.3.5.(d) Compensated for mesh quality per Section 5.3.3.(1-6) Number referring to the (3)95 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 15c. Limiting alternating stress ratios, SR-a, on welds at EPU conditions with frequency shifts. Locations are depicted in Figure 20.Location SRF(a) Location (in.) node Stress Intensity (psi) Stress Ratio % Freq. Dom.x y z Pm Pm+Pb Salt SR-P SR-a Shift Freq. (Hz)1. Top Thick Plate/Side Plate/Exit Top Perf/Inner 0.8 -15.6 -119 86.5 101861 571 4859 4084 2.87 1.68 7.5 78.7 Side Plate(6)2. Outer End Plate/Outer Hood 101.9 -63.3 24.6 94509 795 4368 3233 3.19 2.12 10 82.0 3. Thin Vane Bank Plate/Inner Base Plate 15.6 114.4 0 99635 3531 5532 3194 2.52 2.15 7.5 82.0 4. Inner Side Plate/Inner Base Plate -2.3 -119 0 99200 4026 8778 3167 1.59 2.17 7.5 123.5 5. Side Plate/Top Plate(2) 0.77 49.6 -108.6 88 103080 1406 5009 3046 2.78 2.25 10 82.6 6. Hood Support/Outer Cover 0.8 -102.8 28.4 0 95267 5054 5081 2996 1.84 2.29 10 13.7 Plate/Outer Hood(6)7. Side Plate/Brace(4) 0.64 79.7 -85.2 31.2 87633 3526 3734 2976 2.64 2.31 10 96.3 8. Thick Vane Bank Plate/Thin Vane Bank 87 -85.2 11.6 90786 907 10028 2941 1.39 2.34 10 82.0 Plate/Side Plate/Side Plate Ext/Outer End Plate 9. Tie Bar 17.6 59.8 88 137575 4549 4549 2906 2.04 2.36 10 82.6 10. Side Plate/Top Plate(2) 0.77 81.1 -85.2 88 91055 858 3611 2811 3.86 2.44 10 86.8 11. Submerged Drain Channel/Submerged Skirt 76.7 -100 93488 498 4963 2811 2.81 2.44 10 123.5 12. Side Plate/Brace 85.7 85.2 31.2 89614 1762 3957 2798 3.52 2.45 10 82.6 Notes: (a) 1[(3)]] Entry is empty if no SRF is applied.(b)(c)(d)Full penetration weld so that weld factor, WF=1.4.Corrected for support lug contact area per Section 5.3.5.Compensated for mesh quality per Section 5.3.3.(1-6) Number referring to the (3)96 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information SR-P 4.9 4.6 4.3 4 3.7 3.4 3.1 2.8 2.5 Figure 20a. Locations of minimum stress ratios, SR-P<5, associated with maximum stresses at non-welds for EPU operation with frequency shifts. The recorded stress ratio is the minimum value taken over all frequency shifts. The numbers refers to the enumerated location for SR-P values at non-welds in Table 15a. This view shows locations 1 and 3.97 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 20b. Locations of minimum stress ratios, SR-P<5, associated with maximum stresses at non-welds for EPU operation with frequency shifts. The recorded stress ratio is the minimum value taken over all frequency shifts. The numbers refers to the enumerated location for SR-P values at non-welds in Table 15a. This view shows locations 1 and 2.98 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information z SR-P 4.9 4.6 4.3 4 3.7 3.4 3.1 2.8 2.5 Figure 20c. Locations of minimum stress ratios, SR-P<5, associated with maximum stresses at non-welds for EPU operation with frequency shifts. The recorded stress ratio is the minimum value taken over all frequency shifts. The numbers refers to the enumerated location for SR-P values at non-welds in Table 15a. This view shows location 4.99 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information K4~x SR-a 5 4.8 4.6 4.4 4.2 4 3.8 3.6 3.4 3.2 2.8 2.6 2.4 2.2 2 Figure 20d. Locations of minimum alternating stress ratios, SR-a<5, at non-welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-a values at non-welds in Table 15a. View showing locations 1-4.100 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Y SR-P 4 3.8 3.4 3.2 3 28 2.4 2.2 2 1.8 1.6 1.4 Figure 20e. Locations of minimum stress ratios, SR-P<4, associated with maximum stresses at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-P values at welds in Table 15b. This view shows locations 1, 3, 6, 7 and 9-11.101 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information z¥Figure 20f. Locations of minimum stress ratios, SR-P<4, associated with maximum stresses at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-P values at welds in Table 15b. This view shows locations 2, 4, 5 and 10.102 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information z Y ý x Figure 20g. Locations of minimum stress ratios, SR-P___4, at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-P values at welds in Table 15b. This view from below shows locations 4, 5, 8 and 10.103 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 20h. Locations of minimum alternating stress ratios, SR-a_<4, at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-a values at welds in Table 15c. This view shows locations 1, 2, 4, 5 and 7-10.104 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 20i. Locations of minimum alternating stress ratios, SR-a_<4, at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-a values at welds in Table 15c. View showing locations 1, 4-6, 8, 9 and 11.105 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Y SR-a 4 3.6 3.4 3.2 3 Figure 20j. Locations of minimum alternating stress ratios, SR-a.4, at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-a values at welds in Table 15c. View around locations 3, 9 and 11.106 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information I 3 4.6.2.6 Figure 20k. Locations of minimum alternating stress ratios, SR-a_<4, at welds for EPU operation with frequency shifts. The recorded stress ratio at a node is the minimum value taken over all frequency shifts. Numbers refer to the enumerated locations for SR-a values at welds in Table 15c. View around locations 1, 2, 5, 9, 10 and 12.107 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

5.6 Frequency

Content and Filtering of the Stress Signals The frequency contribution to the stresses can be investigated by examining the power spectral density (PSD) curves and accumulative PSDs for selected nodes having low alternating stress ratios. The accumulative PSDs are computed directly from the Fourier coefficients as 1(con) = Yl(cOk2 k=l where 8(ok) is the complex stress harmonic at frequency, (ok. Accumulative PSD plots are useful for determining the frequency components and frequency ranges that make the largest contributions to the fluctuating stress. Unlike PSD plots, no "binning" or smoothing of frequency components is needed to obtain smooth curves. Steep step-like rises in X(c 1) indicate the presence of a strong component at a discrete frequency whereas gradual increases in the curve imply significant content over a broader frequency range. From Parsival's theorem, equality between E(O)N) (where N is the total number of frequency components) and the RMS of the stress signal in the time domain is established.

The selected nodes are the ones having the lowest alternating stress ratios (at a weld) in Table 15. These are: Node 101861 -Entry 1 in Table 15c located on the weld joining the tie bar and inner vane bank top plate. The associated PSDs are shown in Figure 21a.Node 91651 -Entry 1 in Table 15a located where the lifting rod contacts the upper lifting rod brace. The associated PSDs are shown in Figure 21b.Node 94509 -Entry 2 in Table 15c located on the weld connecting the outer hood and its end plate. The associated PSDs are shown in Figure 21c.Node 99200 -Entry 4 in Table 15c located on the weld joining the inner side and base plates.The associated PSDs are shown in Figure 21d.Node 95267 -Entry 6 in Table 15c located on the welded common junction between the outer hood, hood support and outer cover plate. The associated PSDs are shown in Figure 21 e.These nodes are respectively labeled as 1 in Figure 20h, 1 in Figure 20d, and 2, 4 and 6 in Figure 20h-k.In each case, since there are six stress components and up to three different section locations for shells (the top, mid and bottom surfaces), there are a total of 18 stress histories per component.

Moreover, at junctions there are at least two components that meet at the junction.The particular stress component that is plotted is chosen as follows. First, the component and section location (top/mid/bottom) is taken as the one that has the highest alternating stress. This narrows the selection to six components.

Of these, the component having the highest Root Mean Square (RMS) is selected.

For comparison the PSDs and cumulative PSDs are also shown for the CLTP load examined in [11] at the shifts producing the highest stress intensity at that load.108 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information The first node (101861) is dominated by several peaks, with the dominant one centered at 84.5 Hz for the +7.5% shifted case. From the accumulative PSD it is evident that frequency shifting increases this peak, but does not shift its frequency.

This is indicative of a peak in the signal moving closer to a structural resonance.

The second (91651) and fifth (95267) nodes are both dominated by low frequency peaks that do not change significantly with frequency shift.These are associated with lifting rod vibrations that are transmitted to the upper brace and supporting structure on the outer vane bank. For node 94509, a broad peak at 90 Hz and a more narrow one at 106 Hz dominate the response and both differ significantly from the non-shifted curve. Finally, for node 99200 multiple peaks are present that do not shift significantly with frequency shift.Another way to characterize the dominant frequencies is to plot the dominant frequency over the dryer surface. For each finite element node the frequency associated with the largest stress harmonic (at any frequency shift) is recorded.

A contour map of this dominant frequency is shown in Figure 22. This map is useful in a qualitative sense for identifying what dryer components appear most responsive to particular frequencies.

Low frequency responses dominate the vane bank side plates, hood supports, lifting rods and parts of the outer cover plates. For most of the dryer hoods, the dominant frequencies are near 70 Hz as indicated in Figure 22b. The skirt response is dominated by higher frequency responses in the 80-130 Hz range, as are the vane bank top plates, hood side plates, and inner closure plates. It should be noted that loadings above 200 Hz play no role in the dynamic response of this dryer.109 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Node 101861, aw CL W.E:3 1000 800 600 400 200 0 0 50 100 150 200 Frequency

[ Hz ]Node 101861, a 250 0.65 106 105 104 1000 100 10 1 0 50 100 150 200 250 Frequency

[ Hz ]Figure 21a. Accumulative PSD and PSD curves of the cyy stress response at node 101861.110 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Node 91651, a YY C-ci U)C-75 E:3 U)0.U)U)I £uu 1000 800 600 400 200 0 107 108 101 104 1000 100 10 1 0-~----------

• *.... iý .. ... ...... .. ~ ... ... ... ... # .... .- P.~.. ....i! ..-~ No shiJ I---% shf 50 100 150 Frequency

[ Hz]Node 91651, a 200 250 0 50 100 150 200 250 Frequency

[ Hz ]Figure 21b. Accumulative PSD and PSD of the (Yyy stress response at node 91651.111 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Node 94509, a 500.i;E~:3 400 300 200 100 0 101 10 4 0 50 100 150 200 Frequency

[ Hz]Node 94509, a YY 250"I" N.1000 0 a.100 10 0 50 100 150 200 250 Frequency

[ Hz]Figure 21c. Accumulative PSD and PSD of the c;yy stress response at node 94509.112 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Node 99200, a d (0 a-75 E U)U)9N 6 600 500 400 300 200 100 0 106 104 1000 1000 100 10 1-No5shift l_ J_ .~ +.... s .. ..50 0 50 100 150 200 Frequency

[ Hz ]Node 99200, a 2 0 50 100 150 200 250 Frequency

[ Hz ]Figure 21d. Accumulative PSD and PSD of the yzz stress response at node 99200.113 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Node 96267, a.E.E 600 500 400 300 200 100 0 T No shift+10% Sqhift 50 100 150 200 250 0 Frequency

[ Hz Node 96267, %No shift 100 10 1000 10 0 50 A-100 150 200 250 Frequency

[ Hz ]Figure 21e. Accumulative PSD and PSD of the axx stress response at node 95267.114 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information z Y Dom. Freq. [Hz]*50 40*30 20 10 0 Figure 22a. Contour map showing the dominant frequencies (i.e., the frequency with the largest stress harmonic).

This shows locations with dominant frequencies in the range 0-50 Hz.115 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 1 /2Dorm. Freq. [Hz]~JQ1 80 70 60 I 50 Figure 22b. Contour map showing the dominant frequencies (i.e., the frequency with the largest stress harmonic).

This shows locations with dominant frequencies in the range 50-80 Hz.116 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 22c. Contour map showing the dominant frequencies (i.e., the frequency with the largest stress harmonic).

This shows locations with dominant frequencies in the range 80-200 Hz.5.7 Real Time Analysis Without U-Stiffener As discussed in [2] it was established that during power ascension the MSL B loads were affected by the reactor core isolation cooling (RCIC) steam line configuration.

As a result, additional tests and data collections were conducted to define the loads with the RCIC line isolated as described in Appendix C of [2]. In this section the stresses resulting from two additional loads corresponding to different RCIC line configurations are calculated and compared against the baseline EPU loads used in the full steam dryer stress analysis.The first load corresponds to an off-normal condition corresponding to the drain trap out of service. This condition gives rise to a 92.5 Hz peak that was first observed in the MSL B data collected at 110% CLTP. This load was not immediately reproducible, but has occurred numerous times since. The second load definition was collected at 115% CLTP with the inboard 117 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information RCIC valve temporarily closed. This resulted in a narrow peak on MSL B at 89.3 Hz. This also is considered an off-normal condition corresponding to a short-duration technical specification limiting condition or to a transient loading associated with an unusual steam line drain configuration.

In addition to these two additional loads, the reanalysis using the ACM 4.1 R model resulted in new bias and uncertainty values resulting from benchmarking of this model. While it is reasonable to use these values, there has also been interest in calculating the acoustic loads by applying the more conservative of the bias and uncertainty values associated with the ACM 4.1 and ACM 4.1R models in each frequency interval.

Hence there are a total of six different operating condition and bias+uncertainty combinations.

These six loads are summarized in Table 16 together with the limiting stress ratios obtained with the real time evaluation summarized below. The Load Names (corresponding to MSL entrance signal files generated by CDI) are referenced in the subsequent tables (Table 17 and Table 18).In the present section real time stress evaluations are performed for the nodes listed in Table 10 of Section 5.2 using each of the six load combinations.

In the real time assessment the stress intensities are calculated in the same manner as when analyzing the entire dryer and the limiting stresses with frequency shifting included are reported.

The complete list of resulting stress ratios is given in Table 17. Note that one additional node (91651, entry 54) has been added. This location had the limiting alternating stress ratio at a non-weld location (see Table 15a) which is sufficiently close to the 2.0 margin that it was deemed of important to include it.For all load cases, the minimum alternating stress ratio is below 2.0 with the limiting stress locations associated with the {45-53} entry set located near the top of the inner, vane bank/side plate junction.

These locations are addressed in Section 5.8 using the U-section stiffener.

For convenience, Table 17 lists separately the minimum stress ratios for this group (entries 45-53)and remaining points outside this group (entries 1-44 and the added entry, 54). This shows that when entries 45-53 are omitted the alternating stress ratios are acceptable (above 2) for all load combinations except the RCIC valve closed conditions.

In Section 5.8 it is shown that adding the U-section stiffener raises the alternating stress ratios of all entries 45-53 to 2.5 or higher so that set of entries 1-44 and 54 becomes the limiting set.Table 16. Load & Summary for Dryer Without Addition of U-section Stiffener Load Condition Bias & Limiting Stress Ratios Uncertainty Peak, SR-P Alternating, SR-a N55 Baseline (Normal) ACM 4.1R 1.39 1.68 N59 Drain Trap ooS 1.43 1.74 N58 RCIC closed 1.34 1.35 N61 Baseline (Normal) Max {ACM 4.1, ACM 4.1R} 1.37 1.61 N62 Drain Trap ooS " 1.42 1.67 N63 RCIC closed " 1.32 1.29 118 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 17. Alternating stress ratios obtained at 53 locations using: (i) different load and bias+uncertainty combinations.

Entry Node Load N55 Load N59 Load N58 Load N61 Load N62 Load N63 SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a 1 37229 2.5 5.4 2.5 5.9 2.4 4.1 2.5 5.3 2.5 5.7 2.4 4.0 2 113286 4.2 4.5 3.3 3.1 4.0 4.0 4.2 4.5 3.3 3.1 4.0 4.0 3 37592 5.0 2.9 4.9 3.0 4.4 2.5 4.8 2.8 4.7 2.9 4.3 2.4 4 94143 1.4 4.5 1.5 4.9 1.4 3.6 1.4 4.4 1.5 4.8 1.4 3.5 5 85191 1.8 3.3 1.8 3.5 1.7 2.9 1.7 3.2 1.8 3.4 1.7 2.8 6 143795 1.6 3.9 1.7 4.0 1.6 3.0 1.6 3.7 1.7 3.9 1.5 2.9 7 113508 1.9 3.2 2.2 4.0 2.1 4.2 1.9 3.2 2.1 3.9 2.1 4.2 8 98968 1,8 3.1 1.8 3.2 1.6 2.4 1.8 3.0 1.8 3.1 1.6 2.3 9 94498 2.2 4.8 2.1 4.8 2.1 4.7 2.1 4.6 2.1 4.7 2.1 4.5 10 90468 1.8 3.8 2.0 3.6 1.9 3.5 1.8 3.7 2.0 3.5 1.9 3.5 11 101600 2.4 2.5 2.6 2.6 2.1 1.9 2.4 2.4 2.5 2.6 2.1 1.9 12 88639 1.8 3.6 1.8 3.0 1.7 3.0 1.8 3.5 1.8 2.9 1.7 2.9 13 92995 2.5 3.3 2.5 3.2 2.5 3.3 2.5 3.3 2.5 3.1 2.5 3.3 14 141237 2.2 4.0 2.1 3.7 1.8 2.6 2.1 3.9 2.1 3.6 1.8 2.5 15 89317 2.3 6.8 2.1 6.0 1.7 5.3 2.2 6.8 2.1 6.0 1.6 5.3 16 87784 4.0 2.5 4.0 2.6 3.2 2.1 3.9 2.5 3.9 2.6 3.2 2.1 17 94509 3.2 2.1 3.4 2.2 2.7 1.6 3.1 2.0 3.3 2.1 2.6 1.5 18 99635 2.5 2.2 2.6 2.4 2.4 1.9 2.5 2.1 2.5 2.3 2.4 1.8 19 99200 1.6 2.2 1.7 2.5 1.6 2.0 1.6 2.1 1.7 2.4 1.5 2.0 20 113554 1.8 3.5 1.7 2.9 1.7 3.1 1.8 3.5 1.7 2.9 1.7 3.1 21 103080 2.8 2.3 2.9 2.1 2.4 1.7 2.8 2.2 2.9 2.0 2.3 1.7 22 113400 2.6 2.3 3.0 3.0 3.1 3.1 2.6 2.3 2.9 3.0 3.1 3.1 23 95267 1.8 2.3 1.8 2.3 1.7 2.1 1.8 2.2 1.8 2.2 1.7 2.1 24 90786 1.4 2.3 1.4 2.4 1.3 1.8 1.4 2.3 1.4 2.3 1.3 1.8 25 137575 2.0 2.4 2.0 2.2 1.6 1.6 2.0 2.3 1.9 2.1 1.6 1.5 26 95172 2.5 7.9 2.5 6.6 2.2 7.3 2.4 7.9 2.4 6.6 2.2 7.6 27 91055 3.9 2.4 3.5 2.0 2.8 1.8 3.8 2.3 3.4 2.0 2.8 1.7 28 93488 2.8 2.4 3.1 2.9 2.8 2.2 2.8 2.4 3.0 2.8 2.7 2.1 29 98956 2.2 2.5 2.0 2.4 2.0 2.1 2.1 2.4 1.9 2.4 1.9 2.0 30 85631 1.4 2.5 1.4 2.9 1.3 2.0 1.4 2.4 1.4 2.8 1.3 1.9 31 101818 2.6 2.5 2.6 2.9 2.6 2.1 2.6 2.4 2.6 2.8 2.5 2.0 32 98624 2.8 2.5 3.0 2.9 2.6 2.1 2.8 2.4 3.0 2.8 2.5 2.0 119 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 17 (cont.). Alternating stress ratios obtained at 53 locations using: (i) different load and bias+uncertainty combinations.

Entry Node Load N55 Load N59 Load N58 Load N61 Load N62 Load N63 SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a 33 94514 3.6 2.6 4.1 2.7 3.1 1 3.5 2.5 3.9 2.6 3.0 34 95428 1.5 2.6 1.5 2.5 1.4 2.3 1.5 2.6 1.5 2.5 1.4 2.2 35 84090 2.2 2.6 2.3 2.9 2.0 2.3 2.1 2.6 2.3 2.8 2.0 2.2 36 99451 2.3 2.7 2.3 3.1 2.0 2.5 2.2 2.6 2.3 3.1 2.0 2.5 37 93451 5.1 2.7 5.2 2.8 4.4 2.2 4.9 2.6 5.1 2.7 4.3 2.2 38 98172 2.1 2.7 2.1 3.1 2.0 2.6 2.1 2.7 2.1 3.0 2.0 2.6 39 90924 3.8 2.8 3.8 3.1 3.6 2.8 3.8 2.7 3.7 3.0 3.5 2.8 40 100989 4.8 2.8 4.7 2.8 4.3 2.4 4.6 2.7 4.5 2.7 4.1 2.3 41 90926 4.9 2.8 5.2 3.3 4.1 2.5 4.8 2.7 5.1 3.2 4.1 2.4 42 99931 5.1 2.8 5.7 3.1 4.7 2.5 4.9 2.7 5.5 3.0 4.5 2.4 43 91091 1.9 3.0 1.9 3.3 1.8 2.5 1.9 2.9 1.9 3.2 1.8 2.5 44 88702 2.6 6.9 2.5 6.4 2.0 7.0 2.5 6.8 2.5 6.3 2.0 7.1 MIN (1-44) 1.4 2.1 1.4 2.0 1.3 1.6 1.4 2.0 1.4 2.0 1.3 1.5 Locations near top of inner vane bank/side plate/tie bar junction Results WITHOUT U-Section Stiffener installed 45 101861 2.9 3.1 2.4 2.8 L 2.9 2.3 46 95197 3.3 3.2 2.7 3.2 1 3.1 2.5 47 99407 2.4 2.1 2.8 2.4 2.2 2.4 2.1 2.7 2.3 2.1 48 98442 3.1 3.1 2.5 3.0 3.0 2.4 49 98444 3.1 2.9 2.5 3.0 2.8 2.4 50 98451 3.6 3.3 2.6 3.4 1 3.2 2.5 51 98452 2.6 2.5 I 2.1 2.5 1 2.5 2.1 52 99408 3.1 3.3 2.0 2.8 3.0 L 3.2 2.7 53 91240 2.6 2.6 2.1 2.4 2.5 I 2.5 2.0 2.4 MIN (45-53) 2.4 1.7 2.5 1.7 2.1 1.4 2.4 1.6 2.5 1.7 2.1 1.3 Added Real Time Node (Non Weld)544.1 2.2 4.1 2.1 4.6 2.4 12.2 Load N55 Load N59 Load N58 Load N61 Load N62 Load N63 MIN (overall)

SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a SR-P SR-a_______1.4 1.7 1.4 I1.7 1.3 1.4 1.4 1.6 1.4 1.7 1.3 1.3 120 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 5.8 Real Time Analysis Adjusted with U-Section Stiffener Included Section 5.7 shows that for all load combinations the alternating stress ratio is below the 2.0 target level. The limiting stresses all occur on the set of entries, {45-53 1, associated with the top of the inner vane bank/side plate junction also involving the inner tie bar. To address these locations a U-section stiffener is installed 11" from the top edge of the side plate using six bolts.To analyze the stresses with the channel installed a new steam dryer model was developed with the channel attached and unit stress solutions generated over the 60-130 Hz frequency range.The solutions are combined with the six loads in Table 16 and the stresses evaluated at the limiting locations using the real time analysis.

Specifically, the nodes in the new FEA model that are near the set of entries, {45-53}, are identified, and the stresses calculated and appropriately adjusted in accordance with the SRFs for embedded model 6 as described in Section 4.5. For nodes on welds the weld factors are also applied. Finally the stresses and stress ratios are adjusted to account for contributions from the remaining frequencies outside the 60-130 Hz range, Since the steam dryer meshes with and without the U-section stiffener are different it is difficult to develop stresses at the exact same node locations.

However, by taking the maximum stress over all nodes in the same neighborhoods it is straightforward to infer the limiting stress (for each load combination) in the region represented by entries. The associated limiting stress ratios are listed in Table 18 showing that for all but the RCIC valve closed conditions, the limiting alternating stress ratios are above 3.0. For the RCIC valve closed, these stress ratios are 2.4 or higher thus confirming that the U-section stiffener is effective in restoring adequate margin to these locations.

When compared against the minimum stress ratios from the remaining entries, (1-44, 54 1, obtained from Table 17 it is clear that these latter locations now constitute the limiting set. The minimum stress ratios for each load combination with the U-section stiffeners installed are listed in the last row of Table 18. Note that only the alternating stress ratios are listed since the peak stresses are already demonstrated to meet margin (SR-P>l for all load conditions) in Table 17. These show that for the normal (baseline) and drain trap out-of-service conditions the target alternating stress ratios are met. For the RCIC valve closed cases however, the limiting alternating stress ratios remain below 2.0. The nodes where the alternating stress ratios are below 2.0 are identified in Table 17 by the greyed cells associated with entries 1-44.These nodes are evaluated separately in [10] using a cumulative fatigue usage analysis.

This analysis is justified since RCIC valve closure is infrequent so that cycle counts are low and fatigue assessment using cycle counting to derive a cumulative usage factor (CUF) is appropriate.

As shown in [10] the CUFs for all these nodes are demonstrated to be well below 1.0.To summarize, these results show that the dryer meets the target alternating stress ratio, SR-a _2.0, for both normal EPU operation and off-normal operation with the drain trap ooS. For the off-normal RCIC valve closed case, the target alternating stress ratio of 2.0 is not met, but since this condition is infrequent a separate cycle counting analysis is appropriate.

This analysis is performed and shows that the CUF<I. Taken in their entirety, these results show that the steam dryer qualifies for fatigue under all the load cases considered.

Note that this conclusion presumes that the installation of the U-section stiffener does not introduce new high stress locations.

This is confirmed by calculating the stresses at the ends of the channels and showing that these remain well above margin (SR-a>6.5 for all end stresses) for all load conditions.

It 121 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information also assumes that installation of the channel does not significantly alter the stress ratios at other locations (i.e., associated with entries 1-44 and 54). This assumption is reasonable since the channel acts mainly to suppress the vibration amplitude of the side plate rather than altering the frequencies.

Hence the dominant effect of the U-section stiffener is to reduce stresses in the vicinity of the inner side plate, with a lesser reduction farther away.Table 18. Limiting alternating stress ratios with addition of U-section stiffener Entries Load N55 Load N59 Load N58 Load N61 Load N62 Load N63 MIN (Locations 1-44, 54) 2.1 2.0 1.6 2.0 2.0 1.5 Limiting Stress Ratios at Locations 45-53 with U-Section Stiffener Installed MIN (Locations 45-53) 3.4 3.4 2.5 3.3 3.3 2.4 MIN (overall) 2.1 2.0 1.6 2.0 2.0 1.5 122 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

6. Conclusions A re-evaluation of the stresses on the NMP2 steam dryer due to acoustic loads, deadweight and static pressure (the Level A service condition) at EPU (115% EPU) conditions has been performed.

This re-evaluation was necessary due to an identified inconsistency in the previous acoustic loads predictions

[2] and also the detection of additional signals in recently collected data on the MSLSs whose impact on the dryer required assessment.

The acoustic loads are inferred from strain gage measurements on the MSLs and a calibrated acoustic circuit model (ACM, Rev. 4.1 R) that processes these measurements to define acoustic pressure distributions on the steam dryer surfaces.

The ANSYS FEA package is then used to obtain the dryer stress response resulting from these acoustic loads and the stress results post-processed to obtain the limiting alternating stress ratios. The results account for all biases and uncertainties identified for both the ACM Rev. 4. 1R and the FEA harmonic analysis.The acoustic loads are prepared using the acoustic circuit model (ACM) version 4.1 R [2, 8, 9] for three load conditions:

Baseline:

The normal EPU operating load.Drain Trap Out-of-Service:

An off-normal condition occurring when the drain trap in the RCIC system is isolated.RCIC Valve Closed: An infrequent off-normal operating load occurs when the RCIC valve is intentionally closed.For each load two bias and uncertainty combinations were considered.

The first uses the biases and uncertainties of the ACM 4. 1R model and the other uses a conservative combination of the total uncertainties from the ACM 4.1 and ACM 4.1 R models.A complete dryer stress evaluation using the Baseline load identified limiting locations with alternating stress ratios that were below the target of 2.0, but above the ASME allowable of 1.0.These locations were examined in more detail by improving local mesh quality, identifying welds at lifting rod/brace junctions, developing an embedded model, and processing stress singularities at the ends of selected welds. The complete dryer stress evaluation was also used to develop a list of nodes for real time stress evaluation at the other load combinations.

The stress evaluations with frequency shifting at all load combinations identified that the best option for maximizing long term operating margin for the inner side plate connecting the inner vane banks was the installation of the U-section stiffener that is bolted on to each of these inner side plates (2 total). With this modification in place, all nodes have a peak stress ratio, SR-P, of 1.3 or higher at all load combinations thus meeting the required margin for this stress type. With regard to alternating stresses, all of the nodes on the steam dryer have an alternating stress ratio of 2.0 or higher under the baseline and drain trap out-of-service loads so that dryer qualifies for these conditions.

For the RCIC valve closed, certain locations have alternating stress ratios below 2.0 with the minimum value being SR-a=I.5.

However, since this condition occurs infrequently, it is appropriate to assess fatigue using cycle counting.

The cumulative usage factors (CUFs) for these locations are calculated in [10] and shown to all lie well below 1.0.123 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Taken in their entirety, these results show that the dryer qualifies for all of the level A service operation with the U-section stiffener installed.

124 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

7. References

1. Continuum Dynamics, Inc. (2013) Non-Conformance Report (NCR) 03-43. 23 September.

2. Continuum Dynamics, Inc. (2014) Acoustic and Low Frequency Hydrodynamic Loads at 115% CLTP Target Power Level on Nine Mile Point Unit 2 Steam Dryer to 250 Hz Using ACM Rev. 4.1R. C.D.I. Report No.14-09P (Proprietary), December.3. Continuum Dynamics, Inc. (2010) Design and Stress Evaluation of Nine Mile Point Unit 2 Steam Dryer Modifications for EPU Operation.

C.D.I. Report No.10-12P (Proprietary), July.4. Continuum Dynamics, Inc. (2012) Stress Evaluations of Nine Mile Point Steam Dryer Modifications.

C.D.I. Technical Memo 12-03, Rev. 0, January.5. Continuum Dynamics, Inc. (2012) Stress Evaluation of Nine Mile Point Unit 2 Steam Dryer at 115% CLTP. C.D.I. Report No.12-18P (Proprietary), Rev. 0, Oct.6. Westinghouse (2011) NMP2 Steam Dryer Modifications and Repairs. FCN-MODS-NMP2-11, Rev. 1.7. ASME (2007) Boiler and Pressure Vessel Code,Section III, Subsection NG.8. Continuum Dynamics, Inc. (2011) ACM Rev. 4.1: Methodology to Predict Full Scale Steam Dryer Loads from In-Plant Measurements (Rev. 3). C.D.I. Report No.10-09P (Proprietary), November.9. Continuum Dynamics, Inc. (2013) Design Record File DRF-CDI-338 A.10. Continuum Dynamics Inc. (2014) Computation of Cumulative Usage Factor for the 115% CLTP Power Level at Nine Mile Point Unit 2 with the Inboard RCIC Valve Closed.C.D.I. Technical Note No.14-04P (Proprietary), April.11. Continuum Dynamics, Inc. (2011) Stress Evaluation of Nine Mile Point Unit 2 Steam Dryer Using ACM Rev. 4.1 Acoustic Loads. C.D.I. Report No.11-04P (Proprietary), Rev. 0, May.12. Continuum Dynamics, Inc. (2011) Sub-Modeling in the Nine Mile Point Unit 2 Steam Dryer, Rev. 0. C.D.I. Report No.11-03P (Proprietary), June.13. Continuum Dynamics, Inc. (2007) Methodology to Predict Full Scale Steam Dryer Loads from In-Plant Measurements, with the Inclusion of a Low Frequency Hydrodynamic Contribution.

C.D.I. Report No.07-09P (Proprietary).

14. Continuum Dynamics, Inc. (2005) Methodology to Determine Unsteady Pressure Loading on Components in Reactor Steam Domes (Rev. 6). C.D.I. Report No. 04-09 (Proprietary).

15. NRC (2012) Safety Evalution by the Office of Nuclear Reactor Regulation Related to Amendment No. 149 to Facility Operating License No. NPF-69 Nine Mile Point Nuclear Station, LLC, Nine Mile Point, Unit No. 2 Docket No. 50-410. ML1 13560333.16. Structural Integrity Associates, Inc. (2010) Flaw Evaluation of Indications in the Nine Mile Point Unit 2 Steam Dryer Vertical Support Plates Considering Extended Power Uprate Flow Induced Vibration Loading (Rev. 0). SIA Calculation Package No.1000814.401, July.17. Structural Integrity Associates, Inc. (2014) Flaw Evaluation of Indications in the Nine Mile Point Unit 2 Steam Dryer Considering Current Licensed Thermal Power Flow Induced Vibration Loading. Report No. 1400467.401 Revision 0, April.125 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

18. Continuum Dynamics, Inc. (2010) Acoustic and Low-Frequency Hydrodynamic Loads at CLTP Power Level on Nine Mile Point Unit 2 Steam Dryer to 250 Hz Using ACM Rev.4.1 (Rev. 2). C.D.I. Report No.10-10P (Proprietary), January.19. ANSYS, Release 10.0 Complete User's Manual Set, (http://www.ansys.com).

20. Continuum Dynamics, Inc. (2007) Response to NRC Request for Additional Information on the Hope Creek Generating Station, Extended Power Uprate. RAI No. 14.110.21. Continuum Dynamics, Inc. (2008) Stress Assessment of Hope Creek Unit 1 Steam Dryer Based on Revision 4 Loads Model, Rev. 4. C.D.I. Report No.07-17P (Proprietary).

22. Press, W.H., et al., Numerical Recipes. 2 ed. 1992: Cambridge University Press.23. Continuum Dynamics, Inc. (2009) Stress Assessment of Nine Mile Point Unit 2 Steam Dryer at CLTP and EPU Conditions, Rev. 1. C.D.I. Report No.09-26P (Proprietary), December.24. Continuum Dynamics, Inc. (2012) Real Time Monitoring of the Nine Mile Point Steam during Power Ascension.

C.D.I. Technical Note No.12-17P (Proprietary).

June.25. Continuum Dynamics, Inc. (2014) Stiffening Channel Bolt Analysis.

C.D.I. Technical Note No. 14-05, Rev. 1, March.26. Continuum Dynamics, Inc. (2010) Stress Assessment of Nine Mile Point Unit 2 Steam Dryer Using the Acoustic Circuit Model Rev. 4.1. C.D.I. Report No. 10-i1P (Proprietary), June.27. Structural Integrity Associates, Inc. (2008) Flaw Evaluation and Vibration Assessment of the Nine Mile Point Unit 2 Steam Dryer for Extended Power Uprate Operating Conditions.

Report No. 0801273.401.

28. Continuum Dynamics, Inc. (2008) Stress Assessment of Browns Ferry Nuclear Unit 1 Steam Dryer, Rev. 0. C.D.I. Report No.08-06P (Proprietary).

29. O'Donnell, W.J., Effective Elastic Constants For the Bending of Thin Perforated Plates With Triangular and Square Penetration Patterns.

ASME Journal of Engineering for Industry, 1973. 95: p. 121-128.30. de Santo, D.F., Added Mass and Hydrodynamic Damping of Perforated Plates Vibrating In Water. Journal of Pressure Vessel Technology, 1981. 103: p. 175-182.31. Ideichik, I E. and E. Fried, Flow Resistance, a Design Guide for Engineers.

1989, Washington D.C.: Taylor & Francis. pg. 260.32. Continuum Dynamics, Inc. (2007) Dynamics of BWR Steam Dryer Components.

C.D.I.Report No.07-11P.33. U.S. Nuclear Regulatory Commission (2007) Comprehensive Vibration Assessment Program for Reactor Internals During Preoperational and Initial Startup Testing.Regulatory Guide 1.20, Rev. 3, March.34. Weld Research Council (1998) Fatigue Strength Reduction and Stress Concentration Factors For Welds In Pressure Vessels and Piping. WRC Bulletin 432.35. Pilkey, W.D., Peterson's Stress Concentration Factors, 2nd ed. 1997, New York: John Wiley. pg. 139.36. Lawrence, F.V., N.-J. Ho, and P.K. Mazumdar, Predicting the Fatigue Resistance of Welds. Ann. Rev. Mater. Sci., 1981. 11: p. 401-425.37. General Electric (GE) Nuclear Energy, Supplement 1 to Service Information Letter (SIL)644, "B WR/3 Steam Dryer Failure, "September

5. 2003.38. Tecplot, Inc. (2004) Documentation:

Tecplot User's Manual Version 10 Tecplot, Inc., October.126 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

39. Scholl, Rod, Size Matters, in ANSYS Advantage.

2014, ANSYS. p. 50-53.127 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information

[II (3)]](3)]]128 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]129 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]130 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information (3)]]131 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Appendix B. U-Section Stiffener The analysis in Section 5.7 shows that the limiting alternating stress ratios at all load combinations consistently occur on the junction between the top of the inner vane bank, the inner vane bank end plate and the inner side plate connecting the two inner vane banks. The location is depicted in Figure 14b. The high stress is associated with vibration of the inner side plate at approximately 85 Hz. To suppress vibration of the side plates joining the inner vane banks, a U-section stiffener is bolted on to these side plates at a distance (top edge to top edge) of 11" from the top edge of the side plate. A bolted rather than welded attachment is preferred to minimize diver exposure to (radiation) dose during the 2014 RFO installation and also to avoid heat-induced stresses in the materials.

The present Appendix summarizes the design, finite element analysis and stress evaluation of the U-section stiffener.

Design of the U-section stiffener proceeded by first investigating the impact on stress due to addition of a rectangular cross-section strip welded to the inner side plate. To facilitate analysis and simplify combining of unit solutions, the presence of the strip was first modeled as a thickness change of selected elements in the side plate. This leaves the existing finite element mesh and node indexing unchanged so that: (i) the new unit solutions can be swapped in easily (the old ones being simply overwritten) thus allowing the use of existing post-processing software tailored to the Nine Mile Point evaluation, unaltered; and (ii) the effect of the stiffening on the global model, including the changes in the modal properties, are fully accounted for without having to regenerate a complete unit solution sweep using a significantly larger and computationally time consuming model. Several combinations of strip thickness and vertical location were considered.

For each one, a series of unit solutions was generated over a frequency range centered at 85 Hz and the combined with the MSL entrance signals to estimate stresses.

At first it seemed natural to locate the stiffening beam near the top edge of the side plate since this is the elevation where the high stress occurs. However this location proved less effective in suppressing the dominant stress-inducing mode than placing it further down. A 5" high, 0.5" thick strip spanning the horizontal length of the side plate and located 11" below its top edge was found to be most effective in suppressing the stresses at Entry 2 in Table 9c.Next the optimal thickness and location developed in the preceding design process were expressed as an equivalent stiffness and mass distribution that in turn were translated into a U-section stiffening beam design. The final design bolts the U-section stiffener on the interior side (exposed to the steam path) of the side plate using 6 bolts. Bolting the beam to interior side was necessary to avoid interference with the vertical guide rail intended to guide the steam dryer in and out of the RPV during an outage. The final design is depicted in Figure 24 and consists of a 5"xO.5" vertical leg lying on the inner side plate and a pair of 2"xO.65" horizontal legs providing most of the stiffening.

The part is machined out of 304 SS qualified for usage as a steam dryer component.

The differences between the modeled (Figure 24a) and machined (Figure 24b) U-section stiffeners reflect alterations made to the manufactured part during and after analysis to simplify installation and tooling requirements as well as a preference for using a shell element-based representation (to facilitate connection to the side plate and generation of a good quality mesh). However, the cross-sectional second area moments controlling bending in the horizontal plane are in close agreement so that the modeled element model accurately represents the as-built component.

132 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information The final design was analyzed using FEA. The finite element model was developed by first creating a clean mesh on the side plate such that it contains straight lines corresponding to the reinforcement strip edges and the elements are regularly shaped. Figure 25a shows this mesh on the outward side of the inner side plate. Next the channel is modeled using shell elements as shown in Figure 25b. The mesh is designed such that each node in the reinforcement strip is associated with a node on the inner side plate that is located immediately below it. This one-to-one correspondence ensures accurate imposition of the contact conditions between the channel and side plate. Here the two parts are conservatively (since it minimizes net bending stiffness) assumed to be capable of sliding relative to each other. Thus, associated nodes (i.e., nodes on the strip and inner side plate that lie on top of each other) are constrained to have the same deflection in the normal direction.

However, the other degrees of freedom on associated nodes are left unconstrained.

The bolts themselves are represented with 0.5" long beam elements that connect the side plate and U-section stiffener at the cantilevered ends. The exact cross-sectional dimensions of the bolts are not critical since only the shear and tensile forces in these elements are of potential interest.

These forces, when added to the pre-loads associated with bolt tightening, and divided by the bolt cross-sectional area yield the bolt stresses.Unit solutions were generated for this finite element model over the 60-130 Hz frequency interval.

This range encompasses most of the contribution to the stress intensity at the high stress location the channel is intended to address (typically 97-98%; occasionally down to 90%for certain nodes). Signal content below 60 Hz is expected to have marginal impact since it lies below the fundamental frequency of the unrestrained plate (and, accordingly, also below that of the reinforced plate). These unit solutions were combined with the six load combinations defined in Table 16 to calculate the stresses at junction of the top of the inner vane bank and side plate. Specifically, a list of nodes located within several inches of the high stress locations identified in the original (i.e., without the channels) global model, is developed.

The alternating stresses at these locations are then calculated using the six load combinations and the limiting values recorded.Since these results are developed over a reduced frequency interval, 60-130Hz, it is necessary to adjust them to be representative of the entire 0-250 Hz range. This is done by evaluating the stresses in the original global model (i.e., without the U-section stiffener) at the same locations and comparing the stress estimates obtained when using the full interval vs. using only the 60-130 Hz interval.

Specifically, if S(fl,f2) is the stress intensity calculated using only the unit solutions in the frequency interval, fl to f2, then one can calculate an adjustment factor: f= S(60,130) (B. 1)S(0,250)This factor, calculated separately at each individual node and load condition using the original global model, is then used to adjust the stresses at the same location computed in the model with the U-section stiffener.

Thus stresses are obtained by dividing by f and stress ratios via multiplication.

Typically f is near unity (0.98, with occasional values as lows as 0.9) since most of the stress contribution originates in the 60-130 Hz range.133 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Where appropriate, the stress reduction factors developed for embedded model 6 in Section 4.5 (SRF=0.80 and 0.66, depending on the particular weld) are also applied. The final limiting stress ratios for each load combination are reproduced in Table 19.Table 19. Limiting alternating stress ratios for nodes located on or near the common junction between the tops of the inner vane banks, the side plates and vane bank end plates. Stress ratios are calculated for the six load combinations defined in Table 16.Load Limiting alternating stress ratios N55 3.41 N59 3.43 N58 2.49 N61 3.28 N62 3.31 N63 2.39 Evaluation of Stresses at End of Beam The installation of the U-section stiffeners has the potential of increasing stresses locally.Therefore a real time stress analysis is also performed for a collection of nodes located about the ends of the U-section stiffeners where the stresses are most likely to be highest. The stress evaluation is performed using the most limiting load case, outN63 in Table 16, (this load generally gives rise to the highest alternating stresses and also has the highest biases and uncertainties applied).

The limiting alternating stress ratio (taken over all frequency shifts) about the end of the U-channel is SR-a=7.69.

Since only part of the complete frequency range was considered, this estimate is not conservative.

However, from the original global model the 60-130 Hz frequency range contributes approximately 85% of the total stress. Multiplying the stress ratio by 0.85 leads to a good estimate of the full stress ratio as SR-a=7.69x0.85=6.54 which is still well above the recommended margin. The conclusion is that stresses at the end of the U-channel are very low.134 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information I-. 29.00 r 4.63 4.63 7.50 -4.63-- 4.63 ---2.00 5.00.50.65 NOTE: CENTERLINE OF CHANNEL LOCATED 13.5 INCHES BELOW TOP EDGE OF PLATE CRAWN --Marc J. Sibila YAM ' -----3IM2014 Continuum Dynamics, Inc.APPOVEDT Channel Stiffener SI.1 CM NO. OWG NO. REV SCALE SMEET Figure 24a. Drawing of the U-section stiffener

-as designed and modeled in the FEA.135 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information I I STY I NOTE PART TiTUS WACWPTWN I MArItAL I ~C NOTES: 1. CENTERLINE OF CHANNEL LOCATED 13.5 INCHES BELOW TOP EDGE OF PLATE.2. BREAK ALL SHARP EDGES AND CORNERS.3. FOR ADDITIONAL FABRICAT1ON PROCESS REOUIREMENTS.

SEE WESTINGHOUSE SPECIFICATION FS-BWR-ENG-20AI003.

4. FINAL SURFACES TO BE ELECTROPOUSHED AND EVALUATED PER APPENDIX C OF FS-BWR-ENG-20AI003.

5. THIS DRAWING IS MADE FROM CONTINUUM DYNAMICS DESIGN DRAWING T1718-001.

D a C/ \I pflLt-r4L, t a mi~m~~ n, qata I SECTION A-AIAýB DETAIL C-CL 1 B a B ISOMETRIC VIEW-----0.81 TYP DETAIL Bý :1 p IRAEI.ECETAMA4STCE4OM4TTISTTTTAIA 1] ................

I I -.4-- ~A*.I I...............-.

I -~ I 4444.44444 RI'AA I WESTTTTAICATAPITTPITETAR'IALAAT A 0& I-NINI-~~~ ZTSIT M T T.4r~iT I-- A O se NINE IALE POINT NUCLEAR STATION 2 CHANNEL STIFFENER I D.05D78: I I I 4 1 2 1 Figure 24b. Drawing of the U-section stiffener

-as built and installed on the NMP stream dryer.136 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 25a. Finite element mesh for the steam dryer FEA model with the U-section stiffener included.

View from outside of the steam dryer.137 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 25b. Interior and top views of the U-section stiffener installed on the inward face of the inner side plate.138 This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information U-Channel Center Line Stresses In order to design the bolts attaching the U-channel shaped stiffener to the inner side plate, the local stresses in the inner side plate are calculated.

Specifically, the stresses are extracted from a line segment that coincides with the vertical line of symmetry of the inner side plate and extends from the bottom to top of the U-section stiffener, or 11" to 16" from the top edge of the side plate (10" to 15" in the modeled inner side plate). The load used to calculate these stresses is the outN58 signal in Table 16 corresponding to the RCIC valve closed condition.

This load selection is conservative for the ACM 4.1R-based cases since it generally produces higher stresses than the baseline EPU load (outN55) or the condition when the drain trap is out of service (outN59).

A slightly higher load occurs when the higher of the bias and uncertainty estimates from the ACM 4.1 and ACM 4.1R over each frequency interval is used (outN63).Stress calculations using this signal produce stresses at the high stress location that are less than 5% higher than obtained with outN58.The maximum alternating stress intensity on this center line is determined to be 651psi and occurs on the inner side plate. This value is further increased by dividing by 0.85 which estimates the effects of using more conservative load (outN63 rather than outN58 -this accounts for an approximate 5% increase as explained above) and extending the frequency interval from the 60-130 Hz range over which unit solutions are available to the full 0-250Hz frequency range.This range extension is conservatively estimated to increase stresses by 10% based on comparisons of full and partial range calculations carried out using the global model to produce the adjustment factors, f, in (B.1). Using this 0.85 scaling the maximum stress is estimated as 766 psi. This is the value used to size the U-section stiffener bolts in [25].139