th Ave Tower

550 SOUTH HOPE STREET, SUITE 1700 • LOS ANGELES, CALIFORNIA 90071 • TEL 213-362-0707 • FAX 213-688-3018 • WWW.NYASE.COM

L O S A N G E L E S • S A N F R A N C I S C O • I R V I N E

September 10, 2019 Mr. Larry Liu Sr. Structural Plans Engineer City of Seattle Department of Construction and Inspections P.O. Box 34019 Seattle, WA 98124 Re: Lateral Peer Review of 2208 4th Ave Tower 2208 4th Ave, Seattle WA Review Completion (17184.00) Dear Larry:

Nabih Youssef Associates (NYA) and Marshall Lew (AMEC Foster Wheeler) have been engaged to provide an independent structural design review of the structural design of the lateral force-resisting system for the new residential tower located at 2208 4th Ave in Seattle, WA. This letter is to provide our opinion on the complete lateral system structural design, including both the foundation and the superstructure.

Project Summary

The overall 2208 4th Ave project consists of an approximately 307’ tall residential tower over 3 below-grade parking levels. The typical floorplates of the tower are roughly rectangular in shape, with an expanded podium floorplate at Level 6 and below. The tower, due to its height and lateral system type, utilizes a non-prescriptive performance-based design (PBD) approach and is therefore the subject of this peer review.

Typical tower construction consists of 7.5” thick post-tensioned concrete slabs supported by concrete columns and an interior reinforced concrete shear wall core, which also comprises the lateral force resisting system. Below grade, slabs are typically 10” reinforced concrete, and a complete perimeter concrete basement wall is provided.

The project-specific design criteria provides lateral performance equivalent to that required by the use of the Seattle Building Code (SBC). The EOR establishes this through the use of a two-tier design including an elastic service-level analysis along with a nonlinear dynamic analysis at the MCE-R shaking level. A design-basis-earthquake (DBE) analysis, consistent with current prescriptive SBC requirements with approved exemptions, was also performed. A primary reference used for the MCE-R level analyses was the PEER Tall Buildings Initiative (TBI 2017) document. Project-specific design-level exemptions to the code include:

• The building height (307’) exceeds the maximum permitted by Seattle Building Code and ASCE 7-10 for reinforced concrete bearing wall systems

• The redundancy factor ρ will be set equal to unity for the design-level analysis

• Member capacities and demands are determined using the procedures described in the project-specific BOD (per the 2017 TBI guidelines) in lieu of ASCE 7-10 Ch 16

• Specifically related to the use of a condition mean spectrum (CMS) approach to the development of ground motions, the most current methodology for selection and modification of ground motions per ASCE 7-16 are referenced

• The response modification coefficient has been assigned R=6 for a bearing wall system



• Overstrength factor per ASCE 7 Table 12.2-1 was not used, and instead demands at the MCE level seismic event determined from the nonlinear response history analysis were used to evaluate the strength of critical elements to which the overstrength factor would traditionally be applied.

Collected Information We have received and reviewed the following information:

Geotechnical: 1. Site Specific CMS Memo (01/08/19; 01/08/19) 2. Ground Motions Procedures Memo (07/16/18; 08/06/18; 09/18/18) 3. SLE and MCEr Spectra Memo (08/27/18; 10/01/18; 10/09/18) 4. Ground Motions for NLRHA (03/01/19) 5. Geotech Report (05/22/17)

Structural Calculations:

1. Basis of Design (06/25/18; 07/18/18; 08/14/18; 12/21/18; 04/05/19; 06/04/19; 07/03/19)

• Supplemental Exhibits 2, 10, 21 (08/14/18)

• Supplemental Exhibits 2-1, 2-2, 2-5, 4, 6, 8 (12/21/18) 2. Calculations - DE/SLE Submittal (12/21/18)

• Supplemental Exhibits 4, 18, 26, 36, 40, 43, 44 (02/15/19) 3. Calculations – MCE Nonlinear Response History Analysis (04/05/19; 06/04/19; 07/03/19)

• Supplemental Foundation Information (04/22/19; 06/11/19)

• Supplemental Exhibits 46, 49, 50, 52, 53, 57, 60, 62, 63, 64, 68, 71, 73, 76 (01/25/19)

• Supplemental Exhibits 17, 82, 83, 85, 87, 88, 90, 89 (06/11/19)

• Supplemental Exhibits 46R1, 48, 51, 64R1, 70 (07/03/19)

• Supplemental Exhibits 46, 51, 62, 81, 86, 90, 93, 95, 96 (08/20/19)

Structural Drawings:

1. Kickoff Reference Drawings (06/25/18) 2. Progress DE/SLE/Wing Drawings (12/21/18) 3. MCE Design Drawings (04/05/19) 4. Foundation Submittal Drawings (04/22/19) 5. Updated MCE Design Drawings (06/04/19) 6. Updated MCE Design Drawings (07/03/19)



Recommendations Throughout this review, we prepared a log of comments which we recorded on a standard form. We submitted these to the EOR, who responded back with discussion and supplemental materials. The process also included multiple conference calls.

A copy of the most recent version of this comment log is attached for your reference. As indicated in the log, all 96 structural comments and 35 geotechnical comments discussed during the review are resolved to the peer reviewer’s satisfaction.

On the basis of the above, the design will provide seismic performance results equivalent to those obtained by the use of conventional structural systems designed per prescriptive requirements of the SBC. In addition, lateral wind performance was also reviewed as part of confirmation of the lateral system.

The review is now complete, and this letter serves as the final summarizing document.

Please call us at 213.362.0707 if you have any questions.

Sincerely,

NABIH YOUSSEF & ASSOCIATES

Nabih Youssef, S.E. Principall

2208 4th Ave - Structural Design Peer Review Comment Log9/10/2019

Reviewers: Engineer of Record:

Nabih Youssef KPFF

Documents Received:

1) Initial Basis of Design dated 05/29/18

2) Updated Basis of Design dated 07/16/18 and associated meeting minutes from 07/16 Kickoff Meeting

3) Updated Basis of Design dated 08/14/18 and responses to 1st round of BOD comments

4) Linear Design Submittal (12/21/19; 02/15/19)

5) MCE Submittal #1 (04/05/19)

6) Foundation Submittal (04/22/19)

7) MCE Responses #1 (06/04/2019)

8) Additional MCE Responses & Updated Foundation Calcs (06/11/2019)

9) Additional MCE Responses & Updated Diaphragm Calcs(07/03/2019)

10) Additional MCE Responses (08/08/2019) and (08/20/2019)

# Date of

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Design Team Response Related

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Date of

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Resolution/

Notes

1 Basis of Design

(07/16/18)

Superimposed Dead

Loads

Pg. 9

07/30/18 Please provide the typical superimposed dead loads that will be utilized for the project. Once available,

please also provide a floor map of the applied loads throughout the height of the tower. Also, provide

indication of what type of cladding load is applied at each floor (if not typical)

08/14/18 Superimposed dead load map will be provided in the drawing set when it is developed.

09/14/18 1) Please provide dead load map when available

2) Prior to drawings being created that provide the dead load map, please include in the BOD the

typical superimposed dead loads that will be utilized for various types of areas. For example, what

superimposed dead loads were used in order to calculate the DBE base shear currently noted in the

BOD in Section 7.6?

12/21/18 1) Dead load maps are provided on drawing sheets S0.11 through S0.14

2) Typical floor and roof dead loads have been added to the basis of design in Section 6.

01/31/19

Comment

Resolved

2 Basis of Design

(07/16/18)

Analysis Procedures

Pg. 11

07/30/18 A service-level (SLE) analysis is not currently included in the BOD.

As noted in PEER TBI Section 2.2.3, the TBI guidelines are intended to result in a design with seismic

performance capability at least equal to that intended for similar buildings designed in full

conformance with the prescriptive requirements of ASCE 7. In order to meet this objective using an

alternative design procedure, PEER provides performance objectives as noted in the commentary of

Section 2.2.3, which includes "Demonstrate that Risk Category II buildings are capable of essentially

elastic response and limited structural damage under Service-Level Earthquake (SLE) shaking having a

return period of 43 years".

While it is understood that a linear DBE (code-based, with exceptions) analysis will be undertaken as

required by Seattle SDCI, this may not necessarily satisfy the performance objective mentioned above,

given the DBE analysis utilizes a different hazard level and implies a level of ductility (given R=6).

Therefore, in order to meet the requirements of PEER, please provide either an SLE analysis per the

requirements of the PEER TBI guidelines, or provide substantial justification that the performance

objective of PEER, demonstrating that there is essentially elastic response and limited structural

damage under a 43yr event, is met.

Per the 07/16 kickoff meeting, it is our understanding that an SLE response spectrum will be developed

for comparison to the code based design. Confirming that the code-based design will control all

aspects of design over SLE will satisfy this request.

08/14/18 A SLE spectrum developed by Geoengineers has been applied to the building model following the procedures of the PEER TBI

Guidelines. The resulting base shears are compared with the base shears from the DE analysis considering the different acceptance

criteria for deformation controlled and force controlled actions and we have found that the SLE forces do not govern the design of

any of the structural elements. Supporting calculations and the preliminary SLE spectrum are included in the comment folder.

#2

09/14/18 Comparison of global demands as provided to justify no SLE analysis is acceptable, given the following

notes/confirmations:

1) Please provide the SLE analysis model for review once available

2) Per Comment #1, it is not clear what seismic weight was used to determine the current SLE vs. DBE

base shear comparison. Once clarified, please confirm the base shear comparisons maintain the same

results (DBE controls).

3) Will all force-controlled elements be designed for the estimated MCE-level demands (2.5xDBE)? Are

there any exclusions?

4) Please clarify how it is intended to perform the sensitivity check per PEER TBI Section 6.2.2.

5) Although the base shear comparison noted above appears adequate to justify lack of SLE evaluation,

a similar comparison for drift is not possible given different stiffness modifiers. Given that a SLE

analysis model has been created in order to determine base shear, please also provide the drift results

from this analysis to confirm PEER TBI Section 5.6.1 is satisfied.

12/21/18 1) The SLE analysis model is being provided with the linear design package.

2) The seismic weights used in the 8-14-18 calculations for the DE and SLE base shears differ by less than 1% per the calculations in

the comment folder. The previous base shear comparisons maintain the same results that the DE controls the linear design of the

building structure.

3) Force controlled elements are designed for the results of the nonlinear analysis in Perform 3D and the design criteria presented

in the basis of design. Initial proportioning of DE forces multiplied by a factor is used to generate initial members sizes to use in the

analysis models until the MCE forces are obtained.

4) The torsion sensitivity check is performed with DE level forces and stiffness assumptions. Because this check is a ratio of

displacements, it shouldn't significantly affect the results if the check is performed with SLE or DE level forces.

5) The drift values from the SLE model show that SLE level drifts are 0.10% or less, which are within the prescribed limits of TBI

section 5.6.1. See attached pdf for additional information.

Comment 2-

1.pdf

Comment 2-

2.pdf

Comment 2-

5.pdf 01/31/19

Comment

Resolved

Topic

& Reference

Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)3 Basis of Design

(07/16/18)

Submittals

Pg. 11

07/30/18 While it is not noted in Section 7.3, is there any intention to provide a nonlinear model input submittal

prior to the full MCE submittal? This would be a smaller package summarizing the input, properties,

and strategies for the nonlinear model along with the model itself without results - similar to what is

provided in the BOD Appendix B but with more information and specific to this project. This is often

recommended so that any issues can be resolved prior to having already fully analyzed/post processed

the models.

08/14/18 At this time we do not plan on providing a nonlinear model submittal prior to the full MCE submittal. Some nonlinear element

calibration and modeling information is included in the Basis of Design document and additional calibration information will be

provided at the time of the linear submittal.

09/14/18

Comment

Resolved

4 Basis of Design

(07/16/18)

MCE Stiffness Modifiers

- Backstay

Pg. 12

07/30/18 Although the extent and configuration of the parking ramps is not currently well understood, and it is

acknowledged that this layout may still be in progress, please provide (or plan on providing, once

available) justification that the parking ramps will not contribute to lateral stiffness and therefore do

not need to be included in the analysis model. This would apply to the linear and nonlinear models.

08/14/18 Evaluation of the relative stiffness of the model with and without ramps will be provided at the time of the linear submittal.

Because the ramps connect elements below the transfer level/seismic base of the tower where elements are designed to be mostly

elastic during a seismic event, the results from the linear ETABS model should be able to be extrapolated to the nonlinear Perform

3D model.

09/14/18 Pending receipt of later study 12/21/18 A comparison model with the ramps modeled is included with the linear design package. Per the referenced PDF, several key

response parameters have been compared between the DE model and the ramp model and shown to have negligible differences.

The biggest changes are to the shear in the core wall below level 1, which are ~5% or less. However, the core has been designed for

the shear just above level 1 and shear reinforcing at this level will be carried down through the parking levels. Consequently, as

these values are significantly lower than the design value the variation is not considered significant to the overall design.Comment

4.pdf

01/31/18 It is agreed that for the aspects noted, there appears to be little sensitivity to inclusion of the ramps. A

few follow-up questions:

1) BOD Section 7.6 notes that openings for the parking ramps in the semi-rigid diaphragms will be

modeled so that the stiffness of the diaphragms isn't overestimated - however in the provided ETABS

model "v7", it appears that the opening is only provided at L1 - please clarify.

2) Section 7.7 notes that the ramps will not be modeled in the nonlinear model as well. Please confirm

that the associated mass of the ramps will still be represented.

3) Please clarify how the ramps will be treated in terms of diaphragm design (e.g. will they treated as

openings/continuous diaphragm, etc..?)

02/15/19 1) Ramps have been added to the model and do not affect the design loading. Previously this was done as a side study (2208 4th

Avenue - v7 Ramps.edb) but has since been incorporated into the primary model (2208 4th Avenue - v8a.edb).

2) Section 7.7 has been modified to include the statement, "Mass of the spiraling ramps will be modeled by applying additional

tributary point masses at the ends of the ramps."

3) Ramps are treated as openings in the overall diaphragm design. Ramps will be designed locally as simple spans between their

supports at each level.

Comment 4 03/05/19

Comment

Resolved

5 Basis of Design

(07/16/18)

General DBE Design 07/30/18 Please clarify if any amplifications to the code based DBE analysis will be voluntarily included, in

anticipation of the larger MCE demands.

08/14/18 Deformation controlled elements are designed for the DE force levels that are not amplified beyond the minimum base shear levels

of ASCE 7-10. Force controlled elements are designed for the results of the MCE analysis and the acceptance criteria described in

the Basis of Design. Force controlled elements are initially proportioned using a multiplier on the DE forces based on our

experience with similar buildings but the final design of these elements won't appear in the drawings or calculations until after the

MCE analysis.09/14/18

Comment

Resolved

6 Basis of Design

(07/16/18)

Core Wall Nonlinear

Modeling

Pg. 17

07/30/18 It is noted that the core wall fiber sections will be modeled with an average steel ratio in lieu of direct

representation of distributed + boundary steel, given that the difference in flexural capacity is expected

to be small. This should be confirmed once actual wall rebar configurations are determined.

08/14/18 Calculations supporting this procedure will be submitted with the linear design phase in order to show the effect for this specific

building design. Typically the core sees an overall difference in moment capacity of 2-3% which is within the bounds of uncertainty

of the design process.

09/14/18 Pending receipt of later study 12/21/18 A study has been performed to compare the capacity of the core with every longitudinal bar modeled against a model with an

average area of reinforcing evenly distributed along the length of the wall. The resulting difference in capacity is less than 1% and is

considered to be within the bounds of reasonable design uncertainty.

Comment

6.pdf 01/31/19

Comment

Resolved

7 Basis of Design

(07/16/18)

Level 1 Diaphragm

Pg. 17

07/30/18 It is noted that there are multiple steps and changes in elevation at the 1st floor transfer slab. Please

plan on clarifying how these steps are designed and detailed, particularly with slab shear and collector

load path in mind.

08/14/18 From the Basis of Design, page 17:

"The Level 1 diaphragm has several steps and changes in elevation. These steps and changes in elevation are modeled directly in

Perform3D using semi-rigid area element.

For design, the total force transferred out at level 1 and below grade parking levels will be recorded from the NLRHA in Perform 3D

using diaphragm section cuts. The diaphragm and collectors will be designed using partial-depth collectors ad described by MEHRP

Seismic Design Technical Brief No. 3 where possible, although it is likely that collectors at level 1 will be needed across the full-

depth of the diaphragm."

Additionally, the shear from the core will generally transfer from the collector elements to the top of the nearest basement wall in

shear. The level 1 diaphragm will be designed for this shear and any changes in elevation will be designed to resist the sliding shear

across joints through shear friction. The collectors and diaphragm will be proportioned to provide a load path that does not

depend on "rolling" of perpendicular beams where possible.

09/14/18

Comment

Resolved

8 Basis of Design

(07/16/18)

Sloping Columns

Pg. 17

07/30/18 It is agreed that a rigid diaphragm is appropriate in the NL model for typical non-transfer diaphragms,

including where column slopes change - however, it is highly recommended to utilize strut and tie

procedures to evaluate/design a clear thrust load path to the core wall.

08/14/18 We plan to design a single slab/beam section with additional reinforcing to carry the horizontal resultant of the column loading and

do not plan on using the distributed slab reinforcing to resist this load.

09/14/18 Clarification would be useful if it is possible to provide sample sketches of the layout of sloping columns

that are expected. If the component that resists the column thrust is not completely parallel to the

demand, strut and tie evaluation of the slab components that resist the thrust will be necessary to

resolve the various components of the force.

12/21/18 Column steps are generally minor and are shown on the plan sheet included in the comment folder. Ties will be added in line with

resulting horizontal tension forces due to column steps to distribute the load to the diaphragm. The diaphragm will be designed to

transfer the horizontal resultant of the load to the core in shear, rather than using the strut and tie method, in order to keep the

slab reinforcing orthogonal to the main building axes for constructability. The resulting diaphragm moments will be considered in

the design.

To clarify the 8/14/18 response, added reinforcing will be used to resist the tension forces in-line with the force; distributed slab

reinforcing will be used to resist shear and the resulting moment in the diaphragm as the force is transferred to the core wall. We

expect that added distributed slab reinforcing above quantities required for local slab forces will be required to resist this load.

01/31/19 General approach is understood and appears acceptable - pending verification of actual calculations to

be provided at later stages of design.

See supplemental calculations for design of these elements (also included in calculation package on page D4).

05/10/19 See comments #46 and beyond below 06/04/19 See response to comments 46 and beyond below.06/18/19

Comment

Resolved

2 of 20

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Slab-Beam Modeling

Pg. 18

07/30/18 In general, the proposed approach to the strength of the slab-beam hinges appears acceptable.

However, note that if slab rotations end up large and near acceptance criteria, this may require a closer

review/refinement to the modeling assumptions.

08/14/18 Understood. Slab-beam hinges and the rotations resulting from the NLRHA will be included in the MCE submittal.

09/14/18

Comment

Resolved

10 Basis of Design

(07/16/18)

CMS Specta

Pg. 18

07/30/18 Regarding the use of CMS spectra input, please provide a general description of the strategy that will be

undertaken in determining the target periods (or range of target periods) that will be provided to the

geotechnical engineer. Include strategy of how it is intended to:

1) Capture both directions of building motion at both the long-period and short period targets.

2) Represent at least 90% participating mass (per PEER Section 3.3)

3) Represent long-period elongation.

4) Given the intention of determining the target periods from the DE stiffness assumptions (per the

07/16 meeting minutes), please clarify the strategy for the confirmation of these target periods during

the MCE analysis.

08/14/18 The long-period CMS will be developed based on the 1st translational modes in the X- and Y- directions and the anticipated

amount of period elongation from DE analysis. The elongated period is estimated to be about 1.2T, where T is determined from DE

stiffness parameters. The long-period CMS will envelope the shorter of the 1st translational modes and the period coinciding with

the average of the longest 1st translational mode and 120% of that mode.

The short-period CMS will envelop the longest 2nd translational mode and the period at which 90% mass participation is achieved

in both directions.

Inelastic periods will be confirmed during the MCE phase of the design process by measuring the building period from the

displacement history of the roof during periods of major ground shaking. A recent example of this calculation for a similar building

is included in the comment folder and shows that the increase in period due to inelastic behavior is 1.2 x DE periods maximum,

with average period values being similar to the linear model with DE stiffness assumptions.

#10

09/14/18 1) Please clarify why the max end of the long-period range is noted as the average of the longest 1st

translation mode and 120% of that mode, instead of just 120% of the longest translational mode.

2) Closed

3 & 4) While 1.2T may be accurate for this project, this will need to be confirmed as noted. Please note

that if inelastic period is determined through measurement of the roof displacement, ensure that

measurement does not start until after peak response of the motion.

Also note that in the example provided, a few of the reported inelastic periods from different records

appear to calc out as shorter than the elastic periods.

It is recommended that if measurement of the roof displacement is used to determine elongated

period, the elongated period is double-checked either by determining the secant stiffness of a

nonlinear static pushover of the tower, or by creation of a response spectrum from a roof node's

acceleration.

12/21/18 1) ) Based discussions with KPFF and our experience during a similar recent peer reviewed and ongoing high-rise building projects,

we propose to modify our approach for determining conditioning periods for the CMS.

The goal when developing the CMS is to capture a reasonable level of shaking across the periods that are expected to be important

to the building, while also maintaining the benefits of the conditional mean spectrum approach. The benefits are developing target

response spectra that are more realistic and less conservative compared to the UHS. It is our opinion that the approach of

developing CMS which are envelopes of individual CMS conditioned at 1.2T1 and T2 for the long-period CMS (where T1 and T2

equal the larger and smaller of the first translational mode periods, respectively), and T3 and T4 for the short-period CMS (where

T3 equals the larger of the second mode translational periods and T4 equals the period where 90 percent mass participation is

achieved in both horizontal directions), is overly conservative. We propose to develop short-period and long-period CMS with a

narrower range of conditioning periods using the following approach:

The conditioning periods for the long-period CMS were selected so that the long-period CMS developed based on the narrower

conditioning period range does not fall below 90% of the KSSI-adjusted MCER response spectrum between 1.2T1 (6.40 sec) and T2

(3.14 sec) (i.e., the more conservative conditioning period range).

Similarly, the conditioning periods for the short-period CMS were selected so that that the short-period CMS does not fall below

90% of the KSSI-adjusted MCER response spectrum between T3 (0.97 sec) and T4 (0.34 sec).

Based on these criteria, the long-period CMS was defined as the envelop of individual CMS conditioned at 3.8 seconds (1.2×T2) and

4.8 seconds (0.9×T1). The short-period CMS was defined as the envelop of individual CMS conditioned at 0.4 seconds (1.2×T4) and

0.8 seconds (0.8×T3). The 90% threshold (i.e., 10% mismatch between the CMS and UHS) is consistent with recommendations by

Carlton and Abrahamson (2014). It is also consistent with the accepted level of mismatch for amplitude scaled ground motions for

NLRHA in ASCE 7-16. Our forthcoming CMS memorandum will provide further details regarding the development of CMS for this

project.

3 & 4)

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)01/31/19 1) Please justify what aspects of the original method were deemed conservative - is the concern that

the area of the CMS near the target spectrums is over-represented, or the area outside of the target

periods for each spectrum? Although it is understood that the intent is to match within a 10% range

between the CMS and UHS, there is some concern that the the demands at T1 and T3 may be under-

represented, given the conditioning periods are set lower than the fundmantal periods (not including

elongation) at 0.9T1 and 0.8T3.

3 & 4) Response appears to be missing

5) As a related note (regarding KSSI) - ASCE 7-16 Chapter 19 notes that while SSI representation is

acceptable, the analytical models should directly incorporate horizontal, vertical, and rotational

foundation and soil flexibility. It is our understanding, per the BOD, that soil springs will not be

considered in the analytical models - please clarify.

02/15/19 1) Our concern is that the original approach (i.e., a short-period CMS that equals the MCER UHS from the period at which 90% mass

participation is achieved in both horizontal directions to the longest 2nd translational mode period, and a long-period CMS that

equals the MCER UHS from the shortest 1st translational mode period to an lengthened longest 1st mode period) imposes MCER

level demands at practically all the periods that contribute to the building response. The period range summarized above comprises

92% to 94% of the total mass in both horizontal directions when considering both CMS.

Therefore, we sought to modify the approach described above based on experience on past performance-based design projects

and available literature. Our modification results in a modestly narrower range of conditioning periods to be enveloped and

maintains a maximum mismatch of 10% as recommended by Carlton and Abrahamson (2014) for developing CMS that capture a

range of periods.

This modification results in ground motions that are about 4% lower than the MCER UHS at T1 and T3. In our opinion, this is a

reasonable trade-off between the conservatism built into the UHS approach and the benefits of adopting the CMS approach, which

are more realistic spectral shapes and demands.

3) This question will be addressed at the time of the nonlinear submittal

4) This question will be addressed at the time of the nonlinear submittal

5) The statement in ASCE 7-16 Chapter 19 identified by the peer reviewer appears to be overly generalized. Based on the ASCE 7-16

Chapter 19 commentary, there are three specific types of soil-structure interaction (SSI) that are accounted for using three separate

approaches. These are: 1) foundation deformations, 2) inertial SSI effects, and 3) kinematic SSI effects (KSSI).

The requirement that the analytical models should directly incorporate horizontal, vertical, and rotational foundation and soil

flexibility is specific to foundation deformation SSI (refer to pg. 703 paragraph 6 in the ASCE 7-16 commentary) and is not a

requirement for considering inertial SSI or KSSI. For the 2200 Block 4th Avenue project, we only consider KSSI and not foundation

deformation SSI or inertial SSI; therefore, the elements required to account for these phenomena (i.e., horizontal, vertical, and

rotation foundation and soil flexibility for foundation deformation SSI, and radiation damping for inertial SSI) are not included in

the structural model.

03/05/19 1) Closed

3) Pending nonlinear submittal

4) Pending nonlinear submittal

5) Closed

xx 3&4) Displacement records for the roof nodes are included in the MCE calculations for each of the 22 time history analysese of the

building. Calculations of the long period elongations and a summary table of the elongations are provided in the submitted MCE

calculations on page B62. In general, the measured elongated building period is less than the elastic period using DE stiffness

assumptions. While this result is surprising, it means that the conditioning periods selected for the acceleration records are

acceptable.

Per an e-mail from Nabih and Daniel on 4/2/19, it is response spectra created from a roof nodes acceleration are not required

because the building displacement is "clean" and is clear enough to evaluate the approximate elongated periods. 05/10/19

Comment

Resolved

11 Basis of Design

(07/16/18)

Ground Motions

Pg. 18

07/30/18 The "Analysis Requirements" section notes that the NRHA will be performed with 2 suites of 11 ground

motions each.

Please clarify if the PEER requirement (Section 3.3) of having a minimum of 5 records per each seismic

source contributing more than 20% to the hazard at a period of interest will be maintained.

08/14/18 The PEER TBI 2.03 criterion will be satisfied such that a minimum of 5 records will be used to represent discretely modeled sources

that contribute to greater than 20% to the total hazard.

Note that in the structural BOD document dated 7/18 that was submitted for review, Section 3.3 specifies 11-15 pairs of ground

motion acceleration histories.09/14/18

Comment

Resolved

12 Basis of Design

(07/16/18)

Backstay Analysis

Pg. 19

07/30/18 It appears that for the above-grade components, the design will be based upon the "expected stiffness"

which per Table 7-3 is equivalent to the "Lower Bound Stiffness". While the intent to capture the

bounded demands at the 1st floor and below is understood as the focus of the bounded analysis,

please justify evaluating the remainder of the structure only at the "expected stiffness", and justify

those values.

08/14/18 This comment appears to relate to the Basis of Design - Early draft dated July 16th which is not intended to be a formal submittal.

The version of the Basis of Design that was submitted for formal review is dated July 18th, after incorporating the results of the

discussion during the kickoff meeting. The July 18th version has the discussion of bounding the backstay stiffnesses removed after

we confirmed that soil springs would not need to be modeled for the building during the kickoff meeting. In KPFF's experience, if

the soil springs are not modeled then the differences that arise from the bounding analysis based only on diaphragm and basement

wall properties are relatively small. Please let us know if you think that a bounding analysis will be worthwhile even if soil springs

are not included in the building model.

The last commentary paragraph to 4.2.4 of the PEER TBI Guidelines states, "it might be necessary to make bounding assumptions

on stiffness properties to bracket the force for which the various components of the podium structure should be designed."

Because the PEER TBI Guidelines do not recommend using the bounding assumptions for the design of the superstructure

elements, expanding the bounding analysis to the superstructure was not included in the design procedure.

Expected stiffnesses for MCE properties are based on PEER TBI Guidelines Table 4-3. Section 4.6.3 recommends the use of this

table in lieu of detailed justification of the stiffness values.

09/14/18 Although the impact is reduced without soil springs, similar projects have still seen non-negligible

differences for the foundation/diaphragm/basement wall design based upon the bounding analysis -

please include.

It is agreed that expanding the bounding analysis to the superstructure is likely unnecessary - it would

be best if this could be justified by a comparison of global results (shear, moment, drift) for one ground

motion with stiff vs. soft bounding properties to confirm that the non-negligible differences are only

seen at grade and below.

12/21/18 The bounding analysis language has been added back into Section 7.7. This section has been expanded to describe the process of

verifying that the bounding analysis is not applicable to the superstructure elements.

01/31/19

Comment

Resolved

4 of 20

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Foundation Design

Pg. 21

07/30/18 1) It is noted that "Foundations are designed in SAFE using the Design Point moments for the core and

basement walls as defined above" (referring to Figure 7-2. Please clarify how the referenced approach

is intended to be implemented for the basement walls.

2) Please confirm that in addition to the core and basement wall demands, column demands will also

concurrently be applied to the SAFE model for foundation design.

3) Please clarify how foundation demands will be imposed upon the SAFE model (point loads, line

moments, distributed line loads, etc...). If not currently available, please plan on providing in advance

of the submission of the foundation design.

08/14/18 1) The foundation design envelope shown on Figure 7-2 will result in a realistic assessment of the directional combination that

should be used for the design of the foundation. Typically the result shows that the directional combination of moments at the

foundation should be 100% + 50-60% in the orthogonal direction. Average maximum moments for the core wall, and for the

basement wall will be calculated from the NLRHA analysis using section cuts in the Perform 3D model. The average maximum

moments from the core and basement walls are applied concurrently in load cases that are then included in load combinations

used for design. The forces in design load combinations are divided by a B factor of 1.05 per BOD Table 7-1 to account for this

factor when using the automated flexural design process within SAFE.

2) Column demands including axial loads resulting from the beam-frame behavior of the NLRHA will be applied concurrently to the

SAFE model. Seismic axial loads from the NLRHA will be applied to that they apply an overall moment to the foundation in the

same direction as the core wall and basement wall moments.

3) The moments from the core and basement walls are applied to the SAFE model through line loading on wall elements that are

included in the model. Walls are divided into different line segments and the overall core moment is deconstructed into axial loads

and in-plane moments for each segment. Axial loads are applied to SAFE line elements as a uniform line load; in-plane moments

are applied using non-uniform line loads that result in the same in-plane moment. Axial loads from the columns are applied as

point loads at the column CG location.

09/14/18

Comment

Resolved - to be

further verified

and confirmed

at later stages

of design when

analysis models

are provided

14 Basis of Design

(07/16/18)

Acceptance Criteria

Pg. 21

07/30/18 #4 on the acceptance criteria list notes that "elements not explicitly modeled do not exceed the

deformation limits at which members are no longer able to carry their gravity loads". Is this acceptance

criteria expected to be utilized for any components on this project?

08/14/18 At this time we do not plan on checking individual components against this criteria and we plan to rely on the interstory drift limits

to qualify the performance of the elements intended to mainly resist gravity loading. This language is included in the BOD to be

consistent with the PEER TBI guidelines and to serve as a starting point should any additional elements need to be evaluated.09/14/18

Comment

Resolved

15 Basis of Design

(07/16/18)

Structural Element

Classification Table 7-4

Pg. 22

07/30/18 Regarding Table 7-4: Please include "Force transfer from diaphragms to vertical elements of the

seismic force resisting system, including shear-friction between diaphragms and vertical elements

(either as it's own separate line, or in addition to the "Collectors" line item), per PEER Appendix E.

08/14/18 In Table 7-3, the entry in the component column "Collectors" has been changed to read "Diaphragm connection to concrete walls".

The entry in the Seismic Action column "Axial" has been changed to read "Axial and Shear".

All collector forces at elevated and transfer diaphragms will be designed as force controlled, critical elements regardless of the load

path used to transfer the forces from the diaphragms to the vertical SLFRS.09/14/18

Comment

Resolved

16 Basis of Design

(07/16/18)

Table 7-4

Pg. 22

07/30/18 For efficiency/clarity of review as the design progresses, it is recommended to include the associated φ

and "B" factors that will be utilized for the various designed components - either as an additional

column in Table 7-4, or it's own separate table.

08/14/18 φ and "B" factors have been included in Table 7-3

09/14/18 Although values provided in table 7-3 appear reasonable, please plan on providing the supporting

calculations that were used to determine the various values in later stages of the review process.

In particular, concrete shear wall "B" will of course need to be verified per the strains in the wall (also

see comment #17).

Please also clarify how the bias factor is intended to be utilized for Soil Bearing.

12/21/18 Calculations for Bias factor will be provided at the time of the MCE calculations. Values for some elements that have strength

contributions from both steel and concrete are dependant on the proportion of the demand resisted by each.

Agreed. Wall strains will be verified at the time of the MCE submittal.

The bias factor for soil bearing will be used to evaluate the soil bearing pressure calculated by the SAFE model for MCE loading

conditions and will increase the acceptable soil bearing values beyond the nominal soil bearing strength for the MCE condition. If P

is the soil bearing pressure, the bias factor for soil bearing is calculated as Pexpected = F.S. x Pallowable = 2.0 X Pallowable. Pnominal = 1.4 x

Pallowable. B = 0.9 x Pexpected / Pnominal = 0.9 x (2.0 Pallowable) / (1.4 x Pallowable) = 0.9 x 2.0 / 1.4 = 1.28.

01/31/19

Comment

Resolved -

Pending MCE

calcs for "B"

justifications

17 Basis of Design

(07/16/18)

Wall Strain

Pg. 23

07/30/18 In reference to the strain-based core wall acceptance criteria, please clarify over what height strains will

be recorded. Will a refined height in the expected plastic hinge zone be utilized?

08/14/18 Basis of design has been modified at this location to specify that the longitudinal wall strains are measured over a single story

height.

09/14/18 It is generally agreed that measurement of strains over a single story height will be sufficient for a

majority of the model.

However, in order to ensure that strain localizations are not being missed within the potential hinge

zone, please include a parametric study that divides the vertical height of the core wall elements (and

strain gages) within the lower stories into ~2 to 3 subdivisions (depending on story height). Comparison

could be made perhaps with just one ground reprentative ground motion to understand if any

localization of nonlinear strain is occurring.

Although it is assumed that tensile strains will still not be close to the typical max strain acceptance

criteria, the study could provide insight into whether the 1.5x amplification to wall shear capacity is

appropriate in these critical zones.

04/05/19 Results from using a more refined area element mesh in a pier immediately above level 1 show that modeling the piers with a

single area element between floors is accurate and that strain localizations due to double curvature do not occur. Strain

localizations observed are attributed to limitations of the fiber modeling process and are not representative of the actual behavior

of the building. See supporting calculations in the comment folder.

05/10/19 The concern regarding the refined strain measurements was not necessarily to understand if double-

curvature occurs, but rather to understand what strains are reported near the base of the wall when a

more refined approach to the strain measurements are observed (across the entirety of the core, not

just an individual pier).

The main intent of this study is to confirm how sensitive the strains are to the refined meshing in the

context of the use of B=1.55 for shear, which is dependent upon tensile strains < 0.01 (in this case, the

bottom of the wall reports strains > 0.02).

While it is understood that there is no prescribed height over which to measure the strain, the refined

approach helps provide a better "index" for the strains at the base of the wall to help confirm that

B=1.55 at the base of the wall is or isn't appropriate.

06/11/19 To better understand the behavior a similar study was performed with the critical wall meshed into 9 elements in the first story,

each 14" tall. As before, the results show a concentration of strain in the lowest element and lower strains in the 8 elements above.

The average of all 9 strain gages is 9%-12% higher than the value observed when a single element is used for the full story height.

Given that the average of the maximum tension strains is 0.0042 for the critical location, the average strain at this level is still well

below 0.01 even if a 12% increase is applied. Thus, the use of equation 4-2 from the TBI Guidelines is appropriate for determining

the bias factor of shear in the core walls.

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)06/18/19 As a clarifying follow-up, please refer to email from 06/07:

"The intent of comment #17 was not necessarily to confirm whether or not strain localizations due to

double curvature occur, but rather how sensitive to the meshing of the walls the nonlinear strains are

in context of justifying the bias value “B” applied to the shear strength of the core walls, which is

dependent upon this strain.

Therefore, as a response to Comment #17, please provide a justification for the value “B” used in the

core shear wall shear calcs, given the study that had been undertaken as part of the 04/05/19

response."

xx

07/29/19

Comment

Resolved per

Comment #17

Exhibit

18 Basis of Design

(07/16/18)

Wall Confinement

Pg. 24

07/30/18 Please clarify the last bullet point on this page. Does this imply that a fully-confined section of core wall

will rise from foundation to a certain level above grade? Is there an approximate idea of that Level

(currently noted as "Level X" in the BOD)?

08/14/18 The average of all 9 strain gages is 9%-12% higher than the value observed when a single element is used for the full story height.

Given that the average of the maximum tension strains is 0.0042 for the critical location, the average strain at this level is still well

below 0.01 even if a 12% increase is applied. Thus, the use of equation 4-2 from the TBI Guidelines is appropriate for determining

the bias factor of shear in the core walls.09/14/18 Please also confirm that ACI Section 18.10.6.2 checked at the DBE level analysis, and that 18.10.6.2(b)

will be checked against the final configuration of boundary zone height extension.

12/21/18 A code exception has been added to the BOD for ACI Section 18.10.6.2. Per the Basis of Design Section 7.7, confinement of the wall

will be provided where the average wall strains in compression exceeds 0.002. This criteria provides a factor of safety of 1.5 against

crushing of the concrete and is consistent with the factor of safety for other non-ductile mechanisms in the building. Note that

measured compression strains from the MCE model are multiplied by a factor of 2 to account for the underprediction of

compression strains per the basis of design.

The code exception states, "• Height of confinement reinforcing per ACI 318-14 18.10.6.2 within the core shear walls. Instead of

complying with the prescriptive requirements of ACI 318-14 18.10.6.2, the compression strains in the core wall are measured

directly from the nonlinear analysis of the building and confinement reinforcing is provided where strains from the analysis show

that it is needed."01/31/19 Proposed exception is preliminarily acceptable - however it would be useful to review the actual MCE

strains that are to be reported from the future analysis to understand how these requirements would

compare to the prescriptive 18.10.6.2 requirements prior to confirming.

02/15/19 See comment folder for calculations of required confinement length and height per ACI 318-14 Section 18.10.6.2 for combined core

and individual wall piers. These values will be compared with the strains measured during nonlinear analysis of the building.Comment

18

03/05/19 Comment to remain open through comparison with strains measured during the nonlinear analysis for

confirmation

04/05/19 See pages B16-B22 of the MCE calculation package for strain results from the nonlinear analysis.

05/10/19

Comment

Resolved - see

MCE notes

regarding strain

comments

19 Basis of Design

(07/16/18)

Wall Design

Pg. 25

07/30/18 For flexural design of the wall, it is noted that "The core wall is evaluated for flexural and axial loads for

the entire rectangular cross section, and for individual piers that are bounded by coupling beams on

both sides". Please confirm that all individual piers, not only those bounded by coupling beams, will be

individually evaluated (for example, the short-direction N-S piers at either end of the core).

Please also confirm that these individual piers will be evaluated utilizing effective wall flange widths per

ACI 318.

08/14/18 Yes, this pier will be evaluated and designed for flexural and axial loads. The C-shaped pier is bounded by the coupling beams at

the elevator lobby at one end any by the door to the stairwell at the other end.

The following language has been added to the Wall Design section on page 25. "Where the effective flange widths requirements

per ACI 318-14 18.10.5.2 break the wall into sections that are not bounded by coupling beams on both sides, the section including

the effective flange width will be evaluated for uniaxial moments and axial loads."

09/14/18

Comment

Resolved

20 Basis of Design

(07/16/18)

MCE Diaphragm Design

Pg. 25

07/30/18 It is noted that "elevated non-transfer diaphragms are designed for forces from ASCE 7-10 Section

12.10".

Ensure that diaphragm design is also evaluated at the average diaphragm demands as observed from

the MCE time history analysis.

08/14/18 Language in the Basis of Design document has been changed to read, "Elevated non-transfer diaphragms are designed for in-plane

shear and flexural forces obtained from the MCE NLRHA analysis using section cuts from the Perform3D model."

09/14/18

Comment

Resolved

21 Basis of Design

(07/16/18)

Appendix B 07/30/18 As noted in comment #3, Appendix B is interpreted as a set of example nonlinear modeling input that

will be further developed with more specificity to this project and it's components later. A few related

comments:

1) On Pg. B-7, it is noted that different shear modifiers for the wall may be used within and outside of

the plastic hinge, referring to Table 7-1. Table 7-1 only notes one shear modifier for core wall shear

properties - please clarify.

2) On Pg. B-1, it is noted that certain modifiers were utilized on the backbone in order to match the

coupling beam test data - please provide the associated data/plots that were utilized to determine

these modifiers.

08/14/18 1) The sentence stating that different shear stiffness multipliers are used has been removed. Only the single multiplier of 0.5 per

Table 7-1 is used.

2) The nonlinear coupling beam model has been calibrated to the test data of Naish, Wallace (2009). Hysteretic loops from the

Perform3D model definition plotted on top of the Naish and Wallace test data is included in the comment folder for review.

#21 09/14/18

Comment

Resolved -

more specific

calibration info

to be reviewed

later per

Comment #3

22 Basis of Design

(08/14/18)

Section 7.6 09/14/18 Last sentence of first paragraph of Section 7.6 references a shared podium and church in Section 7.8 -

please clarify.

12/21/18 This sentence has been deleted as it was a carryover from an earlier project.

01/31/19

Comment

Resolved

23 Linear Design Submital

(12/21/19)

General Calcs 01/31/19 The current submittal provides DBE-related calculations for the design of the flexural strength of the

core wall and the coupling beams (along with initial DBE-amplified confirmation of the wall

thicknesses).

As noted in Comment #2, the DBE analysis intended to provide justification that an SLE analysis was not

required. Will any other aspects of the lateral system (foundation, diaphragm, basement walls, etc...)

be evaluated as part of the linear design? Although it is understood that most if not all force-controlled

components will be controlled by the MCE analysis, this would need to be justified for all components.

Note also that the DBE evaluation is a Seattle SDC requirement.

02/15/19 Force controlled elements will be checked for DE level forces at the time of the nonlinear submittal at which time the force

controlled elements will be designed for both the results of the NLRHA and the DE level forces.

Sequencing the design in this way allows the design process to be more efficient, and avoids the problem of adding a design to the

drawings only to increase the quantity of steel or size of the element at a later date.

Submittal: Linear Design Submittal (12/21/19; 02/15/19)

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)03/05/19 Proposed approach is acceptable given that all aspects of the lateral system design (including

foundation and diaphragms) are at some point provided for the DBE analysis (or that a clear

comparison is provided that shows that DBE design will not govern over MCE design)

04/05/19 DE level forces have been included in the nonlinear calculation package for comparison with MCE results. In general, the MCE was

found to control the design. For example, see pages B184 for gravity columns, B216 for core wall shear comparison and B233 for

above grade diaphragms. Similar calculations are provided in section C for the level 1 diaphragm, basement walls and below grade

diaphragms.05/10/19

Comment

Resolved


(12/21/19)

Seismic Loads

Pg. A2

01/31/19 It is noted that gravity analysis and design is still ongoing. As part of this, is there the potential for

typical slab thickness to be modified? Or is this aspect of the design relatively set (excluding perhaps

local heavily loaded areas or transfer diaphragms)?

02/15/19 Typical slab thicknesses will remain at 7-1/2". A model of the typical slabs has been created and the 7-1/2" slab has been found to

be adequate for the structural design. Modifications to the slab will be performed with the goal of keeping the 7-1/2" thickness

and modifying the reinforcing only.03/05/19

Comment

Resolved


(12/21/19)

Gravity Loads

Pg. B3

01/31/19 Similar to Comment #24, this page notes that the gravity line loads applied in the model will be updated

after the gravity analysis is complete. Please ensure that the final versions of the models (and any

pertinent design aspects that are impoacted) are sent after these gravity demands are updated.

02/05/19 Gravity loads in the ETABS model have been updated based on the latest load maps. Changes to loading are typically small (less

than 5%) and do not change the design.

03/05/19

Comment

Resolved


(12/21/19)

Torsion Irregularity

Pg. B15

01/31/19 1) It is noted that the max/average ratio at each level was less than 1.2 - however the following pages

show ratios up to ~1.4 - was this a typo?

2) It is noted that torsion was evaluated from displacements extracted from the response spectrum

load cases. Given absolute results from modal superposition, these displacements should either be

taken by manual combination of modes

(https://wiki.csiamerica.com/pages/viewpage.action?pageId=4554971), or a static representation of

loading

02/15/19 1) Irregularity has been re-evaluated using static load cases including accidental torsion. Max/average ratios have been found to be

generally less than 1.2, with the exception of levels 2 and 3 where the maximum ratio is 1.22. The sheet B15 has been revised and

included in the comment folder.

2) Torsional displacements have been evaluated using static loading. Updated results are provided in the comment folder.

Comment

26 03/05/19

Comment

Resolved


(12/21/19)


Pg. B16

01/31/19 Please clarify the process of utilizing the link elements in the model to evaluated torsion. How does this

work for some of the more irregular-shaped floors (like L7 or L2)?

Similarly, please clarify the RX/RY restraints that are assigned the slab at these link nodes.

02/15/19 Link elements are used to measure drifts at the corners of the building because the link output can be arranged to output interstory

drift values while including all modal effects. Link output is used instead of subtracting story displacements between floors

because subtracting the story displacements can underestimate the interstory drifts due to higher mode effects. Use of ETABS

output tables report similar results but do not report the values at each building corner which is needed for torsional checks of the

building. Per the response to comment 26, torsion calculations have been changed to use static ELF loads so either method will

yield similar results.

RX and RY restraints are applied to the nodes at the top and bottom of the link so that the link remains vertical at each end, forcing

the link into double-curvature-like deformation. Because the link is held to the vertical at the ends and no rotation of the element

occurs, all X and Y displacement between the ends of the link is recorded as member displacements within the link element.

Because no horizontal members with significant stiffness are connected to the restrained nodes, the restraints do not affect the

lateral stiffness of the building.

At irregularly shaped floors, link elements are carried vertically downward from the floor above and connected to the diaphragm

below at the same X and Y ordinates that they use at the floor above. This positions the link at the corner of the skin of the

building, which is the most eccentric location where reported interstory drifts would be meaningful. At some locations a node will

be connected to a nearby diaphragm using a rigid element because no diaphragm exists immediately above or below. At these

locations, the node has been connected in a way that we think would best estimate the torsion at this location and is a modeling

idealization used for practicality.

03/05/19


(12/21/19)


Pg. B20

01/31/19 While it is noted that the average Ax value is 1.02, individual floors (near the base of the building) see

Ax up to about 1.13, which would represent an eccentricity of around 5.6%. Please clarify why Ax is not

provided on a per-floor basis utilizing these values.

02/15/19 With the revised torsion calculations provided in response to comment 26, Ax = 1.0 for all levels and no amplification is required.

03/05/19

Comment

Resolved


(12/21/19)

Boundary Zones

Pg. C9

01/31/19 It appears that the end of the last paragraph may be missing text - please clarify. 02/15/19 The full text should read as follows (previously missing text in bold):

For WG3 and WG4, the longitudinal steel at levels P3-L6 has been grouped to match the boundary zones calculated for Level 6.

Since the plastic hinge zone will have heavy transverse reinforcing carried through to the foundation, the entrie wall section is

effectively a boundary zone at these levels. Consequently, the longitudinal steel has been located over a reduced length for

economy of design, and need not match the calculated boundary zone length for level 1.

03/05/19 Please clarify response. Response refers to the entire wall section as effectively a boundary zone at

these levels, however the current drawings do show distinct boundary zone regions at the ends of

these walls. Please also clarify what is meant by locating the longitudinal steel over a reduced length.

On a similar topic:

1. Please clarify how the boundary length at the corners of the walls were determined.

2. To prevent buckling of the longitudinal bars, it is typically recommended to provide supporting cross-

ties at all longitudinal bars in the cross section for at least the plastic hinge zone - please provide

justification if this will not be provided.

04/05/19 1) The SBZ length is determined as described and calculated in the linear submittal calculations pages C9 through C25 and in

appendix A. For each wall section, the neutral axis depth "c" is calculated in Spcolumn at the nominal moment capacity of the wall

with the axial load as determined by gravity and DE seismic load case combinations. The neutral axis depth "c" is used in the

equations given in ACI 318-14 Section 18.10.6.4 to determine the length at the wall corners and ends that will be confined with

hoops and ties.

2) Supporting cross ties are provided at all longitudinal bars at level 6 and below. See updated core wall sections on sheets S0310

through S0312.

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)05/10/19 1) In terms of the zones at the corners, was "c" rotated to the angle of the applied demands when

determining the required boundary zone lengths?

2) Confirmed. As a supplemental note, particularly for the floors that do not have the supporting cross

ties at all bars, it would be recommended for the vertical bars to be placed on the inside face of the

horizontal reinforcement, rather than completely unsupported at the exterior face of the horizontal

bars.

06/04/19 1) The distance "c" is determined based on the neutral axis at the nominal moment capacity as calculated about the primary axes

of the core per the requirements of ACI Secion 18.10.6.2.a. Full confinement is provided in the confined levels of the wall so the

extent of the corner rebar "cages" is more for constructability than compliance with ACI confinement extent requirements.

2) The concrete of the wall is expected to provide adequate buckling restraint for the vertical web bars as the concrete is not

expected to spall based on the strain results of the nonlinear analysis and should provide restraint for these bars. Note that the

strains in the building would need to be significantly higher than calculated in the nonlinear analysis for the web concrete to spall,

first due to the 1.5 factor of safety applied to the compression strain limit of 0.002 above which confinement reinforcement is used,

and second due to the fact that the calculated compression strains are measured at the wall corners and edges, not at the wall

webs where the strains should be less.

If the concrete were to spall at the beam webs due to unexpected or larger than expected seismic response, the horizontal bars are

only supported at their ends and are unable to restrain the vertical reinforcing because they are flexible out of the plane of the

wall, and because the spacing of the horizontal bars is too large to provide buckling restraint for the vertical bars even if they were

able to provide restraint.

Placement of the vertical bars on the outside of the horizontal bars is typical practice in Seattle and is done for constructability

reasons. The placement of the vertical reinforcement has not been changed due to the anticipated constructability problems this

would cause with the wall.

05/10/19 1) For bi-axial analysis, spCol provides the neutral axis in terms of a depth and an angle of the neutral

axis - the intent of the comment above is to understand if this angle was correctly applied when

determining required boundary zone length (including at the corners).

2) Although the reviewer has seen other recent PBD reviews in Seattle placing the vertical bars as the

inner layer, given the relatively low tensile strains in the bars, it is agreed that there is likely low chance

of bar buckling along the majority of the height. One note: The ties at every bay are continued up

through L6 presumably to "capture" the intended zone of maximum yielding within the plastic hinge

zone and above, given that ties restrain bar buckling which is initiated through bar yielding. The ties at

every vert. bar stop at L6, however there are actually higher strains in the vertical bars (~yield strain)

from about L14 to L22 - please clarify intent of the tied bars and justify/confirm the current cut-off

point.

07/03/19 1) Per our discussion on the 6/24/19 conference call, from the foundation to level 6 the entire area of the core wall including the

boundary zones at the corners and the entire length of the web is confined as shown on S0310, S0311, and S0312. Calculations

performed to size the boundary zones only considered flexure along the primary building axes, but consideration of the neutral axis

depth cannot increase the area of confined concrete since the entire area of the core is confined.

2) Per our discussion on the 6/24/19 conference call, confining ties extend from the foundation to level 6, which exceeds the

criteria established in the basis of design document based on strains. Per the criteria, confining reinforcing is provided where the

average compression strain exceeds 0.002. Calculation results show that this limit is exceeded only for by the long-period records

at level 1. The confinement reinforcing is not curtailed until level 6 to capture the intended zone of plastic hinging and above and is

a criteria betterment.

For the long period records, compressive and tensile strains do not exceed the values at level 6 higher in the building. For the short

period records, strains higher in the building exceed the values at level 6, but do not exceed the threshold values to provide

confinement reinforcing. Average tensile strains of the long period records are significantly below the yield strain of the

longitudinal steel. Average tensile strains of the short period records just reach the yield strain of the longitudinal steel, indicating

that only minimal yielding of the bars is likely to occur in the upper levels of the building.07/26/19

Comment

Resolved


(12/21/19)

Shear Wall Shear

Pg. C111

01/31/19 Currently only wall thickness is confirmed (through stress comparison to ACI limits). Please clarify if any

horizontal reinforcement design will be included as part of the DBE design.

02/15/19 Per comment 23, DE design of force controlled elements will be submitted along with the nonlinear submittal.

03/05/19 Pending nonlinear submittal 04/05/19 MCE design of the core wall shear and comparison with the DE level forces is provided on pages B198-B224.05/10/19

Comment

Resolved


(12/21/19)

Shear Wall Shear

Pg. C111

01/31/19 Please clarify if the shear demands in the walls are factored loads, including gravity demands (including

impacts of sloping columns)

02/15/19 Estimated shear demands in the wall do not include the gravity load demands because the gravity demands are expected to be

small and within the uncertanty of the seismic wall shear estimates.

03/05/19 Although it is agreed that the typical gravity demands upon the core wall are likely relatively negligible,

this is likely not the case for the horizontal thrust that is formed due to the sloping columns, which must

be resisted by the core wall. Please clarify how these demands are incorporated into the shear design

of the core wall.

04/05/19 The horizontal thrust has been considered directly in the Perform and ETABS models through modeling of the inclined columns.

Horizontal elements have been designed in the slab to transfer the horizontal component of the force (see also response to

comment 8). See supplemental calculations for confirmation of DE wall shears that account for shear due to gravity thrust forces.05/10/19

Comment

Resolved


(12/21/19)

Shear Wall Shear

Pg. C112

01/31/19 Please clarify how the X and Y direction overstrength values of 3.2 and 5 were determined. Although

there are no specific exceptions to these values, it would be helpful to understand if there was a certain

methodology used to come up with the two different values in the two directions. Are these

amplifications intended to represent the required "Vn" that will result from the MCE analysis?

02/15/19 X and Y overstrength values are determined based on previous comparisons of DE wall shear compared to design demand values

obtained through the NLRHA. The smaller X-direction value was chosen because of the two coupling beams along each length of

wall and the large amount of energy dissapation that will occur due to the yielding of the coupling beams. The larger Y-direction

value was chosen because there are no coupling beams along these lengths of walls.

Note that these factors are not used in the final design of the building and are only used for preliminary proportioning.

03/05/19

Comment

Resolved


(12/21/19)

Shear Wall Shear

Pg. C115

01/31/19 As a note: it is understood that the demands shown on these wall shear calculations are amplified

beyond the code-prescribed values and therefore should not have any issue with code-acceptance.

However, piers W3 and W4 do indicate a high shear stress at the base of their walls (below L02),

averaging about 8.4root(f'c) in the Y direction. This is only noted as an area to watch out for, given that

if these values are intended to approximate the required Vn in this direction at the MCE analysis, the

walls in this direction would exceed an 8root(f'c) shear stress.

02/15/19 The areas where the estimated shear demand exceeds 8root(f'c) are noted and will be looked at further during the nonlinear

design phase.

03/05/19

Comment

Resolved


(12/21/19)

Coupling Beam Design

Pg. D3

01/31/19 Pg. D3 reference a phi=0.85, related to diagonally reinforced coupling beams which we understand are

currently not utilized. Is this reference only a typo - or is phi=0.85 utilized anywhere in the design of the

conventionally-reinforced coupling beams?

02/15/19 This is a typo - phi = 0.75 was used for all conventionally reinforced coupling beams.

03/05/19

Comment

Resolved

8 of 20

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)35 Linear Design Submital

(12/21/19)

Coupling Beam Design 01/31/19 The DE Coupling Beam Design info indicates that both B7 and B8 have a L/h ratio of 2.0. This is only

noted in case any potential further geometrical modifications are expected for the coupling beams, as

this ratio is right of the edge of the diag. reinforcement requirement.

02/15/19 Noted. The depth of the coupling beams was deliberately chosen to result in an aspect ratio that would allow conventional

reinforcement for constructability.

03/05/19

Comment

Resolved


(12/21/19)

Coupling Beam Design 01/31/19 Referencing Pg. D10 as an example (for Coupling Beam B8):

1) Please confirm that the demands shown (MDE, Mgravity) are factored loads

2) Please clarify how redistribution was undertaken. If redistribution was utilized, why are the final

values still different from one another?

3) Confirm that redistribution does not modify the demand on any particular coupling beam more than

20%. This appears to be violated for the B8 roof coupling beam, for example.

4) The σDE value appears to be based upon the original, rather than the redistributed shear demands -

please clarify.

02/15/19 1) Confirmed.

2) Redistribution is based on the coupling beam types shown on page D11. The variation in final value is typically due to the gravity

load, which is a small proportion of the total load on the coupling beam. DE forces are often redistributed to the same value but

gravity forces are not redistributed. See supplemental calculation in the comment folder for clarification.

3) Redistribution does not reduce the demand on any particular coupling beam by more than 20%. However, forces have been

conservatively allowed to increase by more than 20%. This increase is typical for the most lightly loaded coupling beams at the roof

and level 1.

4) Please see the comment folder for an updated calculation based on the redistributed shear. Since force redistribution will only

reduce the shear demand on the most highly stressed coupling beams, initial stresses are conservatively checked without the

redistribution. As a comparison, the maximum stress is 6.1*root(f'c) with the original values but drops to 5.1*root(f'c) when the

redistributed shear is used. The average across all coupling beams is essentially the same in either case (4.1*root(f'c)).

03/05/19

Comment

Resolved


(12/21/19)

Coupling Beam Design 01/31/19 Referencing Pg. D10 as an example (for Coupling Beam B8):

Please clarify what the "B" value represents in the φBVs value. This appears to represent B=1.05.

02/15/19 "B" represents the bias factor and is used when checking the capacity based shear design of the coupling beams. B = 1.05 in

accordance with table 7-3 in the design criteria.

03/05/19 These calculations are understood to represent the DBE "Code-based" calculations. While the bias

factor "B" is utilized as part of the MCE force-controlled acceptance criteria as noted in the BOD Section

7.7, it does not appear to be appropriate to be utilized to increase the shear strength of the capacity-

based shear code check. Or, please clarify if this is being utilized in a different fashion.

04/05/19 The DE coupling beam designs have been re-verified with the B factor removed from the calculations. See supplemental

calculations.

04/30/19

Comment

Resolved


(12/21/19)

Coupling Beam Design

Pg. D30

01/31/19 It is agreed that DCR's slightly more than 1.0 are acceptable at DBE, given explicit representation of the

flexural yielding at MCE. However, confirmation of how redistribution was utilized (see Comment #36)

is required to ensure that these values are accurate.

02/15/19 Please see response to comment #36 and supplemental calculations in the comment folder for clarification of force redistribution.

03/05/19

Comment

Resolved


(12/21/19)

Applied Loads

Pg. G4

01/31/19 A typical 25psf SDL is noted in the building mass takeoffs. For a few local floors (such as L2 and L6), the

load maps in the structural drawings indicate a larger,varying SDL is required - please clarify.

02/15/19 Mass has been updated in ETABS for all levels to reflect the latest load maps. The updated mass does not significantly affect the

design loading of the structural elements.

03/05/19

Comment

Resolved


(12/21/19)

Structural Drawings 01/31/19 1) Sheet S1.05: The western edge of the core wall appears to have openings in plan near the corners - is

this just a drafting issue?

2) Sheet S1.09: Please discuss the design procedure for the thin "bridge" portion of slab at the eastern

side of the core. Will this portion be treated similar to the rest of this slab?

3) Is the plan on Sheet S1.12 accidently mislabed L30 instead of L31?

02/15/19 1) Plan openings are a drafting issue. The wall is continuous at the locations where openings were shown. An updated second

floor framing plan is included in the comment folder.

2) The design of the bridge at level 6 is uncertain. For the moment, we plan to include the bridge in the ETABS and Perform models

as a continuous force-controlled portion of the floor plate. If the Perform analysis results in forces that are undesignable, we will

introduce a seismic joint on one side of the bridge.

3) Correct. An updated level 31 framing plan is included in the comment folder.03/05/19

Comment

Resolved


(12/21/19)

ETABS Model - L31 01/31/19 The load map for L31 on S0.14 indicates a varying SDL. In the model, a single additional mass value is

applied - please update or clarify.

02/15/19 Massing in ETABS has been updated to reflect the varying SDL.

03/05/19

Comment

Resolved


(12/21/19)

ETABS Model

Gravity System

01/31/19 Our understanding is that for the linear DBE model, only columns that slope are included in the analysis

model (with their gravity loads per floor applied directly to the column nodes at each floor). It is not

clear though why slab-beams are included in the model. They appear to essentially have no stiffness or

mass - but it is not clear if there any purpose for these elements.

02/15/19 The slab-beams are included so their stiffness can be toggled on in the wind drift model for serviceability checks. Additionally, the

slab-beams will be included in the Perform model to capture their energy dissipation contribution.

It is correct that the linear DBE model does not see any contribution from the slab-beams, as their stiffness has been reduced to

zero.03/05/19

Comment

Resolved

9 of 20

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& Reference

Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)43 Linear Design Submital

(12/21/19)

ETABS Model

Wall Stiffness

01/31/19 Please justify the wall stiffness modifiers (f11,f22,f12) as utilized currently in the model. 02/15/19 Property modifiers for the WALL30-10ksi wall property are as given below

f11 = 1.0 - horizontal axial stiffness of the area element. This degree of freedom of the area element does not significantly affect

the stiffness of the core walls which is why the value is left at 1.0. Reducing this shear stiffness results in a larger discrepancy

between the ETABS example model that compares the displacements of a concrete wall modeled with area elements and a frame

element which is included in the comment folder.

f22 = 0.5 - Vertical axial stiffness of the area element. This stiffness integrated across the length of the core wall results in the wall's

flexural stiffness. The value of 0.5 matches the DE core wall flexural stiffness shown in Table 7-1 of the BOD. An ETABS model in

the comment folder shows how the deflection of a core wall modeled with area elements and a reduced f22 matches the

deflection of a core wall modeled with a frame element and a reduced flexural stiffness for the same loading.

f12 = 0.75 - Shear stiffness of the wall area elements. The value of 0.75 matches the value shown in Table 7-1 of the BOD.

m11 = 0.25 - Out of plane flexure of the area element about the horizontal axis. The out-of-plane degrees of freedom are reduced

to represent cracking in the wall but they don't significanly affect the results of the ETABS analysis.

m22 = 0.25 - Out of plane flexure of the area element about the vertical axis.

m12 = 0.25 - Out of plane shear of the area element.

03/05/19

Comment

Resolved


(12/21/19)

ETABS Model

Walls

01/31/19 In the model, it appears as though pier W2P2 sees a reduction in wall length between L2 and L1 - is this

accurate? If so, please clarify how the detailing (and boundary zone) will be handled at the edge of this

wall.

02/15/19 This is accurate. The opening in the wall at this location is larger due to required access to the freight elevator. The outside face of

the boundary element at the freight elevator opening continues upward for several stories to make sure that the boundary

element "laps" with boundary element for the smaller openings above. See the wall elevation with comments in the comment

folder.

Full confinement of longitudinal reinforcing will be used between L1 and L6 though this reinforcing has not yet been shown on the

core wall sections pending the results of the wall strain analysis for the full wall height.

03/05/19

Comment

Resolved


(12/21/19)

ETABS Model

Walls

03/05/19 The BOD Section 7.6 notes that rigid diaphragms would be used for the linear model above the 1st

floor, with semi-rigid diaphragms at the 1st floor and below. It was noticed that the current model has

semi-rigid diaphragms assigned up through the 6th floor. Although not anticipated to make a

substantial analysis difference, was there a particular reason that the semi-igid diaphragms were

carried up?

04/05/19 Due to the presence of sizable diaphragm openings and/or irregular floor plates at levels 2-6, the diaphragms were modeled as

semi-rigid in the ETABS model to ensure the worst case torsional effects were captured. As noted this did not lead to a substantial

difference in analysis results.

04/30/19

Comment

Resolved

46 MCE Submittal #1

(04/05/19)

Sloping Columns

(Comment #8)

05/10/19 (Refer back to Comment #8)

1) Please provide diagrams that clarify the intended load path at the top and bottom levels of all

sloping columns (preferably shown directly on the current structural plans at the appropriate levels).

2) Is slab bearing utilized at all to transfer the thrust loading?

3) Per the sample calcs shown on the first page of the Comment #8 response, please clarify where

B=1.35 is determined from. This appears to be the "B" value that would be provided for transfer

diaphragm shear in LATBSDC, however this does not coincide with the value noted in the BOD

(additionally, this would be treated as a axial drag action, not a diaphragm shear component).

4) For the diaphragm check shown on the 3rd page of the Comment #8 response:

a) It appears that slab tensile capacity is determined based upon the full representation of the mild

steel in the slab. Please clarify if any reduction was accounted for to represent the persistent gravity

demands upon the slab in the mild steel.

b) For the determination of the slab capacity, φBAsfy is utilized. However, the overall demand

determined on the first page was already determined by dividing out "B" (see #3 above), so it may not

be appropriate for "B" to also be used in the capacity of the components.

5) Where collectors are required to be added to transfer the thrust back into the core walls, are these

in addition to collectors provided for the overall diaphragm design?

06/04/19 1) See comment folder for diagrams showing load path and revised calculations.

2) No. The thrust loading is transferred from columns to walls by dragging the load into the diaphragm with added rebar, then

transferring to the walls via diaphragm shear. See response to item 1 for clarification.

3) The calculation uses B = 1.14, which is the value for axial force in columns per the BOD. This value is only used for comparison

with the DE forces to determine the controlling case. A calculation for this was provided with the response to comment 16 and has

been included in the comment folder for reference.

4a) The intent was for the added bars to lap with collector bars and transfer load directly into the walls. The calculations have been

revised to typically ignore distributed slab steel and provide dedicated rebar specifically for the load transfer.

4b) We agree and have revised the calculations.

5) Yes. For example, column 21 requires (4) #11 to resolve the thrust force and (6) #11 were provided for collector forces. Thus, a

total of (10) #11 bars are shown on S0216R.

Comment

46

Submittal: MCE Submittal (04/05/19; 04/22/19)

10 of 20

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)06/18/19 1)

a) For Col 21 (pg 12), the load path indicates thrust going to the drag steel and then from the drag steel

to diaphragm shear. The next step is the #10 bars lapping with the #11 collector bars - please clarify the

intended load path for the step between the transfer to diaphragm shear and lap with collection rebar.

b) For C27, is the L3 rebar as noted on Pg. 14 not currently shown in the drawings?

c) For C32, do the provided #10 bars at the column have adequate development length to resist the

thrust at the column?

d) For C22/37: Please clarify the calc of rebar for L06 - is this not a compressive, insteat of tensile,

demand?

2) Closed

3) Agreed - initial comment accidetly referenced B=1.35 instead of 1.14 as shown in the calc. However,

note that the bias factor should be applied to the action itself, not the demand from which the action

occurs. Therefore, for the bias factor associated with the in-plane thrust of the sloping columns, the

bias factor based upon a gravity column does not seem as consistent as the bias factor provided in the

BOD for in-plane forces of a diaphragm - please confirm/justify.

4) Closed

5) Closed

07/03/19 1) See below for specific responses, and comment folder for updated calculations.

a) Load is intended to transfer directly from the drag steel into the collector bars, without an intermediate step. The diaphragm

shear is checked to make sure the diaphragm is not overstressed in shear locally as the load is dragged directly to the collectors. See

pages 13 and 14 of new calculation.

b) Apologies for the confusion - the load is being transferred at L4 as shown on page 10. The calculation incorrectly referenced L3

and has been revised. Please see S0214R.

c) The bar length has been increased to 12' to accommodate the required length plus development length.

d) Yes, the load at level 6 is a compressive load between the columns. The rebar has been detailed within the slab to take the entire

compressive force without relying on the concrete area. The concrete in the diaphragm with a strut width equal to the column

width + 2*column depth has been assumed to confirm the compressive stress on the concrete is adequate without confinement.

3) The design of rebar to resist the axial thrust is based on a bias factor of 1.05 which matches the value in the BOD for diaphragm

axial and in-plane loads. The bias factor of 1.14 was only used to compare column axial forces between MCE and DE in order to

determine which case controls (based on the entry for gravity column axial-flexure). For simplicity the bias factor has been changed

to 1.05 for this comparison as well since the MCE still governs for all cases.

Comment

46R1

06/18/19 1)

a) For Col 21 (pg 12), the load path indicates thrust going to the drag steel and then from the drag steel

to diaphragm shear. The next step is the #10 bars lapping with the #11 collector bars - please clarify the

intended load path for the step between the transfer to diaphragm shear and lap with collection rebar.

b) For C27, is the L3 rebar as noted on Pg. 14 not currently shown in the drawings?

c) For C32, do the provided #10 bars at the column have adequate development length to resist the

thrust at the column?

d) For C22/37: Please clarify the calc of rebar for L06 - is this not a compressive, insteat of tensile,

demand?

2) Closed

3) Agreed - initial comment accidetly referenced B=1.35 instead of 1.14 as shown in the calc. However,

note that the bias factor should be applied to the action itself, not the demand from which the action

occurs. Therefore, for the bias factor associated with the in-plane thrust of the sloping columns, the

bias factor based upon a gravity column does not seem as consistent as the bias factor provided in the

BOD for in-plane forces of a diaphragm - please confirm/justify.

4) Closed

5) Closed

07/03/19 1) See below for specific responses, and comment folder for updated calculations.

a) Load is intended to transfer directly from the drag steel into the collector bars, without an intermediate step. The diaphragm

shear is checked to make sure the diaphragm is not overstressed in shear locally as the load is dragged directly to the collectors. See

pages 13 and 14 of new calculation.

b) Apologies for the confusion - the load is being transferred at L4 as shown on page 10. The calculation incorrectly referenced L3

and has been revised. Please see S0214R.

c) The bar length has been increased to 12' to accommodate the required length plus development length.

d) Yes, the load at level 6 is a compressive load between the columns. The rebar has been detailed within the slab to take the entire

compressive force without relying on the concrete area. The concrete in the diaphragm with a strut width equal to the column

width + 2*column depth has been assumed to confirm the compressive stress on the concrete is adequate without confinement.

3) The design of rebar to resist the axial thrust is based on a bias factor of 1.05 which matches the value in the BOD for diaphragm

axial and in-plane loads. The bias factor of 1.14 was only used to compare column axial forces between MCE and DE in order to

determine which case controls (based on the entry for gravity column axial-flexure). For simplicity the bias factor has been changed

to 1.05 for this comparison as well since the MCE still governs for all cases.

Comment

46R1

07/29/19 1)

a) Closed

b) Closed

c) Closed

d) Please clarify this response - has the concrete strut been checked to confirm confinement is not

required per ACI? Also, similar to (c) above, please confirm that if the rebar is intended to participate in

teh resistinance of the compression demands, the the rebar is fully developed by the center o fthe

columns at L06. (Similarly, at L05 and L04, please confirm that the bars have adequate development

length to engage the thrust demands as utilized in the calculations).

3) Closed

08/20/19 Compression struts have been checked for axial stress and do not require confinement per ACI - see comment folder.

At level 6 along grid G, rebar extents have been specified to ensure full development beyond the centerline of columns. Other

locations at levels 4-6 have been checked and bars are already fully developed beyond the column centerlines.

Comment

46 09/10/19

Comment

Resolved

47 MCE Submittal #1

(04/05/19)

Gravity Loads

Pg. A3

05/10/19 It appears that not all basement walls are considered as part of the self weight assignment - please

clarify or update

06/04/19 All basement walls have been added to the self weight load pattern, see revised model.

06/18/19

Comment

Resolved

48 MCE Submittal #1

(04/05/19)

Mass

Pg. A4

05/10/19 Per Section 7.7 of the BOD, mass was to be included at the first floor and the subgrade levels, however

it is noted in the calcs that mass is not applied at these levels. Please update or clarify.

06/04/19 Mass at level 1 and below has been added to the model. See updated model and revised calculations on page A5 for verification of

the design based on the additional mass.

11 of 20

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)06/18/19 1) Please summarize if there was noticeable global impact from this inclusion of mass?

2) In the PERFORM model, it appears that the L1-and-below mass was applied upon free nodes that are

then slaved to nodes connected to slab elements. Please clarify this strategy, and justify locations

where the mass is slaved to a node relatively far from it's applied location in plan (e.g., Level P1, slaving

name "P1-10").

07/03/19 1) Inclusion of the below grade mass did not substantially change the behavior or design results. See comment folder for further

discussion.

2) The below grade mass was applied to nodes representing the locations of gravity columns, as well as the corners of the core

walls and evenly distributed points along the basement walls. We consider this to give a reasonable spatial distribution of the mass.

Since the column coordinates do not necessarily align with the slab mesh nodes, any free nodes were slaved to the nearest

available slab node to ensure the mass was linked to the diaphragm. As pointed out there are mass nodes relatively far from

diaphragm nodes, generally at columns intersected by the ramps. Slaving the nodes to the unrestrained basement wall edge was

not desirable, and so linking the mass to the nearest diaphragm level was determined to be appropriate. Given the aforementioned

minimal impact of below grade mass overall, the contribution of mass at these few nodes is even smaller. Consequently, this

modeling is considered acceptable for these select nodes. Comment

48 07/29/19

Comment

Resolved

49 MCE Submittal #1

(04/05/19)

Fiber Elements

Pg. A11

05/10/19 Using only 3 fiber elements in steel and compression seems like it may be light - has a study been done

to confirm that 3 fiber elements will provide adequate behavior/information compared to 4 or more

fibers?

06/04/19 A comparison model was run with the number of fiber elements increased to 5 in each wall area element. The resulting drifts,

strains and wall shears were compared to the initial results and found to have minimal (< 5%) difference. Consequently, the use of

3 fibers appears to give sufficient accuracy for the analysis.

See comment folder for comparison of the analysis results.

Comment

49 06/18/19

Comment

Resolved

50 MCE Submittal #1

(04/05/19)

CB Modeling

Pg. A16

05/10/19 Please clarify where MY was originally calculated 06/04/19 See comment folder for sample calculation clarifying value on page 16 and supporting information for other values of My.Comment

50 06/18/19

Comment

Resolved

51 MCE Submittal #1

(04/05/19)

Column Modeling and

Acceptance

A29

05/10/19 1) Section 7.7 of the BOD noted that "All columns are directly included in the analytical model".

Additioanlly, acceptance criteria for the gravity columns was provided in Table 7-3 for both axial-flexure

and shear actions.

However Pg. A29 notes that only limited columns are modeled, and those that are include moment

releases.

It is not clear how the gravity columns are evaluated per the criteria provided in the BOD. Additionally,

on other very similar projects, it has been noticed that inclusion of the columns in the analysis model

(as was expected based upon the BOD criteria) often reveals some non-negligible demand transferring

between the core and columns near the base of the tower.

2) For the column axial checks that are shown on D14, please clarify if these include the outriggering

and vert. acceleration components of the PEER load combo's

06/04/19 1) All gravity columns at and above level 1 have been modeled directly, with gravity loads applied as nodal loads. Columns have

also been modeled as fixed instead of pinned. See calculation page B185 for P-M design of columns based on Perform analysis.

Note that below grade the column nodes do not align with the diaphragm mesh and each node must be slaved to an adjacent

diaphragm node. This increases the time for stiffness matrix factorization by a factor of 10, leading to impractical model run times.

Consequently, below level 1 only columns that form slab frames or are part of sloping columns have been modeled with frame

elements. For the remaining columns nodes have been included to capture the mass only and are slaved to adjacent diaphragm

nodes. In this way the below grade mass has been captured and the outrigger effect above grade is captured, while maintaining

practical analysis times. Seismic effects in non-outrigger columns below level 1 are expected to be small due to small interstory

drifts. The exclusion of the below-grade columns and the inability of Perform to report the column deformation-compatibility

moments is not expected to significantly affect their design.

2) The referenced page shows the gravity design of columns for reference only, with load combinations per ASCE 7-10 (including

vertical seismic effects). Per response to item 1, see calculation package for MCE column design.

06/18/19 1)

a) PLease clarify the boundary condition assumptions that were utilized at grade when columns were

not modeled continuously below grade.

b) Please provide a global plot that provides total core shear vs. all columns shear along the height in

both directions

2) Closed

3) Please provide confirmation of the shear design of the columns. Although actual demands are

anticipated to be large throughout the majortiy of the structure, similar PBD projects have shown a non-

negligible amount of shear distributed to the columns near the base of the core walls.

4) Which model did the demands come from (short/long period)?

5) Based upon the sample spCol input/output, it appears that the max compression & tension was

analyzed in conjuntion with the worst-case moment in both directions (assumed to all be average of

maximums). Please confirm or clarify if the method differed from this.

6) Please provide a breakdown of one of the worst-case columns (say Col 35) that shows the demands

extracted from the model (grav and seis separtely), the load combination to determine the axial and

flexural demands, and how the value "B" was utilized in the evaluation.

7) Is there a unique condition at Col 21 Level P1 that results in such a large Mmaj,max value? Also,

please confirm that this column analysis at the floors below is correct - is it accurate that even though

level P2 and P3 see a substantial reduciton in this Mmaj value with relatively minor differences in axial

load, the DCR remains constant at 0.75? A similar situation appears to be present at Col 27.

07/03/19 1) a) Columns that are not modeled below grade have a vertical support at level 1 and are rotationally unrestrained. Due to the

minimal displacements below level 1 we expect the deformation compatibility demands for columns below grade to be

insignificant.

b) See comment folder for plots of core and column shear over the building height. It is observed that the story shear primarily

distributes to the core except in the lowest stories (~ levels 1-5) where the columns shares a larger portion of the story shear. This is

more pronounced in the H2 direction which is the short direction of the core and aligns with most column strong axes. Separate

plots have been generated for the short and long period time histories, but the results are similar.

3) See comment folder for shear calculations. The columns are adequate to resist the shear demands from the Perform3D model at

all locations.

4) Demands have been taken from both models. For each, the average of the maximum seismic demands is calculated and the

short and long period results are then enveloped to determine the maximum demand. This value is then multiplied by 1.3 per the

acceptance criteria in the BOD.

5) Confirmed. Results are enveloped as described above and the controlling Pu,max and min are evaluated in combination with

Mu. See also calculations provided in response to item 6 below.

6) See comment folder for a breakdown of the demands, the load combinations, and the P-M checks for column 35. Conservatively,

the bias factor was not included in checking P-M DCRs since all cases work without this increase in capacity.

7) At level 2 and below columns 21 and 27 grow in size while also rotating to run parallel to the short direction of the core. At the

same time the level 2 diaphragm steps in, so these two columns are the only support for the west edge of the diaphragm beyond

the core. This combination of location and size causes the columns to act as outriggers and pick up notable shear and moment. See

comment folder for a graphic explaining the demands at level 2 and below. Note that the columns are designed for the maximum

moment at end i or j for each level, so the controlling moment is the same for level 1 and P1.

Since level P1 controlled the design, this DCR was conservatively copied down to levels P2 and P3. The calculations provided for

item 51-6 (pages 11-21) have refined the DCRs for each level and show a significant drop in DCR at P2 and P3.

Comment

51

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& Reference

Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)07/29/19 1a) Previous comment response is understood to note that where columns are not modeled

continuously below the level 1, they terminate at Level 1 with a pinned vertical support. Please clarify if

there is specific detailing that supports the modeling of these columns as pinned, or provide an update

to this study that includes representative fixity at the base of these columns in order to not

underestimate their potential demands.

1b) Closed

2, 4, 5) Closed

3), 6) , and 7): Given 1a) above, please confirm column shears are not under-represented.

08/20/19 1a, 3), 6) and 7) The shear demands from the first story columns have been conservatively increased by a factor of 3, and have been

found to have adequate shear capacity to resist the demands. A single time history run has been performed with the columns

rotationally fixed at level 1 to compare the shear in the column with and without the rotational fixity to verify that this factor is

acceptable to envelope the shear demands - see comment folder for supporting calculations.

Additionally, below grade columns have been evaluated for the shear demand due to imposed seismic drifts and deformation

compatibility. Calculations are provided in the comment folder.

Comment

51 09/10/19

Comment

Resolved

52 MCE Submittal #1

(04/05/19)

Global Shear Results

B8

05/10/19 1) On Pg. B8, it appears that there is a relatively large amount of shear added to the core wall at L2 -

however the floorplate at this level appears relatively minimal compared to the typical floors. Is it clear

what is causing this large increase in shear at this level?

2) For plots comparing shear vs. Ag*root*(f'c), please normalize the demands by B*φ in order to to

have a straight comparison of the required Vn vs. the ACI shear stress limits.

3) Please provide a Vu/W plot along the height of the building in each direction from the NL analysis

06/04/19 1) The shear load is from section cut "CoreL02", which is taken just above the 2nd floor line and reflects the load within the second

story. Since level 3 has the largest above grade floor area, the jump in shear entering the second story is reasonable.

The core wall shear plots incorrectly showed the wall shears one story below where they should have. This has been corrected in

the updated wall shear plots in the calculation package starting on page B8.

2) The plots have been normalized as requested. Note that since B*φ = 1.163, the reported shear stresses are lower than previously

reported. These plots are included in the calculation package starting on page B12.

3) See comment folder for the added plot (also included on page B16 of the revised calculations).

Comment

52 06/18/19

Comment

Resolved

53 MCE Submittal #1

(04/05/19)

Wall Strain Results

B17

05/10/19 BOD Section 7.5 states that the structure will be proportioned in provide a yielding mechanism of

flexural plastic hinging in the shear wall. However, the values shown on Pg. B17 show that there is

essentially no average flexural yielding being reported (other than one corner of L1, which is just

slightly beyond yield strain). Please comment and discuss the wall flexural behavior in light of the

intended behavior.

(Note that this is similarly seen on Pg. B180, where the shear walls visually account for a negligible

amount of energy dissipation)

06/04/19 We agree that there is little yielding associated with flexural hinging at the base of the wall per the results presented in the design

package. Our interpretation of the MCE results is that the rectangular shape of the core helps to resist the overall seismic flexure

by providing a stiff and strong section in the deep direction of the core, and by providing a significant flange length in the other

direction of the core.

The rectangular shape of the core requires two coupling beams along the long wall of the core instead of a single coupling beam

which would be used for a more square-ish shaped core. The energy dissipated by the large quantity of coupling beams helps to

limit the demand on the core wall sections for this "shorter" tall building.

The tension strains shown on MCE calculation package page B18 shows that yielding in the core wall is most significant in the

intended hinging zone, but KPFF has further optimized the design of the core wall section at the base of the wall to ensure that

yielding for first mode behavior occurs as intended. The axial-flexural design of the core wall sections have been revised so that

they have up to a DCR of 1.05 for the DE condition. Updated response plots are provided in the calculation package, and new linear

DCRs are shown in the comment folder.

Comment

53 06/18/19

Comment

Resolved

54 MCE Submittal #1

(04/05/19)

Wall Strain Results

B18

05/10/19 For the strain plots starting on B18, please clarify which strain gauge results this represents? Or is this

an envelope of all of the strain gauges?

06/04/19 The plots for each time history show the maximum of all strain gauge readings at a given level for the given record. The average of

the maximum for each record is plotted in red. This average maximum is used to determine acceptance on the previous page (B18

of the revised calculations). 06/18/19

Comment

Resolved

55 MCE Submittal #1

(04/05/19)

Wall Strain Results

B18

05/10/19 For the strain plots (as well as other parameters like drift, etc…) there appears to be a few records that

act as outliers that result in substantially "larger" results than the other records. Please clarify if it is

understood why these particular records are resulting in more substantial response.

06/04/19 For the long-period CMS records, the Chi-Chi and the Loma Prieta records stand out as the most demanding on the structure.

Review of Figure 3-4 (on page 13) of the ground motion memo shows that the response spectrum for the Chi-Chi record is about

double the average of the response spectra for the 11 ground motions near the 1st period response of the building. Similarly, the

Loma Prieta record extends significantly above the average spectra between 1-2 seconds which could affect the initial response of

the structure before significant yielding and the 2nd mode response of the building after yielding. Both of these records are crustal

records with significant velocity pulses as shown in Table 3-4 (on page 8) which are likely the cause of the larger demands on the

building.

For the short-period CMS records, the San Fernando - Hollywood and the Coalinga records stand out as the most demanding on the

structure with the San Fernando record producing the largest strains lower in the building and the Coalinga records producing the

largest strains higher in the building. Both of these are crustal records but neither has a significant velocity pulse per Table 3-4. Per

figure 3-1, the response spectrum of the San Fernando - Hollywood record is significantly more than the average in the area of the

primary building period, and the Coalinga record has a peak above the average response spectrum around the 2nd mode response

of the building. Geoengineers attributes these peaks in the response spectrum to the natural variability of the ground motion

records due to the fact that the records are not spectrally matched.

06/18/19

Comment

Resolved

56 MCE Submittal #1

(04/05/19)

Drift Results

B22

05/10/19 While understood that it is not an exact comparison, the reported average drifts from the NL model are

essentially the same in the H1 direction, and only slightly larger in the H2 direction, than the drifts

determined from the DBE analysis. Please discuss and justify.

06/04/19 As discussed in comment #53, there is less yielding of the longitudinal bars in the core walls than expected. The stiffness

assumptions in the linear DE model are broad and reduce the flexural stiffness of the wall along its entire height to 50% of its gross

flexural stiffness. The MCE results indicate that the 50% reduction, which assumes some amount of yielding and damage in the

core wall, likely overestimates the stiffness reduction in the core wall for the DE condition.

Use of stiffness based on specified material properties for the DE model and use of stiffness based on expected material properties

for the MCE model also contributes to this discrepancy.06/18/19

Comment

Resolved

57 MCE Submittal #1

(04/05/19)

Column Axial

B186

05/10/19 1) Please clarify the use of (1.2+.12Sms)D in the load combinations in lieu of the (1.2+0.2Sms)D per

PEER Section 6.8.3. Note that the 0.2 value in the load combination is used later on B198 for shear

walls.

2) Was slenderness accounted for the columns? In particular, the taller-height columns at L7?

06/04/19 1) Please note that page B198 erroneously showed 0.2Sms but 0.12Sms was used in calculations, consistent with the referenced e-

mail above. Page B198 has been updated and is included in the comment response folder. The use of the combination

(1.2+.12Sms)D is made at the recommendation of Ron Hamburger who recommended the use of the combination on a different

tower project due to an upcoming change in the provision. An e-mail confirming this conversation is attached in the comment

folder. An updated calculation sheet has also been provided to clarify the use of 0.12Sms.

2) Yes, slenderness effects were accounted for. See comment folder for sample calculation.Comment

57 06/18/19

Comment

Resolved

13 of 20

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)58 MCE Submittal #1

(04/05/19)

Wall Shear Design

B198

05/10/19 1) This section references the 2017 LATBSDC, although this code reference was not clear from the BOD

(in lieu of PEER TBI). Please clarify.

2) Please clarify if Sms in the load combination came from the mapped or site specific value. Is there a

large discrepency between the two?

3) Is something unique occuring at L1, Wall W2 and W2P3? These two piers appear to see a large

increase in shear directly at L1.

06/04/19 1) This section has been adjusted to reference the basis of design in the updated calculation package page B202. Per response to

item #57, the previously submitted calculations were in accordance with the basis of design.

2) The mapped value (Sms = 1.355) per ASCE 7-10 was used. While not directly provided, a value of 1.716 could be inferred from the

10/9/18 geotechnical memo (table 1-7, KSSI adjusted at 0.2s). Though this is notably higher than the mapped value, the relative

dead load factors for seismic combinations (1.2+0.12Sms) are 1.36 using the mapped value v. 1.41. Since the variation in the dead

load factor is only 3%, the difference in total factored load is even lower and the discrepancy is not significant.

3) The wall opening between these two piers is not present below level 1, and the solid wall below the piers provides more fixity to

the base of these piers as compared to the same piers further up the building. Additionally, since the openings on the south face of

the wall continue down there is more fixity to the north wall overall at this level, compounding this effect.

06/18/19

Comment

Resolved

59 MCE Submittal #1

(04/05/19)

Diaphragm Design

B225

05/10/19 1) This section (and the one on C2) references the 2017 LATBSDC, although this code reference was not

clear from the BOD (in lieu of PEER TBI). Please clarify.

2) This page notes that section cuts were used at the interface between core walls and slabs to

understand the diaphragm shear demands and force transfer. Was this done in a side model that

included slabs? Please clarify the process of determination of the diaphragm demands at MCE.

3) It is noted that the slabs have not yet been evaluated for gravity loads, and that the capacity of the

diaphragmwill be increased for teh combined loads where required. Does this just mean that the

gravity reinforcement will be added in addition to the diaphragm reinforcement?

4) Please clarify if chords were designed.

06/04/19 1) The load combinations for these calculations are per section 7 of the basis of design. The narrative has been revised to reference

the BOD as seen on page B229 of the updated calculation package.

2) MCE diaphragm demands were determined by taking section cuts at the interface of wall piers above and below a given floor

level. The narrative has been revised to read: "Section cuts were taken at the interface between walls elements above and below a

given level to assess the diaphragm shear demands and force transfer. " See graphic on page B230 and revised calculation section

for additional clarification about which model elements are included in the section cut. The section cuts as shown report the

maximum difference in the wall shear above and below a level, which is the same as the force transferred in at that level and can

be created in a model that has rigid diaphragms.

3) Yes, any reinforcing required from the gravity analysis will be additive to that required for diaphragm reinforcing.

4) Diaphragm chord calculations have been added to the calculation package, starting on page B252.

06/18/19 1) through 3): Closed

4) Please clarify what rebar was entered into spColumn for the chord checks - is this the same rebar

that is also utilized in the diaphragm shear calcs? Or is this a separate mesh, to avoid double-counting

for both shear and chords?

07/03/19 4) During the 6/24/19 conference call it was agreed that flexural DCRs of 0.80 or below will allow for potential interaction of shear

and flexural demands in the diaphragm. Per pages B254-255 of the previously submitted calculations, the flexural DCRs for typical

diaphragms satisfy this criteria, with a max value of 0.72 and typical values below 0.50.

The diaphragm designs at levels 2 through 7 have been re-checked and provided in the supplemental diaphragm calculations

package. See page 67 of this package for a summary of diaphragm flexural DCRs, showing values limited to 0.80 per the conference

call discussion. Chord steel has been added where required to maintain this DCR.

07/29/19

Comment

Resolved

60 MCE Submittal #1

(04/05/19)

Diaphragm Design

B226

05/10/19 1) Is the full length of the wall (including CB length) utilized for shear transfer through shear friction? If

so, please justify transferring over the CB length (and clarify how the axial loads in the CB are resolved).

Aslo, justify that the load will viable load path through the CB's given that these components will be

undergoing substantial yielding.

2) Please clarify where the shear friction from slab-to-wall calc is provided.

3) Are the demands shown on this page already factored through the required load combinations?

4) In the spreadsheet, the V/Side appears to be 0.6*the H1 shear - please clarify the intent behind this.

06/04/19 1) Diaphragm design in the H1 direction has been revised to ignore the length of the coupling beams. See updated calculation

package starting on page B231.

2) See comment folder for calculation.

3) Yes, loads have already been factored per the load combinations described in the basis of design.

4) See revised calculations based on section cuts on each face of the wall at each level starting on page B231 of the updated

calculation package. The 0.6 factor was meant to be a conservative estimate, assuming equal diaphragm shear on each side of the

wall and then increasing by 20% to account for torsion. In the revised calculations the shear is taken directly from section cuts at

each level on each side of the wall.Comment

60 06/18/19

Comment

Resolved

61 MCE Submittal #1

(04/05/19)

Diaphragm Design 05/10/19 Please provide a diaphragm design for L2, and L31 06/04/19 See comment folder for L2 and Roof diaphragm designs. Note that L31 design was previously provided and updated calculations are

included on pages B231 and B232 of the updated MCE calculation package.06/18/19

Comment

Resolved

62 MCE Submittal #1

(04/05/19)

Diaphragm Design

General

05/10/19 1) In most instances, it is unclear if diaphragms were designed for load potentially occuring in either

direction (in lieu of just one direction) - please clarify and confirm.

2) All diaphragms appear to have an opening adjacent to the core wall along the southern edge of the

wall. Please clarify the load path to utilize the full length of the remaining section of wall in shear

friction to transfer diaphragm forces when the load is applied towards the west (e.g. when load would

be applied towards the bottom of the page in plan). Is there a load path provided that allows for the

shear to be dragged back up to engage the full length of this wall in shear friction?

06/04/19 1) Diaphragm loading in either direction was considered. The calculations on B226 and B227 use the following hierarchy to evaluate

the diaphragm:

a) Deliver entire load to wall via direct diaphragm shear over the available length of wall

b) If the wall length is insufficient, the remaining force is delivered entirely through compression bearing on one side of the wall

only. This force is capped at 0.5f'c over the bearing length of the diaphragm to avoid the need to provide confinement reinforcing

in the diaphragm. The diaphragm is then verified to have sufficient length available to deliver this bearing force to the wall end. The

bearing checks are performed on each side of the wall to account for load reversibility.

c) Finally, any load that cannot be delivered through direct shear or slab bearing is taken out through the addition of tension steel.

These bars are added to both sides of the wall to account for load reversibility.

2) The H2 shear delivered to the wall parallel to E.4 has been checked against the compression bearing capacity of the slab against

one end of the wall. Typically the bearing capacity exceeds 50% of the shear, indicating the tributary load north of the core can be

delivered directly through bearing. The dowels are still conservatively sized assuming the entire load along this line will be

delivered through direct shear before the addition of collectors. At level 3 only, additional collectors have been provided. See

comment folder for supplemental calculations.

Comment

62

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)06/18/19 1) Closed

2) Please clarify response. It is understood that a large amount of the shear transfer is able to take

place through bearing at one edge of the wall, but it appears that the tensile collector at the other end

is still determined based upon assuming that the full length of the shear wall can be engaged in shear

friction.

For instance, at L6 @ the top of Pg. B237, with a shear transfer of Vu=1943k, a bearing capacity of

1229k, and a tensile collector capacity of 384k, the resulting required shear friction transfer would be

1943-1229-384 = 330k. Dividing out by the 25.6klf capacity requires a length of 12.9' (the full length of

the available wall segment), which of course makes sense since the tension collector was based upon

this assumption in the calcs.

Therefore, with an example like above, please clarify how the response resolves the load path around

the opening.

07/03/19 2) Please note that the load delivered to the wall in shear friction has been limited to the 12.9' length of wall (The length excluding

the opening) for all cases. In the supplemental diaphragm calculations package the spatial distribution of load on the east side of

the core has been considered to justify delivery of load through bearing and direct shear without adding reinforcing to drag load

around the opening. Locally, tension in the available slab steel is relied upon to deliver load to the shear/bearing regions of

diaphragm on the wall. See for example page 34 of the supplemental calculations.

07/29/19 2) In terms of the justification of load path around the opening (given the example on Pg. 37 of the

supplemental diaphragm calcs):

It appears that all #4 bars within the slab between the opening and the edge of slab were added up and

utilized to determine a φBRn capacity to justify dragging load around the opening. This appears to

calculate this action more like one wide collector compponent. It is not clear that all of the bars,

particularly those closer to the eastern edge of the slab, will engage in this fashion - please update or

clarify.

Also, it is expected that due to the resistance of other actions (including slab shear), the full capacity of

these bars would not be available in this fashion - please justify or provide a representation of the

available capacity of these bars.

08/20/19 The slab steel is no longer utilized for this check. At levels 4-6 collectors have been added adjacent to the wall to replace the

previous force that was delivered by slab steel. The dowels into the wall along grid E.4 have been revised to accommodate the

added load.

At level 7 the existing collectors have been increased to take the load that would have been delivered by the slab steel.

At levels 2 and 3 the slab steel was not utilized, so no changes have been made.

Comment

62 09/10/19

Comment

Resolved

63 MCE Submittal #1

(04/05/19)

Diaphragm Design L3

B228

05/10/19 1) For the EW direction, Pg. 226 showed the 600k collector force all in bearing at the end of the wall.

On Pg. B228, it is shown in both tension and compression - please clarify. Also, the compression calc is

checked using rebar only - please clarify why this was not done checking the concrete in compression

(and confirming whether confinement per ACI is required).

2) Under"Collector Lengths", please clarify what is meant by including the "3' opening" in the

compressive collector length

3) Shear friction calc assumes slab and wall poured monolithically - is it confirmed that this is the case

(applies to all diaphragms)

4) Please clarify how the load path for the offset collector at the southeast wall corner works (sim at L4

through L6). Is the load from the drag transferred (and added) to the sheaer friction bars? What is the

offset from the center of these drag bars to the wall? Is there perpendicular rebar that resolves the

eccentricity?

06/04/19 1) The previous calculations showed the entire collector force being taken in bearing for the positive H1 direction. In the negative

H1 direction, bearing was conservatively ignored and the total collector force was taken in tension and compression steel.

The calculations have been revised for new loads as shown on page B231 in the updated calculation package. Concrete

compressive stresses have been checked and are below the threshold for confinement per ACI (see for example page B233).

2) To deliver the 246k load to the wall, the required diaphragm length is 246k/(2*15.7klf) ~ 8'. However, at the 3' wide opening near

E.4 the diaphragm shear capacity is ignored, so the length required is increased by the opening width to 11' from face of wall.

3) In our experience, local practice is to use jump forms for towers under ~400' and slip forms for taller towers. As this tower is well

under that threshold, it is assumed that jump forms will be used above grade and diaphragm calculations have assumed the slab

and wall are poured monolithically.

Due to the changing ramp elevations below grade, a slip form option is likely to be used at these levels. Consequently, the

calculations for P3-P1 have been revised to account for a lower shear friction capacity.

Calculations will be revised for the lower capacity should the contractor request to change the approach for above grade levels.

4) The load described is resolved through shear friction at the diaphragm/wall interface. See comment folder for calculation

regarding eccentricity.

Comment

63

05/10/19 1) Closed

2) Closed

3) Is there a particular note or detail in the drawings currently that specifies this assumption regarding

construction sequence above and below grade?

4) Per the Comment #63 exhibit, it appears that the load from the collector is transferred to the wall

through additional shear friction. Please confirm that these additional shear friction bars are provided

over the length of the collector overlapping with the wall. Similarly, response to Comment #60 noted

that Coupling Beam Lengths are not considered as part of the shear friction transfer length. For Floors

4-6, almost half of the offset collector length is adjacent to a coupling beam - has this been taken into

consideration?

06/04/19 3) Notes have been added to 2/S0320 and 4/S0320 stating that the slip formed connection is to be used below level 1, and the jump

formed connection is to be used at level 1 and above. If the contractor decides to vary from this approach, they will have to

contact KPFF which will give us the opportunity to adjust the design.

4) Confirmed - the extents of the T5P2 dowels along this wall cover the enitre length of the collector. See S0214R, S0215R and

S0216R.

The added bars at this corner are not required to drag load to the wall, as force in this direction can be resolved entirely through

bearing on the opposite end of the wall and direct shear (see pages 38, 41 and 43 of the calculations). Since collectors are required

at the opposite side of the wall, nominal collector bars were added here for redundancy only. Please note that per response to

comment 92 the length required to develop collector forces into the walls has been calculated and is typically under 5'. Thus, the

15' provided is sufficient to transfer the load while ignoring the ~4' length of coupling beam where it occurs.

15 of 20

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Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)07/29/19 3) Closed

4) See response to Comment #92, as the embedment calculation method currently shown in the 07/03

calcs is not recommended (compared to current NEHRP guidelines). Additionally, the method

calculated in #92 (the current response as well as the recommended update) is based upon

embedment of the collector into the thickness of the wall. In this case, the collector is offset from the

wall and demand is transferred through shear friction dowels which presumably do not occur along the

length of the coupling beams.

Reponse notes that in the specific example location, these offset collector bars are actually redundant

and not necessary. Therefore, the above conversation may be moot at this location - however please

clarify at other locations where offset collector bars are required and provided.

08/20/19 4) See response to comment #93. All locations with offset collector bars have been checked to deliver load from slab to walls

through shear friction, and the length of the coupling beams has been ignored when calculating the force transfer.

Comment

93 09/10/19

Comment

Resolved

64 MCE Submittal #1

(04/05/19)

Diaphragm Design

Lower Levels

05/10/19 At the lower level diaphragm design with the notches at the north and south edges, please clarify

if/how shear and flexure were evaluated along the strips of slab on either side of the notches.

06/04/19 See comment folder for calculations of shear and flexure in these areas.Comment

64

06/18/19 1) For the shear calcs at L6 and 5/4 in the Comment #64 exhibit, please evaluate the upper and lower

sections of the diaphragm segments independently.

2) For the chord check, is the spCol analysis utilizing slab bars that are also utilized for shear resistance?

07/03/19 1) See comment folder - shear calculations have been provided that check diaphragm segments independently at levels 4 through

6. The shear at critical sections is calculated by hand using the area of slab regions to estimate the spatial distribution.

2) See comment 59 for discussion of slab rebar for shear/flexure based on the 6/24/19 conference call.Comment

64R1 07/29/19

Comment

Resolved

65 MCE Submittal #1

(04/05/19)

Diaphragm Design -

General

05/10/19 Please clarify if bearing is being utilized to transfer diaphragm shear at the southeast wall corner

adjacent to the slab opening.

06/04/19 At this location bearing is utilized only over the available wall width of 8", resulting in a capacity of φBRn = 0.65*1.05*8"*7.5"*6ksi =

246 k.06/18/19

Comment

Resolved

66 MCE Submittal #1

(04/05/19)

Diaphragm Design - L7

B231

05/10/19 Please clarify how the required tension collector length calculated here correlates with the collector

lengths determined on B227

06/04/19 The length shown on B227 was the length required to develop the bearing force into the diaphragm. Page B231 showed the

required length of tension collector. For forces in the positive H2 direction there is insufficient diaphragm length to develop the

entire bearing force from edge of diaphragm to face of wall, so the bearing capacity is limited and the balance of the load is taken

in tension. This has been clarified on page B234 of the updated calculation package.

06/18/19

Comment

Resolved

67 MCE Submittal #1

(04/05/19)

Diaphragm Design - 8-

30

B232

05/10/19 The shear friction calculation on this page appears to the use the diaphragm capacity determined at the

top of the page, rather than a shear friction calc - please clarify.

06/04/19 The section "shear transfer at wall" is based on the controlling shear capacity. In this case, the diaphragm shear capacity (15.7 klf)

controls over the shear friction capacity (20.5 klf). See also response to comment 60 for clarification of shear friction capacities.06/18/19

Comment

Resolved

68 MCE Submittal #1

(04/05/19)

Diaphragm Design

MCE vs DBE

B233

05/10/19 1) Was MCE demand only compared to Fpx? Or was it also compared to the maximum demands from

the DBE analysis?

2) Does MCE still govern over drag components that would utilize overstrength load combinations in

the DBE code design?

06/04/19 1) Fpx typically governs over the DE story forces. However, the calculation has been revised to show story forces at each level for

comparison. See comment folder for updated calculation.

2) The MCE collector forces govern - see comment folder. At each level the DE collector forces have been calculated and, where

collectors are required, amplified by Ωo. Comparing to the required MCE collector forces, the amplified DE forces are still at least a

factor of 3.3 lower.

Comment

68 06/18/19

Comment

Resolved

69 MCE Submittal #1

(04/05/19)

Foundations

B235

05/10/19 1) Did column demands on the mat foundation include the seismic outrigger + vert. acceleration

demands?

2) How were the foudnations under the basment walls determined?

3) Please clarify if PEER TBI or LATBSDC was utilized as the reference document.

06/04/19 1) Yes, seismic forces from the MCE analysis were included in design of the mat foundation.

2) Basement wall foundations are being added to the previously submitted SAFE model, and design calculations will be included

with the response to the foundation comments.

3) Load combinations and acceptance criteria are per the basis of design.06/18/19 1) Closed

2) Pending updated SAFE model and associated updated analysis

3) Closed

xx/xx/xx

07/29/19

Comment

Resolved

70 MCE Submittal #1

(04/05/19)

L1 Diaphragm

C2

05/10/19 1) How were the slab steps accounted for in the L1 diaphragm design? Are these major steps?

2) On C3, the sheet is labeled "Wall Forces". Please confirm if the values in the diagram represent wall

forces, or diaphragm transfer forces.

06/04/19 1) Level 1 slab steps are generally small with the highest elevation of the level 1 slab at el 130'-9" and the lowest elevation at 129'-

0". The level 1 diaphragm is idealized as level at elevation 130'-9" in the Perform and ETABS models.

The BOD has been changed on page 22 to state that the level 1 diaphragm is modeled as a flat plate due to the minor changes in

the diaphragm elevation.

2) The values are diaphragm transfer forces per the section cuts DL01N, DL01S, etc.

06/18/19 1) Are there any locations where a single step is greater than the thickness of the slab itself? If so, is

any special detailing provided at this lcoation for continuinty of the transfer diaphragm?

2) Closed

07/03/19 1) Detail 4/S0443 has been added to provide shear transfer across slab steps larger than the slab thickness. See comment folder for

calculations. Note that at locations where a concrete beam has already been provided, the in-plane shear transfer will be OK by

inspection due to the enlarged section, and no special detailing has been provided for these conditions.Comment

70 07/29/19

Comment

Resolved

71 MCE Submittal #1

(04/05/19)

L1 Diaphragm

C7

05/10/19 1) For the Force Transfer calculation on this page, FN,2 determines the maximum compressive stress

allowed without confinement. Note that this would be the acceptable stress, however the capacity

upon which should be subtracted from the design demand should be φBRn (where Rn=0.5f'cAc)

2) Pg. C8 notes a step beam at the tension drag. Please clarify how deep this beam is, and how it

interacts with the drag steel.

06/04/19 1) Calcs have been revised to reduce capacity by φBRn, see page C7 of updated calculations. Note that as forces at the core have

not changed substantially in the new models, the previous design forces are still used. The forces at the basement walls have been

updated.

2) A 24" wide x 32" deep beam has been provided per plan for the slab step. The required drag steel will be located within this

beam depth, and is in addition to the beam longitudinal reinforcing. See comment folder for clarifying sketch. Note that the

location of the slab step was shown incorrectly on the drawings submitted with the earlier designs, implying that the beam did not

intersect the core walls. The update drawings included in this submittal make clear that the beam and the longitudinal drag bars in

the beam are in the plane of the core wall.

16 of 20

# Date of

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Date of

Resolution

Resolution/

Notes

Topic

& Reference

Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)06/18/19 1) Although representation of FN,2 has reduced, the resulting reinforcement has actually decreased,

although this appears to be due to a typo in the Vw,n calc (from the 04/05/19 submission) - is this

correct?

Similarly, please confirm that demands are not noticeably larger even though mass is now included at

L1?

2) Closed

07/03/19 1) Confirmed. The revised calculations submitted on 6/4/19 corrected Vw,n and the required reinforcing. The original results

miscalculated Vw,n as (Fn,1+Fn,2)/Lw,n, despite the intent to use the equation (V-Fn,1-Fn,2)/Lw as written.

The change in forces was compared on page C3 and showed that no demands were noticeably larger. To summarize the changes:

North - 3328k (3% reduction compared to 3444k)

East - 1887k (11% reduction compared to 2126k)

South - 4857k (2% increase compared to 4740k)

West - 3218k (19% reduction compared to 3989k) 07/29/19

Comment

Resolved

72 MCE Submittal #1

(04/05/19)

Basement Wall Force

Transfer

C13

05/10/19 1) Please clarify where the dowels are shown on plan for the north and south wall force transfer.

2) Please confirm that the callouts for the east and west wall are correct. Our understanding of the

drawings shows dowels at 18", not 9".

06/04/19 1) Dowels have been added to plan.

2) Drawings have been revised to match calculations.

06/18/19

Comment

Resolved

73 MCE Submittal #1

(04/05/19)

Basement Wall Design

C16

05/10/19 While it is understood that the flexural design of the basement walls are considered to be acceptable

be inspection, please clarify how the load path for the out-of-plane soil loads was handled. Is it

assumed to be resisted by the vertical bars in the basement walls? If so, is the out-of-plane demand

substantial engouh that the stress between flexural and out-of-plane soil needs to be evaluated?

06/04/19 The out-of-plane soil loads are assumed to be supported by the basement walls spanning vertically between the below-grade

diaphragms like a beam. The vertical bars in the walls resist the flexure that is created by this behavior.

Calculations for the out-of-plane design of the basement wall have been included in the comment folder.06/18/19

Comment

Resolved

74 MCE Submittal #1

(04/05/19)

Basement Wall Design

C18

05/10/19 Given that the basement walls are considered "Ordinary" force-controlled components, it would be

expected that they would utilize a φ=0.9 and B~1.05 (given the h/lw of the walls)

06/04/19 Calculations have been updated for the referenced φ and B values and are included in the updated calculation package starting on

page C16. Wall reinforcing has been adjusted for the northwest wall.06/18/19

Comment

Resolved

75 MCE Submittal #1

(04/05/19)

Below Grade

Diaphragms

C25

05/10/19 For the diaphragm shear checks on this page, was any accomodations made for stresses already in the

reinforcement due to gravity loading?

06/04/19 Stresses due to gravity loading were not considered at this time. Any reinforcing required from the gravity analysis will be added to

that required for diaphragm reinforcing. See also response to comment 59-3.06/18/19

Comment

Resolved

76 MCE Submittal #1

(04/05/19)

Panel Zone

D2

05/10/19 Please provide a short narrative for this calc and what capacity is being determined 06/04/19 See comment folder for narrative and calculation.Comment

76 06/18/19

Comment

Resolved

77 MCE Submittal #1

(04/05/19)

Perform Model 05/10/19 Please note that it appears that a few elements are missing from the "Core" section cuts - please

update associated results if this provides a non-negligible difference in the reported core response.

06/04/19 The "CORE" section cut has been updated to capture all relevant nodes. Please see updated results starting on page B3 of the

updated calculation package.

06/18/19

Comment

Resolved

78 MCE Submittal #1

(04/05/19)

Perform Model 05/10/19 Please clarify how gravity demands on the basement walls were determined for foudnation design.

Was this determined externally (given no applied demands in the PERFORM model)?

06/04/19 Yes, gravity demands on basement walls were determined externally for input into the SAFE model. See also response to comment

69. Clarification will be provided with the response to foundation comments.

06/18/19 Pending further calcs (also see Comment #69) xx/xx/xx07/29/19

Comment

Resolved

79 Foundation Submittal

(04/22/19)

Foundation Drawings 05/24/19 S0200: Please clarify the construction of the two column spread footings along Grid B. On S0200, these

appear to be directly adjacent to the mat foundation - please confirm that these are still poured

separately, and these are not considered as extensions to the mat foundation?

06/11/19 Confirmed - the column footings at B/2 and B/5 are poured separately and are not extensions of the mat foundation.

07/29/19

Comment

Resolved


(04/22/19)

Foundation Drawings 05/24/19 S0200: Near Lines E and E.4 along the northern side of the core, there are two concrete walls that

connect the core to the basement wall and it's foundation. Please clarify/justify that these walls will

not impact the design of the foundation (given that they are currently ignored), as the foundations for

these walls do appear to tie together the mat and continuous basement footing.

06/11/19 The 10" walls form the sides of a detention vault and are approximately 6' tall from foundation to underside of P3 ramp. These

walls are not designed to resist seismic forces and while they may create a load path for shear connecting the mat foundation and

the foundations at the perimeter wall they are not designed to resist these loads so they will experience localized damage during a

significant seismic event.

Because of the walls function as a detention vault, the addition of a joint between the walls and the foundation could compromise

the ability of the vault to retain water. Because the use of joints would compromise the function of the walls, accepting the

localized damage which does not affect the seismic performance of the building is the best option.07/29/19

Comment

Resolved


(04/22/19)

Foundation Drawings 05/24/19 Will more specific sections/details be included in the drawings beyond the typical details of S0411? 06/11/19 Additional sections and details are not planned to be added to the foundation drawing sheets.

07/29/19 1) At the mat thickness transition details (such as 2/S0411), it is recommended to include vertical ties at

the initiation of thickness transition to resolve the vertical resultant caused by the change in angle of

the longitudinal bars.

2) Is there a detail provided that describes the base condition for the columns landing on the mat?

3) What is the specified reinforcement through the thinner slabs at the pits within the core? How was

this reinforcement determined?

08/20/19 1) Where the slab transitions from the thinner section, the bottom bars slope in a way that results in a vertical tension resultant.

Based on the damands in this area given in previous foundation calculations, the vertical tension resultant can be resisted by the

unused capacity of the bottom reinforcing that slopes upward from the deeper foundation section. See the comment folder for

supporting calculations.

2) See attached detail 10/S0412.

3) Clarification has been added to detail 2/S0411.Comment

81 09/10/19

Comment

Resolved


(04/22/19)

SAFE Model 05/24/19 Please confirm that given the specified reinforcement (and associated # of layers), the assigned cover to

centroid of rebar in the SAFE design preferences is adequate.

06/11/19 Confirmed. Within SAFE the specified clear cover under "Design > Design Preferences > Min. Cover Slabs" was increased to account

for the number of layers. See comment folder.Comment

82 07/29/19

Comment

Resolved


(04/22/19)

SAFE Model

Applied Loading

05/24/19 Please clarify/provide calcs for how the moments shown on Pg. 4 were converted to the distributed

sloping line loads representing the EQ loads in the SAFE model.

06/11/19 The conversion was done using a proprietary KPFF spreadsheet. The core section properties at the base are defined in the

spreadsheet, along with the bi-directional moments for each design point. The program then resolves these forces into the

statically equivalent line loads using the wall section cut properties, and outputs them in a format for input into SAFE. See comment

folder for supporting technical documentation.

Comment

83 07/29/19

Comment

Resolved

17 of 20

# Date of

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Date of

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Resolution/

Notes

Topic

& Reference

Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)84 Foundation Submittal

(04/22/19)

Foundation Design

Loading

05/24/19 Similar to Comment #57, please clarify/justify use of (1.2+0.12Sms) in lieu of (1.2+0.2Sms) for the dead

load factoring on the "Mat Design Acceptance Criteria" page.

06/11/19 Please see response to comment #57.

07/29/19

Comment

Resolved


(04/22/19)

Foundation Design

Loading

05/24/19 Please justify use of reference LATBSDC while use of "B" factors that vary from the recommended

values provided in App. B of that document.

06/11/19 Per LATBSDC section 3.6.3.2.1, there are three acceptable approaches for determining B:

1) B = 1.0

2) B = 0.9Rne/Rnem

3) B taken from Appendix B

We have used the second option to calculate appropriate values of B as presented in the basis of design. The calculations are

repeated in the comment folder.

Comment

85 07/29/19

Comment

Resolved


(04/22/19)

Foundation Design

Rebar Design

05/24/19 Please clarify the strategy behind averaging the results of multiple design strips in each direction from

the SAFE model in order to design the reinforcement. Does this potentially result in overstress at

individual design strips?

06/11/19 The mat foundation is expected to behave more as a rigid body than a series of discrete strips. For the area under the core the

loading and behavior is similar in each strip, and so these results have been averaged. The variation in peak rebar requirements in

these strips is +/- 7%, indicating the response is comparable in each strip and the approach does not overlook any extreme peaks.

Where the mat transitions to 6' thick the foundation response is driven by the columns instead of the core. Given the unique

loading condition design strips here have been treated independently and are not averaged with the core response.

07/29/19 Although the differences are not substantial, it is not clear that averaging of the strips is appropriate, as

it nulls the intent of the design strips (NERHP design guide also notes to check at critical sections and

has no reference to averaging of multiple strips). Presumably, the maximum flexure at one corner

comes from combined loading at an angle that does not necessarily engange the other end of the core

at the same time in the same flexure, which would indicate that averaging would not be appropriate.

Particularly for this core which is very long in one direction, it does not seem advisable to average the

results from one side of the core to the other, regardless of extremity of difference.

08/20/19 Mat flexure design has been evaluated for each strip independently, without averaging. Additional flexural reinforcing has been

added where required. See comment folder and S0200B and S0200T.

Comment

86 09/10/19

Comment

Resolved


(04/22/19)

Foundation Design

Rebar Design

05/24/19 For bottom bars in both directions, S0200B shows (2) layers @ 12", and (1) layer @ 6". However page

44 appears to indiciate that there should be (2) layers @ 6", and (1) layer @ 12". Please clarify and

confirm all other rebar callouts.

06/11/19 The rebar callouts on plan agree with the calculations. See comment folder for clarifying sketch.

Comment

87 07/29/19

Comment

Resolved


(04/22/19)

Foundation Design

Shear

05/24/19 Please provide a sample hand calc for the worst-case of the shear DCR's shown on Pg. 54 and 55 (also,

see #90)

06/11/19 See comment folder. Note that the bias factor has been included in calculations.

Comment

88 and 90 07/29/19

Comment

Resolved


(04/22/19)

Foundation Design

Shear

05/24/19 Please provide a punching shear check of the mat 06/11/19 Punching shear calculation has been added to the comment folder. Note that the strength provided by the vertical shear

reinforcing is conservatively ignored in this check.Comment

89 07/29/19

Comment

Resolved


(04/22/19)

Foundation Design

Shear

05/24/19 It appears that the shear calculations on Pg. 54 and 55 utilize the entire width of the slab to check one-

way shear. It may not be reasonable to utilize the entire width of the mat for this check, and it is

recommended to instead utilize an effective width of the core plus on mat thickness on either side

(refer to NERHP Design Guide 7 for similar recommendations).

06/11/19 For the region below the core, the design width has been limited to the width of the 10' thick area of mat when evaluating one-way

shear. This is equivalent to the design strips that were averaged for flexure (see comment 86). Compared to the NEHRP

recommendation of core width + 2*tmat, the design widths used are:

- N-S strips: 880" (v. 806" from NEHRP)

- E-W strips: 590" (v. 525" from NEHRP)

Though these widths are larger, the NEHRP commentary discusses that the limited width is a "conservative approach, perhaps

overly so." It is agreed that the entire mat width would be inappropriate, and we believe our design widths are reasonable in this

context.

For the 6' mat regions outside of this effective width, calculations have been provided to show the shear DCRs for each individual

strip. The average DCR was reported for reference only. See comment folder.

Comment

88 and 90

07/29/19 From the Comment 88 and 90 response PDF, it appears that in the N-S direction, although CSB1 is not

considered as part of the strip averaging, CSB6 is. Please justify inclusion of this strip within the

averaging, as it's location is more similar to CSB1 than the strips running underneath the core, and the

demands appear noticeably lower than CSB2 through 5. Does removing this strip from the averaging

change the results?

08/20/19 Mat shear design has been evaluated for each strip independently, without averaging. Additional shear reinforcing has been added

for strip CSB2 between grids C and C.7. See comment folder and S0200B.

Comment

90 09/10/19

Comment

Resolved

91 MCE Submittal #1

(06/04/19)

Revised Diaphragm

Calcs (L2)

06/18/19 1) On Pg. B246, it appears that the the length of the small segment of slab south of the core wall is

utilized to calculate the bearing capacity - please clarify this calculation. Given the thin strip of slab, it

would seem appropriate to not depend on bearing at this edge of the wall for diaphragm transfer at

both the west and east walls of the core (while at the north end, it would still seem reasonable to

consider).

2) Please clarify if general diaphragm shear checks are provided for the L2, considering the multiple

small slab widths resulting from the irregular shape/openings (and if any local chords are required

within these irregular areas)

07/03/19 1) Agreed. Calculations have been revised so that only the thin strip of slab south of the core transfers load through bearing. See

page 53 of the supplemental calculations. As noted, the load transferred in bearing is not limited on the north side of the wall.

2) See pages 62-63 of the supplemental calculations for local shear checks at critical locations in the level 2 diaphragm. See page 67

for a summary of the chord checks and pages 81-82 for the critical flexural sections and loading/moments.

07/29/19

Comment

Resolved

18 of 20

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Date of

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Notes

Topic

& Reference

Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)92 MCE Submittal #1

(06/04/19)

Revised Diaphragm

Calcs - General

06/18/19 Please justify that the embedment lengths of all collector rebar into the shear walls are adequate to

transfer the demands into/out of the walls. It is strongly suggested to carry all of, or at least a majority

of, the collector steel through the full length of the wall, for a few reasons

1) Ensures adequate embedment depth of the collection steel itself

2) Ensures that that tension collection reinforcement will reach the compression zone of the wall

(given that when the tension collector is engaged, it is only embeded into the tension zone of the wall)

3) Given flexure-shear interaction in the wall,which is currently not captured in the analysis, the

majority of the shear demands in the wall will be localized in the compression zone of the wall at one

edge of the wall. This requires a mechanism in which to "drag" the shear force back through the length

of the wall to be distributed to the shear friction dowels as well as the tension collector. The collection

rebar continuing through the length of the wall could serve this purpose.

07/03/19 Per the conference call on 6/24/19, we have taken the following approach:

a) At levels 2-7 the shear capacity of the wall was used to determine the required embedment length to develop the collector force

into the wall. Typically the length is small (less than 5') and the bar development length will govern. For typical cases a minimum 10'

embedment has been provided. In the supplemental calculation package see page 64 for a sample calculation and page 65 for a

summary of the required collector lengths.

b) At level 1 the rebar has been detailed to extend through the entire length of wall. Splices have been indicated at locations in the

wall web outside of the boundary zones to avoid congestion.

07/29/19 a) Method appears to refer to the previous NEHRP method which has been superceded in the latest

version of NEHRP - please reference the current version (Second Edition) Section 6.8 for the

recommended approach. As noted in this reference, if at least a portion of the collector steel does not

continue through the entirety of the wall: either reserve capacity of the horizontal rebar in the wall can

be used to splice with the collector steel to distribute the collector load through the length of the wall,

or supplemental bars can be added.

b) Closed

08/20/19 a) See response to comment #93. Collector steel along grids 3 and 4 has all been offset from the wall and run continuous for the full

length of wall. Dowels are provided to transfer the load into the wall, neglecting the length of the coupling beams.

For the walls parallel to grids C.7 and E.4, the collectors are continuous through the wall.

Comment

93 09/10/19

Comment

Resolved

93 Supplemental

Diaphragm Calcs

(07/03/19)

General

Collection/Load Path

07/29/19 (Also see comment #92 above for similar discussion)

As noted in the updated calculations, most of the diaphragm load paths to the core walls include

collection into the walls either through tension drag bars or compression bearing at the end (corner) of

the core in either direction.

In some cases (such as the southwest or northeast corners), there is a coupling beam that is located

very close to the corner of the core, such that the only load path to deliver the collected load to the full

length of the wall (as the rigid diaphragm analysis model assumes) is through the coupling beams.

For example, at L3, for the western edge of the south wall of the core: When the diaphragm load is

pushing to the west, the collector along Grid 4 is in tension. This collector rebar extends into the

second wall pier of the core. Is this load able to get resisted by just the 1st and 2nd wall pier, and not

the longer 3rd (easternmost) pier? Or is there a load path that delivers some remaining load to the 3rd

pier?

Similarly, when the diaphragm load is pushing to the east, compression bearing on the western corner

of this wall is depended upon to deliver the load from the diaphragm to the core. However, the only

load path for this collected compression out of the westernmost short pier is through the coupling

beams. Please justify that either:

a) The initial wall that the collected load enters can resist the collection demand, without distribution

to the remaining 2 piers along this line.

b) The collected load can travel through the coupling beam(s) and into the next pier(s) - however, the

capacity based shear of the coupling beam should then take into account the increased capacity due to

this compression.

c) Load path has an alterante rouote, and the demand in the collector can be transferred through other

means.

Please clarify/justify the above point for all similar collection along the south and north walls of the

core that includes coupling beams.

08/20/19 An alternate load path has been checked at these walls. In lieu of compression bearing a compression strut has been detailed

within the slab to collect this load. Shear dowels are then provided to transfer the load directly into the wall, excluding the length

of coupling beams. Moment across the interface has been checked due to eccentricity of the compression strut from the wall. It has

also been confirmed that the load transferred in shear friction does not exceed the ACI limits on Vn across the shear plane. See

comment folder.

Comment

93 09/10/19

Comment

Resolved

94 Supplemental

Diaphragm Calcs

(07/03/19)

Bearing Collection

Adjacent to Slab

Opening at Northeast

Core Corner

07/29/19 When bearing is utilized for collection at the northeast corner of the corner, there is an adjacent slab

edge that prohibits bearing against a majority of the thickness of the wall.

Although it appears that direct slab-to-wall connection length has been limited by the opening in the

slab (referencing Pg. 41 as an example of this location), the total bearing area appears to be multipled

by 2x, similar to other locations where an assumed expanded bearing area is projected into the slab. In

this case, the opening restricts this expansion of bearing area on one side - please clarify if this is

correctly accounted for in the determination of the bearing area.

08/20/19 Per response to comment #93, bearing is no longer used to deliver load to the walls on grids 3 and 4. Instead, a compression strut in

the slab has been evaluated for load transfer. Thus, of the limited slab width adjacent to this opening against the wall is no longer

considered.

Comment

93 09/10/19

Comment

Resolved

95 Supplemental

Diaphragm Calcs

(07/03/19)

L2 Diaphragm Calcs 07/29/19 There is a pinch point in the L2 diaphragm between the small slab opening adjacent northeast corner of

the core and the slab "cut out" directly east of this at Lines F->G.

From the northeast corner of the small slab opening, to the southwest corner of the slab "cut out", only

a small width of slab exists (~8'4").

Is there any special consideration for this "pinch" point of the slab, and was slab shear checked across

this small width? Compared to typical diaphragm shear type behavior, this area would appear to act in

more of a collection (or strut-and-tie) type fashion to deliver load through this thin section of slab,

however no additional rebar is currently specified. Please justify.

08/20/19 Diagonal bars have been added across the "pinch point" to transfer collector forces. Compressive stresses in the slab have been

checked to verify confinement is not required. Previously provided chord steel accounts for flexure here, but the added diagonal

bars will assist in restraining cracks. See comment folder.

Comment

95 09/10/19

Comment

Resolved

19 of 20

# Date of

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Date of

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Resolution/

Notes

Topic

& Reference

Submittal: Basis of Design (05/29/18; 07/16/18; 08/14/18; 12/21/19)96 Current Drawings

(07/03/19)

Core Wall @ L2 07/29/19 As shown on S0212 and 1/S0303, L2 has two openings in the western wall of the core that are

otherwise not present throughout the height of this wall:

1) These openings do not appear to be present in the current Perform model. Please update/clarify.

Did the associated shear calcs for this wall similarly not represent the shortened length at this location?

2) Please clarify how PM design was handled for this pier at L2?

3) Please justify the boundary zone confinement detailing at this wall. It appears as though the

openings are full height, and would interrupt the portion of boundary zone that extend into the

western wall at the corners - please justify that the resulting confined zones at the left side of the core

(both in the flanges and the web) are still acceptable.

08/20/19 There are no openings at these locations, but a drafting error caused a portion of the wall to disappear in plan and elevation. See

the revised sheets S0212 and S0303 in the comment folder for the correct representation.

Comment

96 09/10/19

Comment

Resolved

20 of 20

2208 4th Ave - Structural Design Peer Review Comment Log9/10/2019

Reviewers: Engineer of Record:

Wood (Marshall Lew) GeoEngineers

Documents Received:

1) Geotechnical Master Use Permit Report, by GeoEngineers, dated 9/26/17

2) Downhole Geophysical Tests, submitted by GeoEngineers

3) Memorandum - Site-Specific Ground Motion Procedures, by GeoEngineers, dated 7/13/18

4) Memorandum - Site-Specific Ground Motion Procedures, by GeoEngineers, Rev. 1, dated 8/6/18

5) Memorandum - Site-Specific MCER and SLE Response Spectra, by GeoEngineers, dated 8/27/18

6) Memorandum - Site-Specific MCER and SLE Response Spectra-Rev. 2, by GeoEngineers, dated 10/9/18

7) Design Memorandum 2 - Site Specific Conditional Mean Spectra, by GeoEngineers, dated 1/8/19

8) Design Memorandum 3 - Ground Motions for Nonlinear Response History Analysis, by GeoEngineers, dated 3/1/19

9) E-mail from GeoEngineers, dated 4/3/19, Long Period CMS ground motions (Comments 26-28)

10) Draft Report of Seismic Design Services, by GeoEngineers, dated 5/20/19

11) Report of Seismic Design Services, by GeoEngineers, dated 7/22/19

# Date of

Comment


Response


Exhibit

Date of

Resolution

Resolution

/

Notes

1 Report 9/26/17 MUP 07/23/18 From a Geotechnical Performance-Based Design perspective, the MUP report is

acceptable. The report was not reviewed for static or gravity design considerations,

nor for Code-related design.

XX/XX/XX Response not required 7/23/2018 Comment

Resolved

2 Downhole

Geophysical Tests

07/23/18 GeoEngineers has submitted the results of Downhole Geophysical Tests performed

by Shannon & Wilson for the SR99 Bored Tunnel Project to be used for the 2208 4th

Avenue analyses. The application of these results to the subject site is acceptable

and appropriate.


Resolved

3 Memorandum

7/13/18

Topic / (Pg) 07/23/18 The memorandum presents the proposed approach for evaluating site-specific

response design spectra and ground motions for the project. The approach is based

on 2015 IBC, ASCE 7-10, and PEER TBI 2.03. The general procedure described in the

Memorandum is acceptable. However, the proposed procedure does not include the

determination of the Service Level Earthquake ground motion spectrum.

XX/XX/XX Forthcoming Design Team Response 8/13/2018 Comment

Resolved

4 Memorandum

7/13/18

Cascadia Subduction

Zone Ground Motion

Model

07/23/18 It was suggested in the July 16 Kick-Off meeting that the updated 2018 BC Hydro

GMM be used instead of the 2012 BC Hydro GMM. Peer Review received e-mail

from GeoEngineers on 7/18/18 indicating that the 2018 BC Hydro GMM will be used

and the weights assigned to the Subduction GMMs will be reviewed and revised in a

supplemental submittal.


Resolved

5 Memorandum

7/13/18

Service Level

Earthquake (SLE)

07/23/18 On 7/18/18, GeoEngineers indicated by e-mail that the SLE ground motion

procedures will be added to the memorandum.


Resolved

6 Memorandum

8/6/18

Cascadia Subduction

Zone Ground Motion

Model

08/13/18 The inclusion of the 2018 BC Hydro GMM is acceptable. XX/XX/XX Response not required 8/13/2018 Comment

Resolved

Topic

& Reference

Submittal: Geotechnical Master Use Permit Report

Submittal: Downhole Geophysical Tests

Submittal: Memorandum - Site-Specific Ground Motion Procedures, by GeoEngineers, dated 7/13/18

Submittal: Memorandum - Site-Specific Ground Motion Procedures, by GeoEngineers, Rev. 1, dated 8/6/18

7 Memorandum

8/6/18

Service Level

Earthquake (SLE)

08/13/18 The SLE procedures are acceptable. In addition, the damping is appropriate for

building height of 295 ft in accordance with TBI 2.03.


Resolved

8 Memorandum

8/6/18

Site-Specific

Conditional Mean

Spectra

08/13/18 The proposed procedure to develop the site-specific CMS spectra is acceptable. XX/XX/XX Response not required 8/13/2018 Comment

Resolved

9 Memorandum

8/6/18

Ground Motion

Selection and

Modification

08/13/18 The proposed procedure to select and modify ground motions is acceptable. XX/XX/XX Response not required 8/13/2018 Comment

Resolved

10 Memorandum

8/27/18

Sedimentary Basin

Effects

09/04/18 In this memorandum, three Basin Amplification Factors are used: (1) Crustal using

CB14 NGA-West2 GMM; (2) CSZ Intraslab; (3) CSZ Interface. In the recent 9th/John

Project, GeoEngineers used four Basin Amplication Factors; excluded in this project

was the equally-weighted ASK14, BSSA14 and CY14 using Z1.0. Please justify the

change the exclusion of the 4th Basin Amplification Factor for this project.

09/18/18 During the March 22, 2018 SDCI/USGS basin amplification workshop, Z1.0

basin amplification models were discussed as not credible models for

computing amp factors for the Seattle basin. This conclusion came from

the developers of the models in addition to the other particpants of the

workshop. ASK14, BSSA14, and CY14 Z1.0 basin amplification models were

developed for the Los Angeles basin, which is much lower Z1.0 gradient

than the Seattle basin. Recommendations by the developers were to drop

their Z1.0 amp models in favor of keeping the CB Z2.5 model. The USGS

report documenting the workshop discussions is in-press.

9/19/2018 Comment

Resolved

11 Memorandum

8/27/18

Sedimentary Basin

Effects

09/05/18 CSZ Intraslab - CTR, MAR and UNK Stations Basin Amplification Factor values differ

from the values given for the 9th/John Project for the period range from 0 to 1

second. Table 1-2 has footnote indicating that values were linearly interpolated to

1.0 at T=0 seconds; for 9th/John, the amplification factor was kept constant from 0

to 1 second period. Please explain this discrepency.

09/18/18 The recorded CSZ intraslab basin factors are applicable for about T>0.5 sec

as shown in Figure 1-4 of our memorandum dated 8/27/18. During the 9th

and John project, we made the conservative assumption that the CSZ

intraslab basin factors are constant for T<1 sec using the value at T=1 sec;

however, during the March 22, 2018 SDCI/USGS basin amplification

workshop, the recommended CSZ interface basin amplification factors

were linearly extrapolated to achieve a basin factor of 1.0 at PGA (T=0

seconds). To be consistent with the CSZ-interface factors, we similarly

extrapolated the intraslab basin factor to 1.0 at PGA from the T=0.5 value.

This benefited our approach by removing the discrepencies in basin factors

at short periods between both source-types.

9/19/2018 Comment

Resolved

12 Memorandum

8/27/18

Table 1-5

Modification factor

for embedment

09/05/18 Please provide the calculations for the Modification Factor for Embedment. 09/18/18 Please refer to the supplemental calculations submitted with this comment

log.

9/19/2018 Comment

Resolved

09/05/18 Please provide a published reference for the single-station sigma or a

communication from Norm Abrahamson confirming the 0.6 value.

09/18/18 Please refer to Figure 3-49 from the 2011 BC Hydro Probabilistic Seismic

Hazard Assessment Report (Report No. E658 - Vol. 3, February 2011)

included with this comment log.

9/19/2018 Comment

Resolved13 Memorandum

8/27/18

Single-Station Sigma

Submittal: Memorandum - Site-Specific MCER and SLE Response Spectra, by GeoEngineers, dated 8/27/18

09/20/18 The concept of single-station sigma for this project is indeed intriguing. The

consultant states that “incorporation of site-specific site effects… provides a more

informed ground motion estimate than an ergodic model was derived under.” In this

case, it is argued that the incorporation of site-specific basin amplification factors

provides the justification for the use of the single-station sigma, thus moving towards

a more “non-ergodic” approach. The consultant states “…by using site-specific basin

AFs from recorded CSZ intraslab earthquakes and the M9 CSZ interface simulations

into the ground motion design, we are providing a more informed ground motion

estimate at our site and consequently expect less variability in the ground motion

than what would be predicted using the ergodic GMM.”

While there may be less variability using the proposed approach, the total process of

predicting ground motion at the site is still essentially an ergodic process. The

GMMs are still ergodic models based on global models and not regional models. The

CSZ intraslab basin factors are based on recorded ground motions from only two

events at three stations in downtown Seattle; these events were not large

magnitude events that could occur on the CSZ. The 3D simulations of the M9 CSZ

interface events (Fig. 1-5 of Memo) still show variability in the amplifications. Also,

given that this project is still using one of the legacy methods of ground motion

prediction just prior to the adoption of new guidelines which will be in effect soon, it

is not prudent to adopt the use of single-station sigma for this project.

1/10/2019 Comment

Resolved as

the use of

single-station

sigma will not

be used.

14 Memorandum

10/9/18

Site-Specific MCE R

Spectrum

01/10/19 The site-specific MCER spectrum is acceptable. The spectrum with KSSI effects is also

acceptable.


Resolved

15 Memorandum

10/9/18

Site-Specific SLE

Spectrum

01/10/19 The site-specific SLE spectrum is acceptable. XX/XX/XX Response not required 1/10/2019 Comment

Resolved

16 Conditional Mean

Spectrum for long-

period response

Page 2, Table 2-1 01/10/19 The conditioning periods for the long-period CMS is acceptable. XX/XX/XX Response not required 1/10/2019 Comment

Resolved

17 Conditional Mean

Spectrum for short-

period response

Page 2, Table 2-1 01/10/19 The conditioning periods for the short-period CMS is acceptable. XX/XX/XX Response not required 1/10/2019 Comment

Resolved

18 Seismic Hazard

Deaggregation

Pages 3-4 01/10/19 The deaggregation analyses are acceptable. XX/XX/XX Response not required 1/10/2019 Comment

Resolved

19 Conditional Mean

Spectra

Pages 6-7 01/10/19 The CMS are acceptable. XX/XX/XX Response not required 1/10/2019 Comment

Resolved

20 Conditional Mean

Spectrum with KSSI

Pages 5-6 01/10/19 The CMS with KSSI effects are acceptable. XX/XX/XX Response not required 1/10/2019 Comment

Resolved

21 Ground Motion

Selection and

Modification

Page 8, Tables 2-8

and 2-9

01/10/19 The text refers to these tables as Tables 2-7 and 2-8. Please correct. XX/XX/XX Response not required 1/10/2019 Comment

Resolved

22 Ground Motion

Selection and

Modification

Page 8, Tables 2-8

and 2-9

01/10/19 The recommended number of ground motions corresponding to source type are

acceptable.


Resolved

Submittal: Memorandum - Site-Specific MCER and SLE Response Spectra-Rev. 2, by GeoEngineers, dated 10/9/18

Submittal: Design Memorandum 2 - Site Specific Conditional Mean Spectra, by GeoEngineers, dated 1/8/19

23 Significant duration

D5-75

Page 9 01/10/19 Peer Review is unaware of a requirement in LATBSDC for the duration as defined in

the Memorandum. In past peer reviews, we have asked for D5-95 as a duration

measure.

02/15/19 Both D5-75 and D5-95 are measures of significant duration. However, the

Brookhaven and Walling et al. significant duration models that are

proposed for this evaluation were derived using the D5-75 duration metric

and afterward transformed to the D5-95 metric. By using the D5-75 metric,

we are being consistent with how the duration models were derived and

are also measuring the duration of the most significant (highest energy)

shaking.

2/18/2019 Comment

Resolved

24 Ground Motion

Selection and

Modification

Pages 7-9 01/10/19 The procedure for ground motion selection, in general, is acceptable. The GEOR may

submit candidate seed time histories for review.

XX/XX/XX 2/18/2019 Comment

Resolved

25 Short-Period

Ground Motions

Table 3-4 03/04/19 Ground Motions S-1 to S-11 are acceptable. XX/XX/XX Response not required 3/4/2019 Comment

Resolved

26 Long -Period Ground

Motions

Table 3-5 03/04/19 Ground Motion L-1 has a durations (D5-75) of 4.1 and 4.6 seconds for the H1 and H2

directions. For the long period CMS, these durations appear to be insufficient for the

intended analyses.

XX/XX/XX Will you please clarify why ground motions L-1, L-3, and L-5 are insufficient

for the intended analysis?

The durations are within a plus and minus standard deviation of the

median value, which is the range in durations that we aim for when

selecting ground motions. This recognizes that there is variability in the

significant duration value for a given earthquake scenario.

We’d like to note that the L-1 record was used previously during the 9th

and John project. For that project the significant duration parameter was

D5-95, and this record exceeded the median value by about 7 sec. The

relatively low D5-75 and a relatively high D5-95 suggest that this ground

motion is compact in time when expending 75% of the energy, but takes

longer than average to expend 95% of the energy.

03/26/19 This is the SPRP response to the Response from the Design Team for Comments 26 to

28. According to Table 2-1 of the 1/8/19 GeoEngineers memorandum, the

fundamental periods of the building are 5.33 and 3.14 sec. The durations of Ground

Motions L-1, L-3, and L-5 do not appear to be sufficiently long enough to fully excite

the building. In addition, L-1, L-3 and L-5 all have D5-75 durations that are below the

expected median of 7-8 sec for the Seattle Fault and the Gridded Background

Seismicity. The D5-75 durations should at be equal or greater than the median

shown in Table 2-10 of the 1/1/19 GeoEngineers memorandum for both the H1 and

H2 components for the long period CMS.

4/3/2019 Comment

Resolved


Motions

Table 3-5 03/04/19 Ground Motion L-3 has durations (D5-75) of 6.2 and 3.8 seconds for the H1 and H2


intended analyses.

XX/XX/XX See comment 26 4/3/2019 Comment

Resolved


Motions

Table 3-5 03/04/19 Ground Motion L-5 has durations (D5-75) of 5.2 and 5.0 seconds for the H1 and H2


intended analyses.

XX/XX/XX See comment 26 4/3/2019 Comment

Resolved


Motions

Table 3-5 03/04/19 Ground Motions L-2, L-4, L-6 through L-11 are acceptable. XX/XX/XX Response not required 3/4/2019 Comment

Resolved

30 Documents received 4/3/l9 (1) Ground Motion Durations and Calculated Periods; (2) Long Period Ground Motion

Husid Plots


Resolved

Submittal: E-mail from GeoEngineers, dated 4/3/19, Long Period CMS ground motions (Comments 26-28)

Submittal: Design Memorandum 3 - Ground Motions for Nonlinear Response History Analysis, by GeoEngineers, dated 3/1/19

31 Ground Motion

Durations

04/08/19 D5-95 Durations are provided for the ground motions in addition to the D5-75

durations. The D5-95 durations are satisfactory in demonstrating the suitability of

Ground Motions L1, L3 and L5 for the Long Period CMS.


Resolved

32 Ground Motion Husid

Plots

04/08/19 The Husid Plots are acceptable; Ground Motions L1, L3 and L5 are acceptable. XX/XX/XX Response not required 4/8/2019 Comment

Resolved

33 Scaled Time Histories

for analysis

05/28/19 The scaled time histories are acceptable. XX/XX/XX Response not required 5/28/2019 Comment

Resolved

34 Scaled Time Histories

for analysis

05/28/19 Some of the scaled time histories have long duration. Also the digital sampling rate

was not stated. Please clarify if there will need to be any modification or processing

of the time histories prior to the nonlinear response history analysis (i.e., increasing

the time step or shortening of the time histories).

06/18/19 We provided the ground motion data files to KPFF with a 0.02 sec time

step. No additional modification such as truncating the ground motions is

required. For the longer subduction zone records, KPFF ran the full length

of the acceleration records in PERFORM and the length of the analysis was

not curtailed to shorten the analysis time.

6/19/2019 Response is

acceptable.

Comment

Resolved.

35 Report 07/29/19 The report is acceptable. XX/XX/XX Response not required 7/29/2019 Comment

Resolved

Submittal: Draft Report of Seismic Design Services, by GeoEngineers, dated 5/20/19

Submittal: Report of Seismic Design Services, by GeoEngineers, dated 7/22/19

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