Hotel Churchill: Using Dampers Strategically

2017 SEAOC CONVENTION PROCEEDINGS

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Hotel Churchill: Using Dampers Strategically

Eric T. Lehmkuhl, Principal/Aaron Pebley, Associate/Shaun Walters, Project

Manager KPFF Consulting Engineers

San Diego, CA

Abstract Hotel Churchill, a historic unreinforced masonry infill building, was slated to be remodeled. Built in 1914 as a hotel, the building would be re-purposed into affordable housing. Interestingly, the required renovations did not trigger a code reanalysis and retrofit. Working with the architect and owner, the structural engineers developed a plan to mitigate the most egregious deficiency, the soft story at the street (first) level, within the project budget. Interesting aspects of the project:

Voluntary retrofit

Performance Based Design and full peer review

Mass reduction as part of the retrofit strategy

Modeling of unreinforced masonry infill walls

Nonlinear Response History Analysis

Point nonlinear model to simplify the analysis

Architect chose to expose the dampers

FRP strengthening of diaphragms and unreinforced concrete columns

Introduction Built in 1914, the historic Hotel Churchill was constructed as the City of San Diego prepared to host a large influx of tourists from around the world for the Panama-California Exposition of 1915. Over the subsequent century of use by a variety of different owners, it was even converted into a medieval-themed hotel. The aging building got a chance at new life when it was acquired by the San Diego Housing Commission in 2013 with the intent to renovate it to provide sorely needed affordable housing to the downtown San Diego community.

Preliminary Investigation When KPFF was brought onboard to support the extensive renovation and rehabilitation design proposed by Studio E Architects, the first task was a comprehensive evaluation of the building’s structure to determine if any significant deficiencies existed. Record structural drawings were limited. The drawings available in the City of San Diego’s records consisted of only three sheets: basement/foundation plan, 1st floor framing plan, and a sheet with sections and details through the basement and 1st floor.


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With the drawings only providing a portion of the information needed regarding the structural system and no material strength information whatsoever, it was evident that a comprehensive as-builting and material testing program would be required. Once KPFF had developed the necessary program, Ninyo and Moore was hired to perform the investigations. The most significant determinations of the findings were as follows:

Gravity System – Cast-in-place (CIP) concrete joists supported by CIP beams and columns with a 3” concrete slab.

Lateral System – On the north and east sides of the

building was a CIP concrete shearwall with many window openings. The shearwalls occurred only from Level 2 to Level 6, which was the original roof. The walls did not extend to grade on these sides. The west and south sides of the building were 8” thick hollow clay tile infilled concrete frames. These infilled frames did run from grade all the way to Level 6. See Figure 1.

Level 7 – The original construction in 1914 was only

six stories. In the early 1930’s an additional floor was added. Selective demolition at this story determined that the entire story was constructed of unreinforced masonry (URM) brick bearing walls with a steel framed roof. In some cases the URM walls had been directly stacked on top of the original concrete parapet with no positive attachment.

First floor Columns – Given the discontinuity of the

concrete shearwalls on the north and east sides of the building, of particular interest was the ductility of the concrete columns at this floor. Review of the as-builts and non-destructive evaluations determined that, in general, the columns were not particularly well-reinforced longitudinally; but they did have a reasonable amount of confinement for a building of this age. See Figure 2.

Figure 1. Plan of Building

Figure 2. Typical Column Of particular concern, however, was that the corner columns of the building were specifically identified in the record drawings as being completely unreinforced. Scanning of these columns confirmed our concerns that in fact these columns had been intentionally unreinforced for as yet undetermined reasons. See Figure 3.

Figure 3. Corner Column In addition to the reinforcement issues of the columns, the drawings indicated poor development of the reinforced columns down into the larger basement columns below. Without wanting to damage the columns in the process, there was no way to confirm if better means of development had been built, so we were forced to rely on the record drawings.


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Site Seismology and Hazard The main sources of seismicity at the project site are the Rose Canyon fault that runs north-south through downtown San Diego and the Coronado Bank/Palos Verdes fault that is about 12 miles offshore. The team determined that a reduced hazard level (as compared to a new building) was appropriate for this project. The performance objective used was Collapse Prevention (CP) under a Basic Safety Earthquake-2 for existing buildings (BSE-2E). BSE-2E is a seismic event with a 5% in 50 year return period as compared to a BSE-2N which is a 2% in 50 year return period. The suffix “E” denotes a reduced hazard that is commonly accepted for existing buildings. ASCE 41-13 Chapter 2 commentary was a resource that provided justification of the reduced hazard level for this structure. Initial Analysis – Linear Model In order to get an initial “feel” for the building’s behavior a linear elastic model was created in ETABS. Model Highlights Much of the building lateral system is Unreinforced Masonry Infill Frames. We chose to model the infill according to the criteria specified in ASCE 41. This consists of modeling equivalent struts to represent the actual compression strut that develops in the infill at low drifts. See Figure 4. The 8” clay infills produced equivalent strut width ranging from 12” to 24” according to this criteria. The two limit states of concern were out-of-plane failure of the infill and in-plane compression strut failure by crushing, buckling, or concrete column shear failure. The first limit state, out-of-plane failure, was addressed with two main criteria:

Follow h/t limits in ASCE 41 and arching theory

Do not increase accelerations on the building

Figure 4. (from p. 11-53 ASCE 41) The second limit state was addressed by limiting the drift of infill frames to around 0.5%. This drift is commonly seen as the limit for most infill frames (Riahi, p.652), (ASCE41 Table 11-9), (Shing, p. 10). The first floor has only concrete columns and beams on two sides. Ignoring reinforcing detailing problems, the concrete columns and beams were modeled as frames. Results A deflected shape of the model is shown in Figure 5. This figure clearly shows that this building behaves as a rigid box on top of a soft story. Assuming that the concrete columns and beams hold together, the drift at the first floor is in the range of 5% for the design level earthquake noted above. Clearly the first story is a collapse threat. Interestingly, the soft first floor elongated the period of the building and seemingly protects the brittle infill walls in the upper floors from excessive acceleration in small ground motions.


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Figure 5. 3D Model The linear analysis confirmed the need to mitigate the soft story. Since a retrofit was not code-mandated, KPFF had to convince the team that this was a necessary expense. Fortunately, the design and ownership team quickly understood the issues and agreed to spend money on addressing them. It was then up to the design team to find the best strategy with regard to aesthetic, function, cost, and life safety. Conventional Retrofit Strategies Considered Initially investigated were conventional retrofits for mitigating the weak/soft story, specifically, steel braced frames and concrete shearwalls. While these systems certainly mitigated the large displacements at the 1st floor and therefore the weak/soft story effects on the columns, they increased the stiffness of the 1st floor. This stiffness increase had the undesired effect of decreasing the building’s natural period and thus increasing the overall building base shear. Our evaluation determined that the upper floor concrete shearwalls and clay tile infill frames were overstressed with the increase in seismic loads. Essentially, retrofitting the 1st floor with these conventional methods would result in additional retrofit in the upper floors.

Given how disruptive braces or shearwalls would be in the upper residential floors, and the significant added cost in retrofitting more than just the 1st floor and basement, it was clear that alternate solutions needed to be investigated. The beneficial effect of added damping without much added stiffness became apparent. Figure 6 illustrates conceptually the reduced response of the damped structure versus the stiffened structure.

Figure 6. Effect of Added Damping Selected Retrofit Strategy The chosen retrofit strategy consists of the following elements. Rebuild Top Floor The entire 7th floor was demolished and rebuilt. The new roof system consists of metal deck on structural steel. Cold formed stud walls with sureboard provide lateral resistance. Mass Reduction In addition to removing the brick walls on the 7th level, all the hollow clay tile partition walls in the building were removed and replaced with cold formed studs. This significantly reduced the mass of the building and, therefore, the seismic loads. Dampers at First Floor Four viscous dampers, one at each side of the building, were used to mitigate the collapse potential of the first story and not increase accelerations in the upper floors. In order to damp the bottom floor and protect the URM, the URM infill walls were disconnected from the concrete frame. Other Items New drag plates were installed at the damper frame lines. Vulnerable concrete columns were wrapped with FRP to increase axial and flexural capacities. Diaphragms were strengthened with FRP.


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Nonlinear Model Time History Development The Rose Canyon and Coronado Banks are both strike-slip fault types that are capable of producing an estimated 6.9 and 7.7 magnitude earthquake, respectively. The seismic sources for the time histories were selected by Leighton and Associates, the project geotechnical engineer, and are summarized below:

Name Country Magnitude Distance

from Source

Fault Mechanism

95 Kobe

Japan 6.9 1 km Strike-Slip

99 Duzce

Turkey 7.1 4 km Strike-Slip

99 Hector Mine

USA 7.1 31 km Strike-Slip

Table 1. Seismic Records Record Scaling The seismic records were scaled to the target spectrum in the time domain using spectral matching for periods ranging between 0.8 seconds and 2.4 seconds. All of the record scaling was performed by the project geotechnical engineer. Analysis Approach This project implemented a Nonlinear Response History Analysis (NLRHA), and the design was based on the maximum demands of an envelope of the three earthquake records described previously. In order to simplify the analysis, only the fluid viscous dampers (FVD) were modeled as nonlinear elements. In ASCE 41 any element with a demand to capacity ratio (DCR) less than 150% is defined as essentially linear and is permitted to be modeled as linear. The 2nd floor and above underwent very low drifts and accelerations and the structural elements were well below their elastic limits. The existing concrete columns in the first floor are expected to undergo large drifts and inelastic deformations under the seismic demands evaluated. Rather than perform rigorous testing to determine capacity and ductility or use very low values from ASCE 41, we chose to ignore the contribution of these columns to the lateral system for strength, ductility, and stiffness. While the existing columns will provide some plastic damping benefit, it would have been difficult and costly to quantify and ignoring it greatly simplified the analysis. In lieu of explicitly modeling these columns accurately, it was decided that they would be modeled as “pin-pin” columns and compatibility analyses would be performed to ensure concrete crushing and shear

failure are not likely to occur. In certain locations critical and unreinforced columns were retrofitted which is discussed in the sections that follow. The existing Unreinforced Masonry (URM) infill walls on the first floor were also excluded from the analysis model as they were isolated from the structure by adding separation joints adjacent to the existing concrete frames. These simplified assumptions of limiting the nonlinear element only to the dampers enabled us to utilize ETABS and implement their Nonlinear Modal Analysis also known as Fast Nonlinear Analysis (FNA) solver.

Figure 7. Model Overview

Figure 8. Damper Elements By utilizing the FNA solver versus using the direct integration solver in ETABS or performing the analyses using more robust programs such as Perform3D, we were able to limit run times from about a day to just under five minutes. Using ETABS also enabled quick changes and simple exporting of data for post-processing. ETABS also performed the steel


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design and allowed for shorter time between iterations, allowing us to optimize frame and damper size, along with implementing studies required by the peer review team. Analysis Highlights

Nonlinear Response History Analysis (NLRHA) Maximum demands from suite of three records Nonlinear Modal “Fast Nonlinear Analysis” 5% constant modal damping in addition to the

explicitly accounted for fluid viscous damping Expected material properties used throughout Cracked stiffness modifiers used for all linear

elements Analysis Runs: Main Analysis, Accidental Torsion,

Upper Bound (UB) and Lower Bound (LB) damper properties, UB and LB in-plane diaphragm stiffness, UB and LB out-of-plane diaphragm stiffness

Nonlinear Damper Modeling The Fluid Viscous Dampers (FVD) were explicitly modeled using damper links in ETABS. The dampers were modeled in series with an HSS driver brace matching the actual built condition. The list below summarizes important parameters that were used in our analysis and are recommended when performing NLRHA with a structure containing Fluid Viscous Dampers:

Ritz vector method of modal analysis implemented 500 modes obtained and both vertical and lateral

modes considered Damper mass explicitly modeled to engage dampers

in the FNA Lateral mass was not lumped at the story levels in the

analysis settings 5% modal damping to account for nonlinear actions

not explicitly modeled Four items are needed to specify a damper: Maximum Force, Maximum Stroke, Damping Coefficient (C), and Velocity Exponent (α). The Hotel Churchill project damper specification is summarized in the following table:

Damper

Force Capacity

Stroke Capacity

Damping Coefficient

(C)

Velocity Exponent

(α)

440 kip 5” 160 kip-sec/inch

0.30

Table 2. Damper Specification Analysis and Design Results As previously discussed, the main intention of the design was to significantly reduce the drift of the first floor and provide a ductile lateral system for the non-ductile frame columns to “lean” on. A secondary goal was to maintain the original global behavior of the structure in order to prevent an increase of forces and accelerations in the upper floors. The following sections provide an overview of the analysis results and comparisons with the original structure. Global Building Results The plots provided below show that the story shears, overturning moments, and accelerations were reduced or matched when compared to the original structure. The weak/soft story of the original structure may not have survived a moderate to low seismic event, but it did limit the magnitude of the forces and accelerations to the upper floors. As part of a voluntary retrofit approach this overall behavior was required to be maintained. Altering the existing behavior of the structure would have increased forces in all levels and pushed the project into a mandatory retrofit, and this would have greatly increased construction costs.

Figure 9. Comparison of Story Shears


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Figure 10. Comparison of Overturning Moments

Figure 11. Comparison of Story Accelerations In the figures below it can be seen that the global drift at the 1st floor was significantly reduced and the drift in the floors above were reduced. As noted previously, the brick framed 7th floor was replaced with a more flexible, but significantly more ductile light-gauge shear wall system.

Figure 12. Comparison of Story Drifts

Figure 13. Comparison of Story Displacements This reduction in drift was accommodated without significantly stiffening the structure or altering the natural period as shown in the figure that follows comparing the fundamental periods. Tuning the building behavior and reducing forces and drifts would not be possible without increasing the structural damping and reducing the seismic mass.

Figure 14. Comparison of Building Periods By inspecting the energy balance curves we were able to obtain approximate global damping information. As stated prior, we set the modal damping at 5% to account for plastic damping of nonlinear elements not explicitly modeled and for miscellaneous damping of non-structural elements. The main item that would contribute to the plastic damping that was not


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included in the model as nonlinear elements are the concrete frame columns at the first floor. There is low likelihood of plastic damping in the upper floors since the drifts are low and the elements are largely going to remain linear. It was decided among the team that 5% constant modal damping was a reasonably accurate value for this structure. The normalized energy balance plot below for the Kobe analysis shows that the nonlinear viscous damping accounts for approximately 80% of the energy dissipation and is on the order of 4x the modal damping. The Duzce and Hector Mine analysis results provided similar energy ratios.

Figure 15. Normalized Energy Dissipation Plot – Kobe Record Column Results A concern for this project was unreinforced columns exposed to large drifts and in some cases uplift. In order to keep close track of each column, the axial loads were recorded and plotted at every time step. Data was exported from ETABS and post processed in excel to quickly evaluate the peak uplift and compression for a specific column for the suite of earthquake records. The following is an example of the data that was processed for column ID-1.

Figure 16. Column Axial Force vs. Time While many of the corner columns underwent large reversals in axial load and were exposed to uplift, a majority of columns remained in a state of compression, contained adequate confinement and longitudinal reinforcing, and did not require retrofitting. A summary of the columns that required retrofit is provided below. It should be noted that all columns that were either unreinforced or underwent uplift loads were retrofitted.

Figure 17. Column Key Plan


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Col ID

Col size

Long. Reinf # (Bar Size)

Spiral Ties

Pu (kip)

Tu (kip)

Retrofit*

1 24x30 NP NP 355 -136 C&U 4 18x18 4 (3/8”) NP 183 -93 C&U

5 18x18 4 (3/8”) NP 171 -29 C&U

12 22x22 8 (1”) 1.5” 273 -11 U 20 24x30 NP NP 471 - C 26 30x30 NP NP 630 -121 C&U

*C=FRP Confinement Retrofit, U=FRP Uplift Retrofit, NP=Not Present Table 3. Column Retrofit Summary Damper Results Similar to the column results, damper data was obtained and plotted. Damper force and damper stroke or displacement were plotted as shown below for each damper and a summary of all damper results is provided in the following table.

Figure 18. Damper D-01 Force vs. Time

Figure 19. Damper D-01 Stroke vs. Time

Figure 20. Damper Key Plan

Damper ID

Maximum |Force|

KOBE DUZCE HECTOR MAX

D-01 340 kip 293 kip 316 kip 340 kip

D-03 324 kip 326 kip 322 kip 326 kip

D-A1 322 kip 304 kip 311 kip 322 kip

D-G1 273 kip 306 kip 288 kip 306 kip

Damper ID

Maximum |Stroke|

KOBE DUZCE HECTOR MAX

D-01 1.77'' 1.21'' 2.23'' 2.23''

D-03 1.68'' 1.20” 2.27'' 2.27''

D-A1 2.14” 1.92” 1.91” 2.14”

D-G1 1.02” 1.47” 0.88” 1.47” Table 4. Damper Analysis Results Steel Frame Design and Detailing Implementing ETABS for the nonlinear analysis had the added convenience of being able to directly design the steel moment frames. As previously stated, ASCE 41 defines any element with a demand-to-capacity ratio (DCR) less than 1.5 to be essentially linear. The approach for the steel moment frames was to detail them as ductile as possible while designing them to remain essentially linear. This task was difficult because of the large drifts at the 1st floor. It is logical to deduce that a frame with a DCR of 150% is a failing frame. However, a more appropriate perspective is that the frames are designed to only accommodate a ductility demand of 1.5 since the analyses used actual seismic demands which were not reduced by response modification factors. The inclusion of the damper frame within the moment frame eliminated the possibility of following any specific moment frame detailing specified in the AISC Seismic Provisions. Due to the low ductility demand, it was decided to detail the frame as an ordinary moment frame with special detailing where possible. The damper driver was designed to accommodate the maximum force that the damper could deliver.


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Figure 21. ETABS Design North Frame / DG-1 Foundation Analysis and Design As part of this project a new shear wall and foundation were provided to anchor the damper-frames. The shear wall design was performed directly in ETABS using the BSE-2E demands. The foundation design was performed in SAFE. The design forces were obtained by determining the critical time step for each record and for each direction at which the maximum overturning moment occurred. This foundation design approach, which we denoted as a “snap-shot” analysis, results in four (Mxmax, Mxmin, Mzmax, Mzmin) load cases for each time history record. If sliding shears were of concern then four more cases would occur for each record, since the maximum shear and maximum moment may not coincide. The reactions associated with each “snap-shot” were exported to SAFE, and the foundation design proceeded similar to a linear or dynamic analysis. Excerpts from the foundation calculation are provided below.

Figure 22. Excerpt From Foundation Calculations

Bounding Results In addition to performing internal sensitivity studies to validate assumptions, the peer review panel requested specific analyses be evaluated to bound the results. The following sections briefly describe each analysis alternate and the impact to the analysis and/or design. Accidental Torsion An analysis was performed where the entire super imposed dead load mass was placed at the extreme corner of the diaphragm. The mass shift was done for two different corners of the diaphragm that were determined to be the governing locations. The shift resulted in a center of mass displacement of approximately 4% in the global X direction and 4% in the global Y direction, with a resultant of 5.7%. While this approach is not strictly consistent with the approach described in ASCE 7 where the user is required to offset the mass 5% in each orthogonal direction, it was determined among the design and peer review team to be a simple and equally applicable approach. For this study, it was determined that the dampers were the elements most sensitive to the mass offset. The average change in force among the dampers with the offset was less than 1%. The largest increase in force was 4.6% at damper A-1 under the Hector Mine record. It was determined that the analysis results were not significantly impacted with the inclusion of accidental torsion and the remainder of the analysis was not required to consider it. Upper Bound and Lower Bound Damper Properties In accordance with ASCE 41-13 C14.3.2.4, damper properties should be modified to account for specification tolerances, device characteristics, environment, and aging. The same section recommends an upper bound damper property modifier on the damping coefficient, C of between 1.15 to 1.20 and a lower bound value of 0.80 to 0.85. For this project the damping coefficient C was modified from its actual value of 160 kip-sec/in to an upper bound value of 184 kip-sec/in and lower bound value of 136 kip-sec/in. Implementing lower bound damper properties increased the drift by less than 0.1%. Applying upper bound damper properties increased the transfer/drag forces to the frames by up to 5.2% and increased the governing column uplift by approximately 12%. The forces in the dampers-frame system due to these variations are summarized in the following table:


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Demand-to-Capacity for Various C Values Governing Members Element C=160 kip-

sec/in C=184

kip-sec/in C=136

kip-sec/in Frame Beam 144% 143% 146% Damper Driver/Brace

56% 63% 49%

Column 135% 134% 137% Table 5. Damper Property Bounding Summary Upper Bound and Lower Bound In-Plane Diaphragm Stiffness The diaphragm stiffness modification factor used for in-plane and out-of-plane diaphragm stiffness was 0.25 and for many types of structures this assumption can be critical and greatly impact results. Very large values in the 0.8 range tend to overestimate slab and collector transfer forces, basement wall shear forces, and tend to underestimate foundation overturning forces. Very small diaphragm stiffness modification factors values in the 0.1 range tend to produce the opposite effect. A model sensitivity study was performed to compare the results on damper forces and structural period using an upper bound modification factor of 0.8 and a lower bound value of 0.1. The increase in the first three periods when using the lower bound stiffness modifier was 0.8%, 2.2%, and 1.6%, respectively. The average change in damper forces was 1.2%, and the maximum increase to any damper bay was approximately 4.1%. Diaphragms were designed to remain essentially elastic. Upper Bound and Lower Bound Out-of-Plane Diaphragm Stiffness The modification factor for out-of-plane diaphragm stiffness was 0.25. The analysis results were impacted by these assumptions to a lesser extent. The main concern with out-of-plane stiffness assumptions is that large values may lead to the slab coupling with gravity columns and falsely increasing the lateral stiffness and strength of the structure. A study was performed for stiffness values of 0.25 and 1. In summary, the difference in results between the two cases was minimal and not worth evaluating further.

Summary of Analysis Results

Drifts reduced from over 5% to 2% Mass reduction of approximately 14% Base shear and overturning moment reduced by a

factor of 2 or more in some directions Accelerations reduced in upper floors supporting

hollow clay tile infill Effective global damping ratio of approximately 25%

achieved Summary of Design The main elements of the retrofit consisted of:

Steel collector plates running the length of the 2nd floor diaphragm

Steel moment frame and damper assembly Concrete shear walls in the basement anchoring the

damper-frames New foundations FRP column confinement wraps and FRP hold-down

system FRP local diaphragm strengthening

Damper Location

Column High Beam

Low Beam

Damper Driver/Brace

North W12x305 W12x136 W12x106 HSS10x0.625 South W12x210 W12x106 W12x106 HSS10x0.625 East W12x305 W12x136 W12x106 HSS10x0.625 West W12x210 W12x106 W12x106 HSS10x0.625 Table 6. Damper-Frame Schedule


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Figure 23. Frame Key Plan

Figure 24. Damper Frame Elevation Peer Review While technically the renovation did not trigger a mandatory seismic evaluation and the analysis indicated that the inclusion of the damper frames would reduce seismic forces to the building, it was determined that a 3rd party peer review of the retrofit was required by the City of San Diego. A peer review panel was chosen consisting of three members: Structural Engineer (Nabih Youssef, NYA, panel chair), Geotechnical Engineer (Marshall Lew, AMEC Foster Wheeler) for seismic record selection and scaling, and a Damper Technical expert (Troy Morgan, Exponent Engineering) for review of the

damper design and specifications. The non-structural design disciplines would still go through a traditional permitting process. In order to avoid a lengthy review process impacting the overall project permitting and construction timelines, a phased plan review process was implemented. This consisted of a staggered review process of the following items:

Overall design criteria Time history selection and scaling Review of nonlinear analysis models Review of Construction Documents

Construction Working in any existing building is always a challenge, let alone a 100-year-old building utilizing a non-conventional retrofit. Because of that, there was some natural concerns about unforeseen issues that may come up as the project went into construction. That said, the overall construction went surprisingly smoothly. Items that were a consideration during construction regarding the damper and FRP retrofits were as follows:

Lead time – Dampers can have a very long lead time. This is due in part to the limited number of manufacturers available and their limited use in buildings. There are also extensive testing programs required of the dampers per ASCE 7-10 and ASCE 41-13 before they can be shipped for use. For the Hotel Churchill project, the long lead times were anticipated and were not a significant concern because all of the damper frames were going to be left completely exposed. For projects where dampers are a critical path item, the lead times can be a significant consideration.

Damper installation – While the dampers can be

adjusted slightly in the field in order to help with fit up, they can still be challenging to install for short bay frames with tall aspect ratios. Also, while they can be adjusted, installers must be careful to not unintentionally reduce the available stroke of the dampers during installation. Designers should take installation into account when specifying a damper.


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Figure 25. Installation of Damper

FRP Bond Strength – If FRP is being investigated as a retrofit tool during design, it is important that bond strength tests are performed on the concrete to determine that minimum bond strengths are available. Should sufficient bond not be present in older structures, it can render FRP unsuitable as a retrofit scheme. The concrete on the Hotel Churchill project was tested during design and construction and was just barely above the minimum strengths required.

Conclusion The renovation of the Hotel Churchill is a great example of a design team, ownership, and local jurisdiction working together to ensure the continued service life of a historic structure that had fallen into disrepair. The structural engineering team played a major role in devising affordable and innovative techniques to greatly increase the safety of the new residents.

Figure 26. Completed Project!

References Riahi, Zahra, et.al., 2009, “Backbone Model for Confined Masonry Walls for Performance-Based Seismic Design,” Journal of Structural Engineering, June 2009, pp. 644-654, American Society of Civil Engineers. American Society of Civil Engineers, “Seismic Evaluation and Retrofit of Existing Buildings,” ASCE 41-13. Shing, Benson, et.al., 2009, “Seismic Performance of Non-Ductile RC Frames with Brick Infill,” ATC-SEI Conference.