ML091610110

CDI Report No. 08-24NP, Rev. 1, Stress Assessments of Nine Mile Point, Unit 2 Steam Dryer, Attachment 13.7, Introduction Through Page 77

ML091610110
Person / Time
Site:	Nine Mile Point
Issue date:	05/27/2009
From:	Bilanin A Continuum Dynamics
To:	Constellation Energy Group, Office of Nuclear Reactor Regulation
References
7708631 08-24NP, Rev 1
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ENCLOSURE ATTACHMENT 13.7 CDI Report No. 08-24NP (Non-proprietary)

Stress Assessments of Nine Mile Point Unit 2 Steam Dryer Nine Mile Point Nuclear Station, LLC May 27, 2009

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information CDI Report No. 08-24NP Stress Assessments of Nine Mile Point Unit 2 Steam Dryer Revision 1 Prepared by Continuum Dynamics, Inc.

34 Lexington Avenue Ewing, NJ 08618 Prepared under Purchase Order No. 7708631 for Constellation Energy Group Nine Mile Point Nuclear Station, LLC P. 0. Box 63 Lycoming, NY 13093 Approved by Alan J. Bilanin Reviewed by Milton E. Teske May 2009 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 finite element model and analysis methodology, used to assess stresses induced by the flow of steam through the steam dryer at Nine Mile Point Unit 2 (NMP2), are described and applied to obtain stresses at CLTP conditions. The analysis is consistent with those carried out in the U.S. for dryer qualification to EPU conditions and complies with a standard analysis procedure [1] supported by the EPRI BWRVIP and currently under review by the USNRC. The resulting stresses are assessed for compliance with the ASME B&PV Code 2007 [2],Section III, subsection NG, for the load combination corresponding to normal operation (the Level A Service Condition).

The 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 scalings in the loads without the need for additional finite element calculations. ((

(3))) 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 damped Helmholtz equation within the steam dryer

[3]. 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.

Results obtained from application of the methodology to the NMP2 steam dryer show that at nominal CLTP operation the minimum alternating stress ratio (SR-a) anywhere on the steam dryer is SR-a=2.92. The loads used to obtain this value account for all the end-to-end biases and uncertainties in the loads model [4] and finite element analysis. in order to account for uncertainties in the modal frequency predictions of the finite element model, the stresses are also computed for loads that are shifted in the frequency domain by +/-2.5%, +/-5%, +/-7.5% and +/-10%.

The minimum alternating stress ratio encountered at any frequency shift is found to be SR-a=2.80 occurring at the -5% shift. The stress ratio due to maximum stresses (SR-P) is dominated by static loads and is SR-P=1.35 both with and without frequency shifts.

Since flow-induced acoustic resonances are not anticipated in the steam dryer, the alternating stress ratios at EPU operation can be obtained by scaling the CLTP values by the steam flow velocity squared, (UEPU/UCLTp) 2=l.1782=1.388. Under this approach, the limiting alternating stress ratio becomes SR-a=2.80/1.388=2.02. For the nodes with the limiting maximum stress ratios at CLTP, the corresponding limiting value at EPU is SR-P=1.28. Given that the alternating stress ratio SR-a obtained at EPU remains above 2.02 at all frequency shifts together with the comparatively small dependence of SR-P upon acoustic loads, the Unit 2 dryer is expected to qualify at EPU conditions.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information In order to achieve these stress ratios, the welds on two components, the closure plates and lifting rod braces, require reinforcement. For the closure plates the top 18 inches of the welds connecting the closure plates to the vane banks and to the hoods are reinforced by adding a weld on the inner side of the closure plate. For the top lifting rod braces, increasing the weld size from 1/4 in to 3/8 in meets the target stress ratio.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Revision. Summary The following change was implemented from Revision 0 to Revision 1:

The uncertainty associated with the FEM modeling idealizations has been increased from 21.51% to 25.26% to correctly reflect the values used in the ACM calculations.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table of Contents Section Page Executive Summ ary ......................................................................................................................... i Revision Summ ary ......................................................................................................................... iii Table of Contents ........................................................................................................................... iv

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

2. Methodology .............................................................................................................................. 3 2.1 Overview ............................................................................................................................... 3 2.2 (( (3))) ............................................................................................................................... 5 2.3 Com putational Considerations ......................................................................................... 6

3. Finite Elem ent M odel D escription ......................................................................................... 9 3.1 Steam D ryer Geom etry ..................................................................................................... 9 3.2 M aterial Properties .......................................................................................................... 12 3.3 M odel Sim plifications ..................................................................................................... 12 3.4 Perforated Plate M odel .................................................................................................. 13 3.5 Vane Bank M odel ............................................................................................................ 15 3.6 Water Inertia Effect on Subm erged Panels .................................................................... 16 3.7 Structural Dam ping .......................................................................................................... 16 3.8 M esh D etails and Elem ent Types .................................................................................. 16 3.9 Connections Betw een Structural Com ponents ................................................................ 16 3.10 Pressure Loading .......................................................................................................... 28

4. Structural Analysis .................................................................................................................... 31 4.1 Static Analysis .................................................................................................................... 31 4.2 H arm onic Analysis .......................................................................................................... 31 4.3 Post-Processing ................................................................................................................... 37 4.4 Com putation of Stress Ratios for Structural Assessm ent .............................................. 37 4.5 Finite Elem ent Sub-m odeling ......................................................................................... 40

5. Results ....................................................................................................................................... 42 5.1 General Stress D istribution and High Stress Locations ................................................. 43 5.2 Load Com binations and A llowable Stress Intensities ................................................... 57 5.3 Frequency Content and Filtering of the Stress Signals ................................................... 78

6. Conclusions ................................................................................................................................ 84

7. References ................................................................................................................................. 86 Appendix A Sub-m odeling of Closure Plates ......................................................................... 88 Sub m odel N ode 101175 ..................................................................................................... 90 Sub m odel node 91605 .......................................................................................................... 97 Sub m odel node 95172 ............................................................................................................ 105 Sub m odel node 100327 .......................................................................................................... 114 Appendix B. Sub-m odeling of Lifting Rod Support Braces ...................................................... 121 iv

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1. Introduction and Purpose Plans to qualify the Nine Mile Point nuclear plant for operation at Extended Power Uprate (EpU) operating condition require an assessment of the steam dryer stresses experienced under the increased loads. The steam dryer loads due to 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 purpose of the stress analysis discussed here is to calculate the maximum and alternating stresses generated during Current Licensed Thermal Power (CLTP) and determine the margins that exist when compared to stresses that comply with the ASME Code (ASME B&PV Code,Section III, subsection NG).

The stress analysis of the modified NMP2 steam dryer establishes whether the existing and proposed modifications are adequate for sustaining structural integrity and preventing future weld cracking under planned EPU operating conditions. 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 [5]. Level B service conditions, which include seismic loads, are not included in this evaluation.

(3))) This approach also affords a number of additional computational advantages over transient simulations including: ((

(3))) This last advantage is realized through the use of "unit" solutions representing the stress distribution resulting from the frequency. ((

application of a unit fluctuating pressure at one of the MSLs at a particular (3)))

This report describes the overall methodology used to obtain the unit solutions in the frequency domain and how to assemble them into a stress response for a given combination of pressure signals in the MSLs. This is followed by details of the NMP2 steam dryer finite I

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information element model including the elements used and overall resolution, treatment of connections between elements, the hydrodynamic model, the implementation of structural damping and key idealizations/assumptions inherent to the model. Post-processing procedures are also reviewed including the computation of maximum and alternating stress intensities, identification of high stress locations, adjustments to stress intensities at welds and evaluation of stress ratios used to establish compliance with the ASME Code. The results in terms of stress intensity distributions and stress ratios are presented next together with PSDs of the dominant stress components.

In order to meet target EPU stress levels (i.e., an alternating stress ratio of 2.0), two locations required modification: the closure plate welds and the top lifting rod support braces. In the former case, an additional weld is placed on the interior side of the junction where the closure plate meets the hood or vane bank. In the latter case, the existing 1/4 in weld is increased to 3/8 in. Both modifications were designed using highly detailed solid element-based sub-models of these locations.

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2. Methodology 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 [6] of the fluctuating pressures in the MSLs, predicts the acoustic pressure field anywhere inside the steam dome and on the steam dryer [1, 3-5]. The ACM is formulated in frequency space and contains two major components that are directly relevant to the ensuing stress analysis of concern here. ((

(1)

(2) 3

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((I (3)))j 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 no 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=10 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, o~i and 0o i+l, 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 [8].

Solution Management 6

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

StructuralDamping 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 dampingnmatrix, D, is set to 2z.

SD=2Kco (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, (ok, are shifted to (0+20)k, 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 conditions hold simultaneously: (i) the acoustic signal contains a significant signal at co*; (ii) the structural model contains a resonant mode of natural frequency, ,n,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 (0* 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 7

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.

(3)))

Evaluationof Maximum and AlternatingStress 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 109 calculations per node each requiring the determination of the roots to a cubic polynomial. ((

(3)))

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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 (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 KG1 -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.

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.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 0.00 100.00 (in) 50.00 Figure 1. Overall geometry of the Nine Mile Point Unit 2 steam dryer model.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 2. Modify the figure to eliminate inner hood strips. On-site modifications accounted for in the model and associated geometrical details.

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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 1.

Table 1. 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 5501F. 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 lbm/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 [11] has been performed to determine whether: (i) they will propagate further into the structure and (ii) their influence uponstructural response frequencies and modes must be explicitly accounted for. To establish (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 05 intensity. For*

05 the USR and skirt cracks the highest stress intensity factors are 1.47 ksi-in and 2.75 ksi-in respectively; both values are below the threshold value (3 ksi-in° 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 ,1/2ithe 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 considers 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:

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" Perforated plates were approximated as continuous plates using modified elastic properties designed to match the static and modal behaviors of the perforated plates. The perforated plate structural modeling is summarized in Section 3.4 and Appendix C of [9].

" 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).

S(3)))

• 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 [12], 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 2 as ratios to material properties of structural steel, provided in Table 1. 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 [9] 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 2) using the. analytical formula for a cantilevered plate and the modified Young's modulus and Poisson's ratio given by O'Donnell [12]. The measured and predicted frequencies are in close agreement, differing by less than 3%.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 2. Material properties of perforated plates.

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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 MASS21. 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 15

This Document Does Not Contain Continuum Dynamics, Inc-. Proprietary Informnation approximating their inertial wpropertigswith 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, thepadded -hydrodynamic mass. This, added mass was found by a separate hydrodynamic analysis (included in DRF-C-279C supporting this report) 1to be.0.143 lbm/in2 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 ibm/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 Da'mping - ,

Structural damping wfas -defined as t1/%:O6f critical damping for all' frequencies.: This* damping:

is consistent with guidance given on pg. 10 of NRC RG-1.20 [16].

3.8 Mesh Details and' EIlemel Tys "' " - , . ...

She'll ele'ments 'were' employed to'-mondel*th eskilt, ,hoods, perforated plates; side 'and end plates,.tough bottomrplates, reinforcements, -base plates and cover plates., ýSpecifically,'the four-node, Shell Element SHELL63, was selected to model these structural,'confponents. '"- This elemehtfimodels'bending and membrane *stressis,butt.d0mits> transverse' sh"ear; 'The use of shell elements is' appropriate -for, most of the structure where the characteristic 'thickness isý small compared to -the other plate dimensions. For 'tiicker structures, such "as the upper, and-lower support r-ings; s'olid biick 'elemehifs w*ere used to, ip6vid 'tie 'full 3D 'stress. lThe elements SURF 154 are used to assure proper application of pressureflbadinfgtothe"striuctutre. Mesh'details and element t'ypes' are shown in Table 3 and Table 4:'

The mesh is generatedaufoinatically by ANSYS-with refinement near e'dges. 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 Andrlthe' hoods. Details of the -finiite element 'mesh are 'shbwn in Figure -5. Numerical experiments carried' out' using the-ANSYS- code "applied i'to simple:

analytic*lly tractable plate structures with dimensions aind mesh 'spacings similai to the ones 'use'l.

for the"'sfeam drier; confirm that the 'natural frequencies' are accurately'recovered (less than 1°/6r..

errors forthe first modes).' These errors are compensatedif6r bylthe'use- 9 f frequency shifting.- '. '"

3.9 Connections Between Structural Components

'Most .co'innections'betwee'n parts are modeled'as node-to'-node' connections. This -is 'th6 correct manner (i.e., within the finite element framework) of joining eltements away .fromfi" discontinuities. At joints between shells, this approach omits the additional stiffness provided by theextra weld material. Also' locall"y 3D: effects are rior e',-jroinoun~ced." The latier:effect is accounted 'for using weld factors'. The deviation in stiffness'due to weld materialis negligible, since weld dimensions are on the order of the shell'thickness. The'ýcon se'quences upon modal 16i

ThisD6cuinent Does, Not Contain Continu* Dynamics,'Inc. Proprietary Information frequiencies -an8 amphitude are, to first order,- pr6portional to t/L, Where t' is the ihickness§ and' L a chafacteristic-shell length. The errors committed by ignoring additional. weld'stiffness air thus small and'readily'coimpensated 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 'bthdiSplacemrtent anid rotationaihtl 'degrees of freedomn at every node -whereas 'solid elemients model only ,the translations. A node-to-node connection'wobuld effectively aippear to the shell element as a simply supported, rather than '(the correct) canitilevered restraint and significantly alter the dynamic iespo'nse of the shell structure.

To address this problem, constraint equations are -used'to pr6perly connect adjacent shell- and solid-elemfient modeled struCtires. Basically, all, such constrairits&expres sthe deflection (and rotatio:i'f&f shellfelements) of a'.dode, RI ¢ on one structural component in terms .of the."

deflecti6ois/rotations of the corresponding- point, P2 , onýi the 6ther icornected component.

Specifically, the element containing P2 is identified and the deformations at P2 ,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 rotatiopal degrees of freedom -when multiple shell:nodes on, a, sufficiently dense grid. are connected to the same solidface. . ,

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, ofrjotations 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 ýuseful 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,. i.iUtiplier-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 matrices (Lagrange multiplier implementations) with poorer conyergence behaviorcompared 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 tthe UsR so that i100% of the, dryer weight is transmitted through the seismic blocks per the FDDR KGl.-265. Because the contact region between the blocks and steam dryer is small, the seismic blocks are considered free to" -

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This Document Does Not Contain Continuum Dynamics, hic. Proprietary Information 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.

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

Point masses Masses are 7 connected to top and bottom supports Gussets to lifting rods connections A Skirt to support rings connections Simply supported restraints

/

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 ((

Table 3. FE Model Summary.

Description Quantity Total Nodes 1 159,793 I Total Elements 1 124,496 1

1. Not including additional damper nodes and elements.

Table 4. Listing of Element Types.

Generic Element Type Name Element Name ANSYS Name 20-Node Quadratic Hexahedron SOLID1 86 20-Node Hexahedral Structural Solid 10-Node Quadratic Tetrahedron SOLID 187 10-Node Tetrahedral Structural Solid 4-Node Elastic Shell SHELL63 4-Node Elastic Shell Mass Element MASS21 Structural Mass Pressure Surface Definition SURF 154 3D Structural Surface Effect Damper element COMBIN14 Spring-Damper 19

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5a. Mesh overview.

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

'Mp -1 A

Figure 5b. Close up of mesh showing on-site modifications.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5c. Close up of mesh showing drain pipes and hood supports.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5d. Close up of mesh showing node-to-node connections between various plates.

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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.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 5f. Close up view of tie bars.

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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.

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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.

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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 [3].

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 alternating stress intensities for assessment in Section 5.

3))) This is useful since revisions in the loads model do not necessitate recalculation of the unit stresses.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information NODES AN PRE S-NORN

-. 101592_ 330 .04099 6.68738 331i .. Y 1 .397447 .540029 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.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information NODES AN PRES-NORM

.394932

.295671 .494193 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.

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This Document Does Not Contain 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, a6m = - Um, are considered for all possible pairs of the stresses co, and am at different time levels, tý and tin. 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 500 psi are rejected. For each remaining stress difference tensor, the principal stresses S 1, S2 , S 3 are computed and the maximum absolute value among principal stress differences, Sm = max IS -S21,IS] -S31,IS2 -S3I , 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 Snm . This alternating stress is compared against allowable n,m values, depending on the node location with respect to welds.

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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

.505E-03 .015 .061254

.008099 ,.05366 .068847 Figure 9a. Overview of static calculations showing displacements (in inches). Maximum displacement (DMX) is 0.069". Note that displacements are amplified for visualization.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information AN 01000 W, 005000 Figure 9b. Overview of static calculations showing stress intensities (in psi). Maximum stress intensity (SMX) is 9,598 psi. Note that displacements are amplified for visualization 33

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 9c. Stress singularities. Model is shown in wireframe mode for clarity.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information NODAL SOLUTION AN STEP=1185 SUB =1 FREQ=50.418 REAL ONLY SINT (AVG)

DMX =.195193 SMN =.081579 SMX =11642

.081579 1294,2 8 9055 10348 11642 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).

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information AN NODAL SOLUTION STEP=305 SUB =1 FREQ=200.446 REAL ONLY SINT (AVG)

DMX =.021716 SMN =.177944 SMX =5801

.177944 644.744 12 rp"-4512 5157 5801 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.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information, 4.3 Post-Processing IThe 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.

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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 5) 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 III, subsection NG.

Table 5. 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 AlternatingStress 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 5 by 0.55.

" For alternating stress intensity, the calculated weld stress intensity is multiplied by a weld stress intensity (fatigue) factor of 1.8, before comparison to the Sa value given above.

The weld factors of 0.55 and 1.8 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, [17], and stress concentration factors at welds, provided in [18] and [19]. In addition, critical welds are subject to periodical visual inspections in accordance with the requirements of GE SIL 644 SIL and BWR VIP-139 [20]. Therefore, for weld stress intensities, the allowable values are shown in Table 6.

These factors (0.55 and 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.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 6. Weld Stress Intensities.

Type Notation Service Limit Allowable Value (ksi)

Maximum StressAllowables.

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

Peak = Primary + Secondary + F 'Salt Sa 13.6 Comparison of CalculatedandAllowable 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 sshells),

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 / (1.1

• Salo,

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) / 1.8.

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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 6, Sm=16,900 psi and Sa=13,600 psi. The factors, 0.55 and 1.8, are the weld factors discussed above. 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 [2], the effect of elastic modulus upon alternating stresses is taken into account by multiplying alternating stress Salt at all locations by the ratio, E/Emodell. .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 1)

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 [21]). These nodes are tabulated and depicted in the following Results Section.

4.5 Finite Element Sub-modeling In order to meet target stress levels at EPU in the NMP2 steam dryer, weld reinforcements are required at two locations: (i) the top 18" of the welds connecting the closure- plates to the hoods and vane banks and (ii) the weld between the vane bank side plates and top lifting rod support brace. These, reinforcements are developed using high resolution solid element-based sub-models of these locations. The use of localized sub-models is motivated by. the need to maintain computational costs at a feasible level. To this end the global steam dryer model is predominantly comprised of shell elements. These elements are well suited for structures such as the steam dryer consisting of shell-like components and tend to produce conservative estimates of the stresses. In some cases however, such as welded junctions involving multiple components, shell element models can overestimate the nominal stress intensities in the vicinity of the junctions. In such cases a more refined analysis using solid elements to capture the complete 3D stress distribution, is warranted. Therefore, to efficiently analyze complex structures such as steam dryers, a standard engineering practice is to first analyze the structure using a shell-based model. Locations with high stresses are examined in greater detail using 3D solid elements to obtain a more definitive stress prediction.

Both locations were examined using detailed 3D solid element sub-models as reported in Appendix A. Based on these models, the nominal stress intensities computed by the 3D solid element model are lower than those obtained with the shell-based FEA used to analyze the complete steam dryer by the factors summarized in Table 7. The stress intensities predicted by the shell element-based analysis at these locations are therefore first multiplied by these factors to obtain more accurate estimates of the nominal stresses. These are then multiplied by the 1.8 weld factor before comparing against allowables to obtain the alternating stress ratios.

For the closure plate the welds connecting the closure plate to the vane banks and hoods experience significant vibratory stresses due to a plate response in the 125-135 Hz frequency range. Though stresses remain well above allowable levels for all frequency shifts at both CLTP and EPU, the margin is below the target level (i.e., a stress ratio of SR-a=2.0 at EPU). Therefore a sub-model was developed for each of the locations on the closure plates where stresses 40

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information exceeded target levels. On each closure plate there are four such locations. The first two are on the vertical weld joining the closure plate to the vane bank. The first node is at the top of this weld and the second one lies 13.5" below it. The other two locations are on the curved weld connecting the closure plate to the curved hood. Again the first location is at the top of this weld and the second one lies 14.5" below it. In both cases, the stresses at the top location result from a combination of membrane and bending stresses whereas the stresses at the lower locations are predominantly due to bending. The stresses are induced by a closure plate response dominated by a (1,2) mode (i.e., the mode shape resembles the first mode of a beam in the horizontal direction and the second mode in the vertical sense) which explains the high stress at the lower locations on the welds. Sub-model calculations at these locations show that to achieve the required target stress levels, an interior weld must be added along the top 18" of each weld thus effectively converting it from a single-sided to a double-sided fillet weld along this length.

Additional details are given in Appendix A.

The second location occurs in the top lifting rod support brace where it connects to the vane bank side plate. In the full model a CLTP alternating stress ratio of SR-a=2.02 is predicted at the

+10% frequency shift. A sub-modeling analysis of the high stress location shows that for the current 1/4" double-sided fillet weld the stress reduction is minimal. Repeating the sub-model analysis with an increased weld of 3/8" results in a stress reduction factor of 0.72. Hence, to meet EPU target stress levels it is recommended to increase the weld to this size.

Table 7. Summary of stress reduction factors obtained using sub-model analysis.

Location Stress Reduction Factor

1. Top of vertical closure plate/vane bank weld 0.62

2. 13.5" below location 1 on the same weld 0.88

3. Top of closure plate/hood weld 0.86

4. 14.5" below location 3 on the same weld 0.71

5. Lifting rod support brace/vane side plate junction 0.72 (assuming an increased 3/8 "weld)

Note: For locations 1-4 it is assumed that an inner weld has been to the top 18" of the welds joining the closure plate to the hoods or vane banks, thereby replacing the existing single-sided fillet weld by one that is double sided.

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This Document Does NotContain Continuum Dynamics, Inc,. Proprietary information 5..Results.,

The stress intensities and associated stress ratios 'iesulting,-fromn, theý Rev. 4 acoustic/hydrodynamic loads [4] with associated biases and uncertainties factored in, are presented below. The bias'due to finite frequency idiscretization, and uncertainty associated with the .finite element model itself, -are also, factored in. ,In: the "following sections the ,highest rnaximum.and alternating stress.--intensities are presented to indicate., which points on 'the dryer experience) significant .stress, concentration and/or ,modal response .(Section 5.1). 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 also reported (Section 5.2). Finally the frequency dependence of the stresses at nodes experiencing the lowest stress ratios is depicted in the form.of accumulative PSDs,(Section 5.3).

In each section results are presented both at nominal conditions (no frequency shift) and with frequency shift included. Unless specified otherwise',freqiuency shifts aregenerally'perfoi-med at 2.5% increments.; ý,The tabulated

' Ostresses and stress ,ratiosl are obtained, using a,, 'blanking!

procedure. that is designed, tq,,preyent, 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 The relevant sress itensies are first computed at every node and then nodes sorted

'rlinws.

according to stress level. The higheststreiss node is noted and'all ieighboring nodes within 10 inchs 'of the highest stiress' node and ts symmetric images'(i.e., reflectidns across the x=0 and y=0 planes) are "blanked" (i.e.,' ekcluded frohm' the search for' subsequet hbigh stress locations)

Of the remaining nodes, the next highest .stress node is identified and its neighbors (closer than 10 inches) blAnkied. The third highiest stress node is simiarly located ani thesearch continiued in thi s fashion unitil' all nodes are eitfier bltanked ori have srsiress kle thandhalf the highest value, on the', ýtii t re For ",:'e, -. - ilo aj .- 'ý"' ,ý ýý , ýIs,, .. ..

thestructure. Forstress rats, an hal"analogous,blanking procedure s-apphied.' Thus the lowest stress ratio of a partic ,ular type min a 1-0" neighborhood and itss symmetric, images is identified ard all otheriode m these region excluded from istig i the'tabl. Of the reaining nodes, the one with the lowest I tess th is reported I'ad its, neighboring points si*niilarly excluded, and so oniui'ntil all no6des are either blanked diorhae a stri6es rati hih`ghr t 'an4' The measured CLTP strain gage signals contain significant contributions, from non-acoustic sources such as sensor noise, MSL turbulence and pipe bending vibration that contribute to the hoop strain measurements. The ACM analysis does no$t distinguish betweenthe,'acoustic and non-acoustic fluctuations in .the MSL. signals that: could ilead'-to,4sizeable,.butfictitious acoustic loads&and, resulting ýstresses'onthe dryer. One. way toremove these' fictitious loads is to collect data with the system maintained ýat operating pressure (1000 ,psi).and temperature, but low power

[22].. By operating, the.recirculation ,pumps at this, condition, the jbackground.plant,. noise. and vibrations,-remain 4present., At.,these conditions the acousticý loads, are known ,to be negligible, so thatý collected: dataý referred to. as the low, power, data;, originate ,entirely from non-acoustic sources such as. sensor noise and mechanical ývibrations.,,' Thisinformation is. valuable, since it allows oneto now, distinguish between the acoustic and, non-acoustic.. content in the CLTPsignal and, therefore, modify the CLTP loads 'so that- only, the, acoustic; component is: retained. , For consistency, the. low: power strain, gage signals:are, filtered in ýthe same manner as'the CLTP data and are fed into the ACM model to obtain the (( .. *i] signals at,,the MSL inlets. Since there is negligible flow, these signals are fictitious, i.e., the hoop strains measured 42

This Document Does NotfContain ContinUd Dynamics, inc. Proprietary Information by the strain gages are not due to pressure -fluctuations, but rather due to noise. However, under the supposition that these signals are acoustic in origin the hypothetical stresses due to these signals can nevertheless be computed.

The contribution, of ,background' noise in, theý Nine Mile Point, Unit;2 steam-dryer was quantified by taking strain" gage' measurements at. 25% power. At, this level there' are ýno significant, acoustic sources i since these scale, as, velocity,. *and hence power, squared, .; To compensate for ithe non-acoustic noiseý source'represented in, the low powerdata, the CLTP MSL inlet pressure'signals are, modifiedaccording to [1,22]:

[ 0.5 (8) wheref isAthe frequency (in Hz), P0 (f) is the MSL inlet pressure (( , (3))) at CLTP conditionts, before; correction,P(f) is the corresponding post-correction pressure and N(f) and P0(f' re the pressue amplitud with the lowvpowerdata pissoower, and CLTP' data respectively. The noise subtraction procedure is, identical to thiat used in other steam dryer analyses'(e.g., [23]) 'and'oUtlined'in the:LTR [11]. No that this modification:'leaves the 'phase information in the original CLTP signal unchanged. Note also that thevaluie of 13=0.5 used here is conservative relative to the value of 13=0.8 recommended in [1].

The appliedload includes all biases and iincertainties fr both the ACM (summarized in [41) and the FEM. For, the lattert arethee main contributors to the bias and uncertainty. The first is an .uncertainty (25.26%) that accounts, for modeling idealizations (e.g., vane bank mass model), geometrical approximations and other disciepancies between the modeled and actual dryer such as neglectingof weld mass and stiffness in the FEA. The'second contributor is a bias of 9.5.3% 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% '.'

5.1 General Stress DiStribution an4 High Stress Locations, -.

The maximum stress intensities* obtained bypost-processing the ANSYS stress histories for CLTP at nominal frequency and'with',frequency shift operating conditions ,are' listed in, Table. 8 Contour plots of the stress 'intensities8POver the steam, dryer structure are .shown. on ,Figure 11.,

(nominal frequency) and ,Figure 12 (maximum stress,; over allt-nine frequency shifts including nominal).: The figur'es are oriented to emphasize the ,high' stressý regions. ,'Note that these, stress intensities do notaaccount foruweld factors but; include"'end-to-end bias and,'uncertaintyý. ,Further; it should be noted that since 'the a'llowable? stresses vary~with 'location, stress',intensities! do not necessarily correspond to, regions of primary structural concemrn.' Intead; structural evaluation, is more accurately made in'terms /,of the stress. ratios which compare the computed 'tresses to allowable levels with due account 'made for' stress' type and weld. Comparisons on the basis of stress ratios aremadei in Section 5.2.

43

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information The maximum stress intensities in most areas are low (less than 500 psi). For the membrane stresses (Pm) the high stress regions tend to occur at: (i) 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); (ii) the welds joining the tie bars to the top cover plates on the vane banks; (iii) the seismic blocks that rest on the steam dryer supports; and (iv) junctions connecting the bottoms of the hood supports. Except for the last location, the stresses are dominated by the static contribution as can be inferred from the small alternating stress intensities (Salt) tabulated in Table 8 for the high Pm locations. From Figure 1 a and Figure 12a higher Pm regions are seen to be in the vicinity of the supports where all of the dryer deadweight is transmitted, the closure plates connecting the inner hoods to the middle vane banks, and various localized concentrations such those along the bottom of the outer hood.

The membrane + bending stress (Pm+Pb) distributions evidence a more pronounced modal response especially on the hood structures. The two locations with the highest stress intensities of this type are the same pair having the highest membrane stress and is dominated by deadweight. High stress concentration is also recorded on the top edge of this vertical plate where it joins to the inner vane bank. Other areas with high Pm+Pb stress concentrations include: (i) the tops of the closure plates where they are welded to a hood or vane bank end plates; (ii) the skirt/drain channel welds; (iii) the outer cover plates connecting to the upper support ring and bottom of the outer hoods; and (iv) the common junction between each hood, its hood support (or stiffener), and the adjoining base plate (see Figure 12c).

The alternating stress, Salt, distributions are most pronounced on the outer hoods directly exposed to the MSL inlet acoustics, and on welds involving the closure plates. All hoods exhibit a strong response (e.g., Figure 12d). The highest stress intensity at any frequency shift occurs at the weld joining the inner hood, hood support and base plate. This location, and similar ones involving the bottoms of the hood supports, are localized as indicated in Figure 12e. These locations have emerged as high stress locations in other steam-dryers also. A finite element substructure analysis [23] where the junction and associated welds are modeled using fine resolution solid elements indicates that the stresses are lower than those (conservatively) predicted with the shell element model here. Other locations with high alternating stress intensities include the tie bar/top cover plate weld and welds involving the closure plate.

Comparing the nominal results (Table 8a) and results with frequency shifting it can be seen that maximum stress intensities, Pm and Pm+Pb, do not differ significantly. The highest alternating stress is approximately 4.2% higher when frequency shifts are considered. For other nodes however the variations are higher. As shown in the next section, all stresses are well within allowable levels.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 8a. Locations with highest predicted stress intensities for CLTP conditions with no frequency shift.

Stress Location Weld Location (in) node(a) Stress Intensities (psi)

Category x y z Pm Pm+Pb Salt Pm Inner Side Plate No 3.1 119 0.5 37229 7445 8745 380 Side Plate Ext/Inner Base Plate Yes 16.3 119 0 94143 6861 9722 368 Upper Support Ring/Support/Seismic Block Yes -6.9 -122.3 -9.5 113554 6268 6268 943 Tie Bar Yes 49.3 108.1 88 141275 5797 5797 729 Hood Support/Middle Base Plate/Inner Yes -39.9 0 0 85723 5769 5944 1772 Backing Bar/Inner Hood Pm+Pb Side Plate Ext/Inner Base Plate Yes 16.3 119 0 94143 6861 9722 368 t Inner Side Plate No 3.1 119 0.5 37229 7445 8745 380 Side Plate/Top Plate Yes 49.6 108.6 88 93256 2463 8314 1021 Middle Base Plate/Inner Backing Bar Out/Inner Yes 39.9 108.6 0 85631 424 6999 1140 Backing Bar/Inner Hood Outer Cover Plate/Outer Hood Yes 102.8 -58.1 0 94498 1027 6989 717 Salt Side Plate/Brace (c) Yes 79.7 85.2 75.8 89649 2021 3257 2350

" Hood Support/Middle Base Plate/Inner Yes 39.9 0 0 88639 5533 5754 2173 Backing Bar/Inner Hood

" Side Plate/Top Plate Yes 81.1 -85.2 88 91055 1131 5108 2063

" Middle Hood No -68.6 69.6 43.7 31149 1255 2053 2041

" Side Plate/Brace Yes 79.7 85.2 53.5 89652 2026 3226 2040 Notes.

(a) Node numbers are retained for further reference.

(b) Appropriate stress reduction factor for the welds on the closure plate listed in Table 7 have been applied.

(c) Stress reduction factor (0.72) for the top-most lifting rod braces has been applied.

45

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 8b. Locations with highest predicted stress intensities taken over all frequency shifts CLTP conditions.

Stress Location Weld Location (in) node(a) Stress Intensities (psi) % Freq.

Category x y z Pm Pm+Pb Salt Shift Pm Inner Side Plate No 3.1 119 0.5 37229 7488 8935 561 10

" Side Plate Ext/Inner Base Plate Yes 16.3 119 0 94143 6894 9731 415 5

" Upper Support Ring/Support/Seismic Block Yes -6.9 -122.3 -9.5 113554 6728 6728 1398 10 it Tie Bar Yes -49.3 -108.1 88 143795 5834 5834 759 5 I Hood Support/Middle Base Plate/Inner Yes -39.9 0 0 85723 5769 5944 2449 0 Backing Bar/Inner Hood Pm+Pb Side Plate Ext/Inner Base Plate Yes 16.3 119 0 94143 6894 9731 415 5

" Inner Side Plate No 3.1 119 0.5 37229 7488 8935 561 5 Side Plate/Top Plate Yes 49.6 108.6 88 93256 2481 8314 1021 0 Outer Cover Plate/Outer Hood Yes 102.8 -58.1 0 94498 1080 7184 893 -10 Middle Base Plate/Inner Backing Bar Yes -39.9 -108.6 0 84197 433 7163 1242 5 Out/Inner Backing Bar/Inner Hood Salt Hood Support/Middle Base Plate/Inner Backing Yes -39.9 0 0 85723 5769 5944 2449 -5 Bar/Inner Hood Side Plate/Brace (c) Yes 79.7 85.2 75.8 89649 2312 3367 2448 10 Side Plate/Brace Yes 79.7 85.2 53.5 89652 2279 3710 2446 2.5 Side Plate/Closure Plate/Exit Mid Top Perf Yes -78.5 85.2 70.5 101873 498 2566 2431 -10

" Side Plate/Top Plate Yes 81.1 -85.2 88 91055 1148 5335 2332 5 Notes.

(a) Node numbers are retained for further reference.

(b) Appropriate stress reduction factor for the welds on the closure plate listed in Table 7 have been applied.

(c) Stress reduction factor (0.72) for the top-most lifting rod braces has been applied.

46

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 1 la. Contour plot of maximum membrane stress intensity, Pm, for CLTP load. The maximum stress intensity is 7445 psi.

47

This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information m I z

xk Pm+Pb [psi]

9000 8250 7500 6750 6000 5250 4500 3750 3000 2250 1500 750 0

Figure 1 lb. Contour plot of maximum membrane+bending stress intensity, Pm+Pb, for CLTP load. The maximum stress intensity is 9722 psi. First view.

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

Pm+Pb [psi]

9000 8250 7500 6750 6000 5250 4500 3750 3000 2250 1500 750 0

Figure 1 c. Contour plot of maximum membrane+bending stress intensity, Pm+Pb, for CLTP load. This second view from below shows the high stress intensities at the hood/stiffener/base plate junctions and drain channel/skirt welds.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure l Id. Contour plot of alternating stress intensity, Salt, for CLTP load. The maximum alternating stress intensity is 2350 psi. First view.

50

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

x Figure IIe. Contour plot of alternating stress intensity, Salt, for CLTP load. Second view showing details of the outer hood and closure plate.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 12a. Contour plot of maximum membrane stress intensity, Pm, for CLTP operation with frequency shifts. The recorded stress at a node is the maximum value taken over all frequency shifts. The maximum stress intensity is 7488 psi.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 12b. Contour plot of maximum membrane+bending stress intensity, Pm+Pb, for CLTP operation with frequency shifts. The recorded stress at a node is the maximum value taken over all frequency shifts. The maximum stress intensity is 9731 psi.

First view.

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

Figure 12c. Contour plot of maximum membrane+bending stress intensity, Pm+Pb, for CLTP operation with frequency shifts. This second view from beneath reveals stresses on the hood support/base plate junctions, outer cover plate and drain channel/skirt welds.

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

Salt [psi]

2500 2250 2000 1750 1500 1250 1000 750 500 250 0

Figure 12d. Contour plot of alternating stress intensity, Salt, for CLTP 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 2449 psi. First view.

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

z Salt [psi]

2500 2250 2000 1750 1500 1250 1000 750 500 250 0

Figure 12e. Contour plot of alternating stress intensity, Salt, for CLTP operation with frequency shifts. The recorded stress at a node is the maximum value taken over all frequency shifts. Second view from below.

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This Document Does Not Contain Continuum Dynamics, Inp. Pr9prietary Information 5.2 Load Combinations and Allowable Stress Intensities The stress ratios computed for -CLTP at nominal fre~quency and with frequency shifting are listed in Table 9.,. 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; are also .depicted in Figure 13 (no frequency shift) and Figure 14 (all frequency shifts included). -The plots corresponding to maximum stress intensities depict all nodes with stress ratios. SR-P<4, and the plots of alternating stress ratios display all nodes with SR-:a_4. .. * -

For CLTP operation at nominal frequency the minimum stress ratio is identified as a maximum stress, SR-P=1.35, and is recorded on the bottom of the vertical plate joining the innermost vane banks. However, this location is only weakly responsive to acoustic loads as can be seen from the high alternating stress ratio at this location (SR-a>16.5 at all frequency shifts).

This is true for all three nodes having the lowest values of SR-P, all having SR-a>4.9 at all frequency shifts. The minimum alternating stress ratio at zero frequency shift, SR-a=2.92, occurs on the weld connecting the upper lifting rod brace to the vane bank end plate.

The effects of frequency shifts can be 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 resulting stress ratios are then processed as before to identify the smallest stress ratios anywhere on the structure, categorized by stress type (maximum or alternating) and location (on or away from a weld). The results are summarized in Table 9b and show that the lowest stress ratio, SR-P=1.35, occurs at the same location as in the nominal case and retains virtually the same value. Moreover, the next four lowest SR-P locations are the same as in Table 9a (the third location has a different node index but is a mirror image of the same location). The lowest alternating stress ratio, SR-a=2.80 occurs at the common intersection point of the bottom of the inner hood, hood support and base plate (see Figure 14d). It is worth noting that sub-modeling analysis of a similar node in another steam dryer [23] indicates that the alternating stresses are overestimated at the limiting node (by a factor of 1/0.79 or 27%). Sub-modeling was not pursued for this location however, since 100%

margin at EPU is already met with the current stress predictions. Hood supports are also involved in locations 11, 13, and 14. The next lowest SR-a location involves the lifting rod support brace (Figure 14 g) involving locations 2 and 3. The remaining low alternating stress ratio locations occur on: (i) closure plates (locations 4, 6-8 and 12); (ii) tie bar ends or their immediate vicinity (locations 5, 9 and 10).

The estimated alternating stress ratio at EPU operation is obtained by scaling the corresponding value at CLTP by the square of the ratio of the steam flow velocities at EPU and CLTP conditions. Since this ratio, (UEpu/UcLTp)2= .1782=1.388, the limiting alternating stress ratio at any frequency shift for EPU is estimated as SR-a=2.80/1.388=2.02. This value qualifies the Unit 2 dryer at EPU conditions with considerable margin. The stress ratio, SR-P, is dominated by the static load and has a weaker dependence on power. When the nodes in Table 9b associated with the limiting SR-P at welds are reanalyzed with the MSL signals increased by 1.388, the limiting SR-P reduces to 1.28 at EPU. The limiting value occurs on the second location (node 113554) due to the stronger contribution from acoustic loads. For location 1 (node 94143) SR-P reduces from 1.35 at CLTP to 1.33 at EPU.

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This-Docurfient D6es Not Contain Continuum Dynamics,,Inc. Proprietary Information:

In summary, the lowest alternaiating stress ratio occurs at the base of the inner hood Supp6rt where it is welded to the middle base plate and vertical vane bank support.- Its'value, SR-a=2.80 at the -5% frequency shift indicates that stresses are well'below allowable level. The loWest,'

stress ratio associated with a maximum stress is SR-P= .35 at' CLTP. This value is dominated by the static component and is only weakly altered by acoustic loads (it reduces-to 1.28 at EPU).

Since acoustic loads scale' roughly' with the square of the steam flow, -the ,limiting alternating.

stress ratio at EPU reduces to 2.02, which given that the applied loads already account for all end-to-end biases and uncertainties, still contains ample margin for sustained EPU operation.

44 t

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 9a. Locations with minimum stress ratios for CLTP conditions with no frequency shift. Stress ratios are grouped according to stress type (maximum - SR-P; or alternating - SR-a) and location (away from a weld or at a weld). Bold text indicates minimum stress ratio of any type on the structure. Locations are depicted in Figure 13.

Stress Weld Location Location (in.) node(a) Stress Intensity (psi) Stress Ratio Ratio x y z Pm Pm+Pb Salt SR-P SR-a SR-P No 1. Inner Side Plate 3.1 119 0.5 37229 7445 8745 380 2.27 32.54

" 2. Thin Vane Bank Plate -15.6 -118.4 0.6 2558 4741 5153 <250 3.56 >14

3. Support/Seismic Block 10.2 123.8 -9.5 113286 4347 4347 1300 3.89 9.51 SR-a No None (All nodes have SR-a > 5)

Y i lteEPae* . 9 - 06.3......................... 36.... . . 6..

2. Upper Support Ring/Support/Seismic Block -6.9 -122.3 -9.5 113554 6268 6268 943 1.48 7.29

3. Tie Bar 49.3 108.1 88 141275 5797 5797 729 1.6 9.42

4. Hood Support/Middle Base Plate/Inner -39.9 0 0 85723 5769 5944 1772 1.61 3.88

....... BackingBar/Inner Hood

___._ 5 Inrer Side&Plate/Inner Base Plate -2.3 '-11i9- -0 99200- '4421 7858 460 1.77 14.92

6. Closure 'Plate/Ihner Backing Bar Out/Inner 39.9 108.6 0.5 93062 5190 5209 811 1.79 8.47

,. Backing Bar/Inner Hood 7.-Hood:Support/Middle Base Plate/Inner -39.9 59.5-- 0 -90468 5145,- 5239 1312 1.811 5.23 S- ...... Backing Bar/Inner Hood " *.. ' "- - - _

8. Hood Support/OuterCoverPlate/0uterHood -- 102.8 28.4 W' 95267 5081 5118 1812 -1.83 3.79

- 9. Thin Vane Bank Plate/Hood-Support/Outer- -87 28.4- - -0 98956- 5072' '5161 1810- 1.83- 3.79

'- --- -Base-Plate ....- . " - - - -

10. Outer-Cover Plate/Outer Hood .. 102.8 -58.1 .0 944987 1027 6989 -717 1.99 9.59 N otes. .. . . .. . . ...-- ,-. . ..

(a) Node numbers are retainedTor further reference.

(b) Appropriate stress reduction factor for the welds on the closure plate listed in Table 7 have been applied.

(c) Stress reduction factor (0.72) for the top-most lifting rod braces has been applied,. -

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 9a (cont.). Locations with minimum stress ratios for CLTP conditions with no frequency shift. Stress ratios are grouped according to stress type (maximum - SR-P; or alternating - SR-a) and location (away from a weld or at a weld). Bold text indicates minimum alternating stress ratio on the structure. Locations are depicted in Figure 13.

Stress Weld Location Location (inm) node(a) Stress Intensity (psi) Stress Ratio Ratio x y .z Pm Pm+Pb Salt SR-P SR-a SR-a Yes 1. Side Plate/Brace (C) 79.7 85.2 75.8 89649 2021 3257 2350 4.28 2.92

2. Hood Support/Middle Base 39.9 0 0 88639 5533 5754 2173 1.68 3.16 Plate/Inner Backing Bar/Inner Hood _ _-

" " 3. Side Plate/Top Plate 81.1 -85.2 88 91055 1131 5108 2063 2.73 3.33

.. .. 4. Side Plate/Brace 79.7 85.2 53.5 89652 2026 3226 2040 4.32 3.37

..,,__ __"__,5. Double Side Plate/Top Plate: '49.3 0 88 93197 1155 2791 200.1 5 3.43

. " 6.-Double'Side.Plate/TOp Plate 17.6 0 88 95617. 1155 2832 1995 4.92 3.44

__" "_"_7.Hood Support/Inner Hood 38 0 36.9- 99522 790 : 1973 1960 7.07 3.50 Notes.

(a). Node numbers are retained for further reference. " m (b) Appropriate stress reduction factor. for the welds on the closure plate listed inTable 7 have been applied.

(c) Stress reduction factor (0.72) for the top-most lifting rod braces has been Applied.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 9b. Locations with minimum stress ratios for CLTP conditions with frequency shifts. Stress ratios at every node are recorded as the lowest stress ratio identified during the frequency shifts. Stress ratios'are grouped according to stress type (maximum - SR-P; or alternating - SR-a) and location (away from a weld or at a weld). Bold text indicates minimum stress ratio of any type on the structure. Locations are depicted in Figure 14.

Stress Weld Location Location (in.) node(a) Stress Intensity (psi) Stress Ratio  % Freq.

Ratio x y z Pm Pm+Pb Salt SR-P SR-a Shift SR-P No 1. Inner Side Plate 3.1 119 0.5 37229 7488 8935 561 2.26 22.02 10 ti .. 2. Support/Seismic Block 10.2 123.8 -9.5 113286 4914 4914 .2076 3.44 .5.96 10 if "it 3. Thin Vane Bank Plate -15.6 -118.4 0.6 2558 4792 5204 <250 3.53 >13 10 SR-a No None (All nodes have SR-a > 5)

SR-P~~'-ýO 6 9 094143--

e,. ý6894 9731ý "ý415i 5 1.4 5

.. .. 2. USR/Support/Seismic Block -6.9 -122.3 -9.5 113554 6728 6728 1398 1.38 4.91 10

3. Tie Bar -49.3 -108.1 88 143795 5834 5834 759 1.59 9.05 5

4. Hood Support/Middle Base Plate/Inner -39.9 0 0 85723 5769 5944 2449 1.61 2.8 0 Backing Bar/Inner Hood

5. Inner Side Plate/Inner Base Plate -2.3 -119 0 99200 4458 7994 612 1.74 11.22 5

6. Hood Support/Outer Cover Plate/Outer -102.8. 28.4 0 95267 5271 5347 -2017 1.76 3.41 5 Hood

7. Hood Support/Middle Base Plate/Inner 39.9 -59.5 0 101435 5250 5449 1640 1.77 4.19 -10 Backing Bar/Inner Hood

8. Closure Plate/Inner Backing Bar Out/Inner 39.9 108.6 0.5 93062 5190 5209 811 1.79 8.47 0 Backing Bar/Inner Hood ,

9. Thin Vane Bank Plate/Hood -87. 28.4 0 98956 5072 5161 1810 1.83 3.79 0 Support/Outer Base Plate

10. Outer Cover Plate/Outer Hood 102.8 -58.1 0 94498 1080 7184 893 1.94 7.69 -10

" ,__ 11. Side Plate/Top Plate 17.6 119 88 91215 888 6968 .1258 2 5.46 5 Notes.

(a) Node numbers are retained for further reference.

(b) Appropriate stress reduction factor for the Welds on the closure plate listed in Table 7 have been applied.

(c) Stress reduction factor (0.72) for the top-most lifting rod braces has been applied.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Table 9b (cont.). Locations with minimum stress ratios for CLTP conditions with frequency shifts. Stress ratios at every node are recorded as the lowest stress ratio identified during the frequency shifts. Stress ratios are grouped according to stress type (maximum

- SR-P; or alternating- SR-a) and location (away from a weld or at a weld). Locations are depicted in Figure 14.

Stress Weld Location Location (in.) node(a) Stress Intensity (psi) Stress Ratio  % Freq.

Ratio x y z Pm Pm+Pb Salt SR-P SR-a Shift SR-a Yes 1. Hood Support/Middle Base Plate/ -39.9 0 0 85723 5769 5944 2449 1.61 2.80 -5 Inner Backing Bar/Inner Hood (d)

2. Side Plate/Brace (c) 79.7 85.2 75.8 89649 2312 3367 2448 4.02 2.81 10

3. Side Plate/Brace 79.7 85.2 53.5 89652 2279 3710 2446 3.76 2.81 2.5

" 4. Side Plate/Closure Plate/Exit -78.5 85.2 70.5 101873 498 2566 2431 5.43 2.83 -10 Mid Top Perf

" _" 5. Side Plate/Top Plate 81.1 -85.2 88 91055 1148 5335 2332 2.61 2.95 5

6. Closure Plate/Inner Hood (e) 28.8 -108.6 87 95172 1710 4500 2155 3.1 3.19 10

7. Side Plate/Closure Plate/Exit -47.1 108.6 70.5 92863 1311 2622 2132 5.32 3.22 -7.5 Mid Top Perf

8. Closure Plate/Middle Hood -64.6 85.2 68.6 91603 536 2110 2085 6.61 3.29 -10

. . 9. Double Side Plate/Top Plate -49.3 0 88 97693 1023 2829 2082 4.93 3.3 2.5

" 10. Double Side Plate/Top Plate 17.6 0 88 95617 1155 2832 2037 4.92 3.37 2.5

11. Hood Support/Outer Cover -102.8 28.4 0 95267 5271 5347 2017 1.76 3.41 5 Plate/Outer Hood

.12. Side Plate/Closure Plate/Exit -47.1 108.6 72.5 90201 1123 2357 1996 5.92 3.44 -7.5 Mid Top Perf (f)

13. Thin Vane Bank Plate/Hood -87 -28.4 0 98950 3563 3766 1978 2.61 3.47 2.5 Support/Outer Base Plate

14. Hood Support/Inner Hood 38 0 36.9 99522 790 1973 1960 7.07 3.5 0 Note S.

(a) Node numbers are retained for further reference.

(b) Appropriate stress reduction factor for the welds on the closure plate listed in Table 7 have been applied.

(c) Stress reduction factor (0.72) for the top-most lifting rod braces has been applied.

(d) Detailed (sub-model) analysis of this location at another plant [23] indicates that the margin is 27% higher (or SR-a=3.54). No submodeling of this location was pursued because 100% margin at EPU is already met at this node.

(e) Stress reduction factor (0.86) for the closure plate/inner hood connection has been applied.

(f) Stress reduction factor (0.88) for the closure plate/inner hood connection has been applied.

62 0

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

SR-P (no weld) 4 3.8 3.6 3.4 3.2 3

2.8 2.6 2.4 2.2 Figure 13a. Locations of nodes with stress ratios, SR-P<4, associated with a maximum stress at non-welds for nominal CLTP operation. Numbers refers to the enumerated locations for SR-P values at non-welds in Table 9a.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information 6%d Figure 13b. Locations of smallest stress ratios, SR-P<4, associated with maximum stresses at welds for nominal CLTP operation. Numbers refer to the enumerated locations for SR-P values at welds in Table 9a. This view shows locations 1, 3, 6 and 10.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information R=' L--j Figure 13c. Locations of minimum stress ratios, SR-P_<4, associated with maximum stresses at welds for nominal CLTP operation. Numbers refer to the enumerated locations for SR-P values at welds in Table 9a. This view shows locations 2, 5 and 9.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 13d. Locations of minimum stress ratios, SR-P<4, associated with maximum stresses at welds for nominal CLTP operation. Numbers refer to the enumerated locations for SR-P values at welds in Table 9a. This view shows locations 4 and 7-10.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 13e. Locations of minimum alternating stress ratios, SR-a<5, at welds for nominal CLTP operation. Numbers refer to the enumerated locations for SR-a values at welds in Table 9a.

Locations I and 3-6 are shown.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 13f. Locations of minimum alternating stress ratios, SR-a<5, at welds for nominal CLTP operation. Numbers refer to the enumerated locations for SR-a values at welds in Table 9a.

Locations 2 and 7 are shown.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14a. Locations of minimum stress ratios, SR-P_<4, associated with maximum stresses at non-welds for CLTP 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 9b.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14b. Locations of minimum stress ratios, SR-P<4, associated with maximum stresses at welds for CLTP 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 9b. This view shows locations 1, 3, 8, 10 and 11.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14c. Locations of minimum stress ratios, SR-P<4, associated with maximum stresses at welds for CLTP 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 9b. This view shows locations 2, 3, 5, 6 and 11.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14d. Locations of minimum stress ratios, SR-P_<4, at welds for CLTP 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 9b. This view from below shows locations 4 and 6-9.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14e. Locations of minimum alternating stress ratios, SR-a_<5, at welds for CLTP 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 9b. This view from below shows locations 1, 8, 11, 13 and 14 all on hood welds.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14f. Locations of minimum alternating stress ratios, SR-a<5, at welds for CLTP 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 9b. This view shows locations 2-7, 9, 10 and 12.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14g. Locations of minimum alternating stress ratios, SR-a_<5, at welds for CLTP 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 9b. Close-up view showing locations 2 and 3.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14h. Locations of minimum alternating stress ratios, SR-a_<5, at welds for CLTP 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 9b. Close-up view around locations 4, 7, 8 and 12.

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This Document Does Not Contain Continuum Dynamics, Inc. Proprietary Information Figure 14i. Locations of minimum alternating stress ratios, SR-a<5, at welds for CLTP 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 9b. Close-up view round locations 5 and 6.

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