Seismic accelerations in architectural precast concrete cladding

Engineering Structures 180 (2019) 742–749

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Seismic accelerations in architectural precast concrete cladding Elide Pantoli a b

a,b,⁎

, Tara Hutchinson

T

a

Department of Structural Engineering, University of California, San Diego, 9500 Gilman Dr., La Jolla, CA 92093, United States Wiss, Janney, Elstner Associates, 2000 Powell Street Suite #1650, Emeryville, CA 94610, United States

A R T I C LE I N FO

A B S T R A C T

Keywords: Precast concrete cladding Cladding connections Architectural concrete Seismic accelerations Seismic forces Component amplification factor Structure amplification factor

Architectural precast concrete (APC) cladding is a type of building facade broadly used worldwide. It is composed of an array of concrete panels attached to the building structural system with steel connections. During earthquakes, the APC system can be damaged both by excessive drifts and accelerations. While research on driftcompatibility has been performed in recent years, to the author’s knowledge there are limited studies aimed to determining the level of seismic forces in APC cladding panels and connections. However, these forces could be larger than anticipated since this system has low natural frequencies due to its large mass and low stiffness in the out-of-plane direction. This paper presents an experimental and numerical study of the seismic accelerations developed in APC cladding panels installed on a building. The design limits provided by the code for ductile and brittle elements of the cladding system indicate that the code is able to meet the desired performance criteria both in case of service and design level motions. However, the satisfactory results obtained for design level earthquakes are due to the combination of significant errors in the prediction of the component and structure amplification factors.

1. Introduction Architectural precast concrete (APC) cladding is a type of facade composed of single concrete panels anchored to the building structural system with steel connections. Panels are separated from each other by joints filled with caulking. The significant use of this type of facade is related to its many benefits, which include durability, low life-cycle costs, aesthetics and the ability to create different colors, textures, and shapes [1]. In addition, the installation of APC cladding on buildings does not require the use of scaffolding, and is faster than that of many other types of facades. Furthermore, the panels can assume a wide variety of geometries. Types of panels include column covers, spandrel panels (Fig. 1a), and U-shape panels (Fig. 1b). The focus of this research is on punched window wall panels, which span from floor to floor and have openings for windows (Fig. 1c). 1.1. Seismic design of APC cladding Engineers need to design cladding connections to both allow interstory drifts and resist forces generated by earthquake motions. Driftcompatibility based design generally determines the overall configurations of the connection. In fact, the first goal of the design is to

provide flexibility between the panels and building in the in-plane direction, to prevent the very stiff APC panels from acting as structural shearwalls. This issue is generally resolved by allowing panels to move rigidly in-plane with one of the floors, typically the lower one, while allowing movement with respect to the other floor, generally the upper one. To achieve this goal, the bottom connections of the panels have large strength and stiffness, while those at the top must be designed to allow in-plane displacements while restricting out-of-plane movements. These two types of connections are named bearings and tiebacks, respectively. The main element of the tieback is generally a steel rod, which allows in-plane drifts by either flexing (flexing tieback) or sliding inside a slotted hole (sliding tieback). Fig. 2 shows an example of an APC panel and the location and details of typical connections as used on the West Coast of the United States. It is noted that a large variety of connection details are used throughout the U.S. and in other countries, but this paper focuses on typical practice in the seismic areas of the U.S. [2]. After drift-compatibility is ensured, designers must make sure that panels and connections are able to resist seismic forces. This is done by determining the design force Fp using Equation 13.3-1 in ASCE-7 [3]:

Abbreviations: APC, architectural precast concrete; BNCS, Building Nonstructural Components and Systems; UCSD, University of California, San Diego; FB, fixed base; PIDR, peak interstory drift ratio; PFA, peak floor acceleration; SDOF, single degree of freedom; PEER, pacific earthquake engineering research center ⁎ Corresponding author at: Wiss, Janney, Elstner Associates, 2000 Powell Street Suite #1650, Emeryville, CA 94610, United States. E-mail addresses: [email protected] (E. Pantoli), [email protected] (T. Hutchinson). https://doi.org/10.1016/j.engstruct.2018.11.062 Received 3 April 2018; Received in revised form 25 September 2018; Accepted 26 November 2018 Available online 06 December 2018 0141-0296/ © 2018 Published by Elsevier Ltd.

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Fig. 1. Typical configurations of architectural precast concrete cladding panels. Photographs courtesy of Willis Construction [http://www.pre-cast.org].

Single APC cladding panel

Tieback

Slab

R od

Panel

Panel

Slab

Bearings

Panel

Fig. 2. Example of bearings and tiebacks in an APC panel.

0.4ap SDS

Fp =

Rp Ip

z ⎛1 + 2 ⎞ Wp h⎠ ⎝

elements classified as flexible (natural period > 0.06 s). The component response modification factor Rp has the goal of reducing the design force to account for the ductility of an element and its potential to dissipate seismic energy. The coefficients ap and Rp for “exterior nonstructural wall elements and connections” are specified in Table 13.5-1 (ASCE-7 2010). The code assigns different values of these parameters for elements of the system that are considered ductile (wall elements and body of wall panel connections) and brittle (fasteners of the connecting systems). In this way, a brittle failure can be avoided, as the design force for brittle elements of the system is more than three times that of ductile elements and brittle failures should be prevented. Table 1 presents the values of ap and Rp for the various elements of APC cladding.

(1)

where

• S = spectral acceleration for short period; • a = component amplification factor; • I = component importance factor, varying from 1 to 1.5; • W = component operating weight, previously found for each connection; • R = component response modification factor; • z = height of the structure at point of attachment of component respect to the base; • h = average roof height of structure with respect to the base. DS

p

p

p

p

1.2. Damage to APC cladding during past earthquakes Damage to precast concrete cladding has been witnessed after many earthquakes in the past, both abroad and in the United States. Type of damage included:

ASCE-7 also specifies the maximum and minimum values of Fp to be 1.6SDSIpWp and 0.3SDSIpWp, respectively. Additional parameters used in this study are defined as follows:

• Permanent misalignment of the panels, crushing and cracking at

• C : seismic coefficient defined as: p

corners, as observed after the 2011 Christchurch earthquake in New Zealand [4], the 2011 El Mayor Cucapah earthquake in Mexico [5], and the 1994 Kobe earthquake in Japan [6]. Photographs showing

Fp

Cp =

Wp

(2)

• a : structure amplification factor, defined in the code formula as: s

z as = ⎛1 + 2 ⎞ h⎠ ⎝

Table 1 Summary of the values assigned by ASCE-7 (2010) to elements of APC cladding.

(3)

Key parameters in the ASCE-7 equation are ap and Rp. The component amplification factor ap recognizes the potential of an element to amplify floor accelerations due to its flexibility, and it is set as 1 for components considered rigid (natural period < 0.06 s) and 2.5 for

Wall element Body of wall panel connection Fastener of the connecting system

743

ap

Rp

1 1 1.25

2.5 2.5 1

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Fig. 3. Typical damage to APC cladding: (a) cracking and crushing of panels (Courtesy of Baird 2014 [4]), (b) failure of connections (Courtesy of FEMA-74 2011 [9]), and (c) collapse of panels (Courtesy of Di Croce et al. 2012 [43]).

• •

APC cladding and its components. It is noted however that very few studies have been conducted to determine the dynamic characteristics of APC panels. Memari et al. [29] used ABAQUS to calculate the natural modes and frequencies of two sample cladding panels, while Merrick et al. [30] achieved this same goal by performing tests on cladding panels installed on buildings. In addition, the authors of the present research conducted a study to determine the natural modes and range of natural frequencies of APC cladding panels as typically installed on buildings [31]. This goal was achieved by using experimental data to validate numerical models of the APC panels installed on a full-scale test building (see next sections), and then varying the characteristics of the panels and connections to represent typical configurations used in practice. They concluded that the movement of typical APC panels is dominated by three modes: out-of-plane bending, rocking in-plane, and vertical in-plane, as shown in Fig. 4. While the in-plane modes tend to have larger natural frequencies (> 17 Hz), the study showed that the out-of-plane bending modes can have frequencies as low as 8 Hz, well below the limit of 16.7 Hz that is conventionally considered the lower limit for a rigid nonstructural component.

this type of damage can be seen in Fig. 3a; Failure of connections, as seen after the 2010 Maule earthquake in Chile [7], the 2009 L’Aquila earthquake [8], the 1994 Northridge earthquake [9] (Fig. 3b); Collapse of panels, as after the 1964 Alaska earthquake, the 1987 Whittier Narrows earthquake [10], and the 2012 Emilia earthquake [11] (Fig. 3c).

It is important to point out that failure of connections and collapse of panels can be caused both by excessive drifts and/or excessive forces, depending on the specific circumstances. In all cases, damage to APC cladding can affect the functionality of a building after the earthquake, it is time consuming and expensive to repair, and it can also be life threatening.

1.3. Past studies on the seismic behavior of APC cladding Since the 1980s, a number of studies were performed to understand the complex building-cladding interaction and determine the possible stiffening effects of APC cladding on structural systems [12–14]. Other efforts have been devoted to developing special cladding connections designed to dissipate energy during earthquakes [15], and included analysis of friction-damped, elastomeric and tapered steel connections [16–18]. While promising, this type of connection has not become mainstream. In addition, a few studies were performed on the possibility of using APC cladding as part of the structural system [19,20]. This solution has also not observed widespread consideration. Some of the most comprehensive studies on APC cladding had the goal of understanding its seismic behavior of APC cladding as typically installed in buildings, with focus on drift-compatibility. Several groups of researchers performed system-level and component tests on tieback connections to achieve this goal. A large study performed by Wang consisted of the pseudo-static testing of a building encased with panels using typical Japanese and U.S. connections [21]. A large test program on panels and connections was also performed by Rihal [22]. This included component tests on connections, racking tests on full-scale panels and dynamic tests on reduced scale model of buildings and panels. Tests on cladding connections using different details were also performed by Craig et al., Pinelli and Craig, and Sack et al. [23–25]. McMullin et al. included two corner panels in a full-scale shake table test of a building [26]. Research on this topic was also conducted by Pantoli et al. [27], who performed shake table testing of a full-scale building including sixteen APC panels and component tests on tieback connections. Recently, tests were also conducted on panels and connections typical of the European practice by Belleri et al. [28]. These and other test programs offer valuable insights to the performance of

1.4. Scope of this paper Recent earthquakes have revealed the seismic vulnerability of APC cladding and the negative consequences that damage to this system could have in terms of life-safety, time, and cost of repair. Seismic damage to APC panels can be attributed to issues of drift-compatibility in the in-plane direction or excessive forces in the connections, especially in the out-of-plane direction. While in recent years several studies were performed to improve the drift-compatible design of connections, there has been little research focused on understanding the level of seismic forces generated by accelerations in APC cladding connections. Namely, there is no evidence that the limits imposed by the code for ductile and brittle elements of the connections are meeting important performance criteria, such as limiting damage during a service level intensity earthquake and avoiding failure during a design level earthquake. The goal of this study is to analyze the magnitude of seismic accelerations developed in APC cladding during service and design level motions, and to compare them with the limits imposed by ASCE-7 (2010). Namely, three parameters are considered in the analysis:

• The seismic coefficient C , which represents the level of acceleration in the panels; • The structure amplification factor a . Much research has been perp

s

formed in the past to investigate as [32–34], and its specific analysis is out of the scope of this paper. However, since as is an important component of Fp, it is relevant to present some key results relative to

744

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Fig. 4. Numerically estimated natural modes of APC cladding panels installed on the BNCS building: examples of out-of-plane modes [The contour shows the out-ofplane displacement]. Courtesy of Pantoli and Hutchinson 2015 [31].

different models were created. The response during service level motions is observed to remain quasi-linear, therefore a linear building model is used, while the building behaved nonlinearly at intensities at and above the design level, a more complex numerical model is utilized to capture the building nonlinear response. The next sections are devoted to the description of the BNCS building and the two models. It is noted that the panels moving in-plane were not analyzed because preliminary analysis revealed that they amplify acceleration less than out-of-plane panels. This is consistent with the findings of the dynamic analysis, which showed that out-of-plane modes have larger chances of being excited during an earthquake.

it;

• The component amplification factor a . Previous research on the dyp

namic characteristics of the APC panel has determined that they might not be rigid nonstructural components in the out-of-plane direction as the code currently classifies them. Thus, the values of 1 or 1.25 attributed by the code to ap are conceptually incorrect. One of the goals of this paper is to establish which are the correct values of ap for different levels of excitation. It is noted that, while the code provides one single value of ap per panel, this does not recognize the fact that panels are connected at multiple floors, thus the amplification from the bottom and top floors are different (Fig. 5). In the following study, ap,top and ap,bottom are defined as the peak acceleration in the panel divided by the peak acceleration in the top and bottom floors, respectively.

2.1. Building Nonstructural Components and Systems (BNCS) project Fig. 6 presents a photograph of the BNCS building including relevant dimensions. This building was subjected to thirteen earthquake input motion tests in the east-west direction (longitudinal building axis). During the first seven tests, the building was in a base isolated configuration, while it was fixed at its base during the last six motions. The six fixed-base (FB) motions were imposed in order of increasing intensity. The first two input motions (FB1, FB2) were selected from the

2. Method of analysis Central to this study is the use of accelerations measured on a fullscale test building central to the Building Nonstructural Components and Systems (BNCS) project. This building consisted of five-floors and was outfitted with a large variety of nonstructural components, including sixteen APC panels. The BNSC building was tested on the unidirectional large outdoor shake table at the University of California, San Diego (UCSD). This study is based both on experimental data obtained directly from the BNCS building and numerical data calculated through models of the BNCS building and panels created in SAP2000 [35]. Both service and design level earthquake motions are considered and, due to the large difference in the behavior of a building in these cases, two

Fig. 5. Schematic showing the amplification of acceleration in the panels compared to their top and bottom floor.

Fig. 6. Photograph of the BNCS building from the north-west corner. 745

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3

motions recorded during the Northridge earthquake in California. They were spectrally matched to achieve service level performance at the fundamental period of the building, and generated a peak interstory drift ratio (PIDR) of 0.56% at the second level and a peak floor acceleration (PFA) of 0.44 g at the roof. The fifth input motion (FB5) was recorded during the Denali earthquake in Alaska and it was scaled to achieve design level performance of the building. During this motion, the structure reached a PIDR of 2.75% at the second level and a PFA at the roof of almost 1 g. Two APC cladding panels were installed on each side of the building at the fourth and fifth floors, for a total of sixteen panels installed on the building. Since this shake table allows only east-west motion, the eight larger panels in the north and south faces of the building moved mainly in the in-plane direction, while those on the east and west faces moved primarily in the out-of-plane direction. This study focuses on the analysis of the out-of-plane panels in the south-east corner of the building at the fourth and fifth floors, since these were the out-of-plane panels installed with accelerometers, and therefore they provide representation experimental data documenting the panel’s response. More information about this test program can be found in Chen et al. 2016 [36] and Pantoli et al. 2016 [27].

Experimental Numerical

2.5

Psa (g)

2 1.5 1 0.5 0

0

1

2

3

Period (sec) Fig. 8. Comparison of the numerical and experimental 5% damped elastic acceleration response spectra for the panel on the fifth floor during service level motion FB2.

the model in SAP2000 including the panel. A linear dynamic analysis was performed using the modal solution with Ritz vectors. The input motions were FB1 and FB2, the pair of service level motions utilized in the BNCS project while the building was fixed to the shake table. The optimal levels of damping were found by minimizing the error between the numerical and experimental results in terms of peak floor accelerations, peak panel accelerations, and component amplification factors. In addition, it was verified that these values led to good comparisons in terms of time history and response spectra. Fig. 8 shows a sample comparison between numerical and experimental pseudo acceleration spectra for the panels, and generally demonstrates a reasonable response. The variables considered for the linear study were:

2.2. Linear model of the BNCS building and panels A linear model of the BNCS building was created in SAP2000. The model included explicit representation of columns, beams, slabs and elevator walls and used the actual geometry of the midline of the elements, as reported in the construction drawings [37]. The weight and the inertial masses of the nonstructural components were added as lumped masses. The properties of the cast-in-place concrete used to construct the different structural elements were determined by compressive tests conducted at the beginning of the earthquake testing sequence. Subsequently, the stiffness of the elements was reduced to account for concrete cracking, with reduction factors selected from the range presented in the literature [38–41] to match the experimental frequency of the main lateral mode of the building. The final stiffness reduction factors were in the range 0.65–0.7 for the columns and 0.45–0.5 for the beams and slabs. The panels were added to the building as single degree of freedom (SDOF) systems created by a mass and two springs connecting the mass to its top and bottom floors. The bottom spring was rigid because it modeled the stiff bearing connections. The top spring was considered rigid in the vertical and horizontal in-plane directions while it was assigned a certain flexibility in the out-of-plane direction. The out-ofplane flexibility of the spring was determined using the natural frequency of the out-of-plane mode and the mass of the panel. Fig. 7 shows

• The location of the panel. The SDOF model of the panel was moved at the fourth and fifth floors; • The natural frequency of the panels, which was shifted at 8 Hz,

•

11 Hz, and 14 Hz. These frequencies were selected because they were within the range of possible frequencies for out-of-plane modes and under the limit value of 16.7 Hz. These different frequencies were achieved in the model by changing the stiffness of the top spring of the SDOF representing the panel; Earthquake motions. Eighteen additional earthquake motions were selected from the Pacific Earthquake Engineering Center (PEER) database (http://ngawest2.berkeley.edu) and scaled to be compatible to a service level demand.

2.3. Nonlinear model of the BNCS building and panels The nonlinear model was created from the linear model by reducing the stiffness of the linear elastic structural elements to account for further cracking and by adding plastic hinges at the base of the columns at the first floor and at the ends of the beams at all levels. The plastic hinges at the base of the columns were modeled as fiber hinges accounting for the presence of the normal force and moment in the two directions. The Park constitutive model was used to capture the behavior of the reinforcing steel elements, with its relevant parameters generally determined using results from tensile tests on the rebar stock used in the test building. Mander’s model was used for the unconfined concrete. The relevant parameters were found by compressive tests on concrete cylinders. The properties of the confined concrete were calculated using the procedure proposed by Chang and Mander [42]. The length of the plastic hinges were estimated using the formula proposed by Pauley and Priestley [41]. The plastic hinges in the beams were modeled as lumped plasticity hinges with assigned moment-rotation curves. The moment-curvature of each hinge was determined via an

Spring Mass

Fig. 7. Model of the BNCS building and a panel in SAP2000. 746

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6

Experimental Numerical

5

PSA (g)

4 3 2 1 0

0

1

2

3

4

Period (sec) Fig. 9. Comparison of the numerical and experimental 5% damped elastic acceleration response spectra for the panel on the fifth floor during design level motion FB5.

Fig. 10. Values of Cp achieved for the panel at the fourth and fifth level during service and design level motions.

independent analysis of the section using a fiber model and theoretical calculations. The type of analysis performed was a nonlinear time history analysis using the broader suite of scaled earthquake motions. This was performed through direct integration and considering Rayleigh damping. A critical step in this analysis was to determine the optimal damping values. These values were found by minimizing the prediction errors between the experimental and numerical results for FB5 in terms of peak accelerations and component amplification factors. Motion FB5 was selected as it imposed design level demands on the building. A parametric study on the influence of the damping parameters was performed by systematically modifying the four damping parameters up to 40%. The time histories and pseudo acceleration response spectra obtained at the floors and panels were also compared. Fig. 9 shows the comparison of numerical and experimental pseudo acceleration response spectra during FB5 for the panels on the fifth floor. These results show that the model tends to underpredict the influence of higher periods, such as that of the main mode of the building, while it overpredicts the influence of the lower periods, such as that of the natural mode of the panel and higher modes of the building. In terms of variables, the location of the panels and the input motions were varied in this case.

bottom of the panels, respectively. In this case, the results for the panels at the fourth and fifth floors are similar, and thus they are presented together. The data (Fig. 10) show that, during a service level earthquake, the achieved Cp are well below the code limit for brittle elements. However, they are very close to the code limits for ductile elements, due to the reduction of the forces by the arbitrary factor Rp. For the panels installed at both floors, the code limit relative to the top floor represents the 70th percentile of the distribution of the actual Cp, meaning that in 30% of the cases the actual Cp is larger than the code limit. This might seem large, however note that the 95th percentile of the distribution of Cp is only 1.1 times the code limit, thus the application of the other safety factors in design should be able to prevent damage to the top connections. The situation is different when comparing the actual values of Cp with the code limit relative to the bottom floor. In fact, in this case the code limit represents the 35th percentile of the distribution, and the value of the 95th percentile is roughly 1.3 times the code limit. Damage should be avoided also in this case when bulky bearing connections are installed at the bottom of the panels, since these forces should still be too small to damage them. However, damage could potentially happen if the location of tiebacks and bearing connections is reversed. Considering the design level motion, all values of Cp are in between the code limit for ductile and brittle elements, thus satisfying the performance goal in this case. For panels at both floors, the 95th percentile of the distribution of Cp is roughly 60% and 70% of the code limit for brittle elements for the top and bottom floors, respectively. This can be considered reasonable levels of conservatism. When compared to the code limit for ductile elements, the 95th percentiles of the distributions are 1.9 and 2.2 times the limits for the top and bottom floors, thus damage to ductile elements can be expected.

3. Results of the parametric analysis The analysis of the results focuses on three parameters: the seismic coefficient Cp, the structure amplification factor as, and the component amplification factor ap. For each of the parameters considered, results are presented for service and design level motions. The performance goals assumed are to prevent widespread damage in the first case and to avoid connection failure in the second. For all the three coefficients, the statistical parameter presented to compare with the code value is the 95th percentile of the distributions. It is noted that, while this is considered reasonable for normal structure, a higher value might actually have been used for very important structures, such as hospitals.

3.2. Structure amplification factor as and component amplification factor ap To fully understand the results relative to Cp, it is important to analyze two of its components: as and ap. It is noted that the study of the parameter as is not the main purpose of this research, however it is necessary to understand its amplitude since the input to the panel is provided by the floor accelerations. Fig. 11 shows the results relative to as for service and design level motions. In case of service level motions, the distribution of the 95th percentiles of as is linear, reflecting the predominance of the first mode of the building as would be anticipated. In this situation, the panels at the two floors receive a similar input, and are expected to have similar values of ap. Very different is the behavior during design level motions, which is characterized by a general

3.1. Seismic coefficient Cp Fig. 10 presents the experimental and numerical values of Cp for both the linear and nonlinear model. It is noted that this study is based on the assumptions that the APC panels can be considered SDOF systems and thus have a single value of acceleration, which is that of the mass of the SDOF located at mid-floor height. The achieved values of Cp can thus be compared to the code limit at the top and bottom floors, corresponding to the design limits for the connections at the top and 747

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percentile of the distribution of ap,bottom being 2.3, while that of ap,top being 1.5. The values of ap,top and ap,bottom for the panel at the fourth floor are closer to each other, with an average of 1.85. The comparison with the code limits are not good in both cases, since the code does not consider a possible amplification of accelerations in the panels. In case of the design level motion, the underestimation of ap values is between 1.5 and 2.3 for the ductile elements and from 1.2 to 1.8 for brittle elements. 4. Conclusions Architectural precast concrete (APC) cladding is a nonstructural component sensitive to both interstory drifts and floor accelerations. While recent research has aimed at improving the drift-compatible design of APC cladding connections in the past couple of years, a systematic analysis of the level of forces generated in APC cladding connections during earthquakes has not been undertaken. Nevertheless, the development of large seismic forces not accounted for in the design phase is a real possibility. In fact, past studies showed that APC cladding can have low natural frequencies in the out-of-plane direction and thus could amplify the floor acceleration, possibility not accounted for in current design codes. The goal of this study was to analyze the seismic accelerations developed in the out-of-plane direction for APC cladding panels installed in building. The effort first involved analysis of measured results from a full-scale building test program, where sixteen APC panels were installed. The experimental results were complemented with those from numerical analyses of the building and panels modelled in SAP2000. The first parameter studied was the seismic coefficient Cp, which is defined as the seismic force divided by the operating weight of the component. In addition, the structure amplification factor as and the component amplification factor ap were analyzed to shed further light on the results relative to Cp. Results were examined in the context of service-level and design-level anticipated demands. In the former, the building is assumed to remain linear elastic, while in the latter the building behavior was assumed nonlinear. Different performance goals were assigned to these two cases, namely to avoid widespread damage during a service level motion and to avoid failure of brittle components of the system during a design motion.

Fig. 11. Values of as achieved for the panel at the fourth and fifth level during service and design level motions.

deamplification of accelerations at the fourth and fifth floors, with 95th percentiles of the distribution of as close to unity. The roof has larger peak accelerations than that of the other floors, a phenomenon well documented in the literature. In this case, the 95th percentile of the distribution is 1.7. This behavior creates two different types of input to the panels. Namely, when the panel is at an intermediate floor, the peak accelerations from the top and bottom floors are similar, and the values of ap,top and ap,bottom are expected to be similar. When the panel is installed at the top floor, it receives much larger input from the roof, thus the amplifications from the bottom of the panel are expected to be much larger. Finally, it can be observed that, while the code formula should be relative to a design level motion, the limits for the structure amplification factor as resemble more closely those relative to a service level motion. The overprediction of as for the design level motion are considerable, with the code limits being between 1.7 and 2.4 greater considering the 95th percentiles of the distribution of as. Fig. 12 shows the distribution of the values of the component amplification factors, again considering the motion at the bottom and top of the floor i.e. ap,bottom and ap,top, respectively. As expected, in the case of the service level motions, the results are similar for the two panels, and at each floor, the values of ap,bottom are larger than those of ap,top. The 95th percentile of the distributions are around 1.4 for ap,top and 1.65 for ap,bottom. The results for the design level motions are consistent with what was expected from the distribution of as. For the top panel there is a large difference between ap,bottom and ap,top, with the 95th

4.1. Implications relative to current design code The results relative to Cp show that both performance goals are met for the cases analyzed. For service level motions, the values of Cp are very close to the design limit of ductile elements, but still small enough to avoid widespread damage. The values of Cp are generally larger for the bottom connections, due to the pulling effects of the top floor. However, the presence of bulky bearing connections at this location is expected to prevent damage. In the case of design level motions, the values of Cp are between the limits for ductile and brittle elements. The 95th percentiles of the distributions are 1.9–2.2 times the limits for the ductile elements, but only 60–70% the limit for brittle elements, thus providing a certain level of conservatism. Results relative to the parameter as were examined since the floors accelerations are the input to the APC cladding panels. These results revealed consistency with previous studies. In the case of service level motions, the values of as had a linear distribution similar to that presented in the design code, thus creating a larger input at the top of the panels than at their bottom. In the case of design level motions, the values of as are close to unity for intermediate floors and larger at the roof due to higher modes effect, however remain very different from those presented by the code formula, which overestimated as by a factor of roughly 2 (from 1.7 to 2.4). The values of ap calculated in this study are much larger than those provided in the code, which does not consider amplification of the accelerations in the panels. In all the cases, it can be observed that the code formula underestimates the value of ap of

Fig. 12. Values of ap achieved for the panel at the fourth and fifth level during service and design level motions. 748

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a factor from 1.5 to 2.3 for ductile elements and from 1.2 to 1.8 for brittle elements. It can be concluded that the resulting experimental and numerically calculated seismic coefficient meet the performance criteria set, however, only by combining a series of large errors in the term of the equation to calculate Cp. In this case, the compensating errors of an overestimation of as of between 1.7 and 2.4 and an underestimation of ap by between 1.2 and 2.3 leads to consistency when comparing the final Cp value with design, and this appears to lead to acceptable levels of conservatism.

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Acknowledgements The development of this paper was supported by the Charles Pankow Foundation Research Grant Agreement #02-11. Experimental results supporting suggestions within this guide were conducted as part of the Building Nonstructural Components and Systems (BNCS) project. The BNCS project is collaboration between four academic institutions (University of California, San Diego, San Diego State University, Howard University, and Worcester Polytechnic Institute), four government or granting agencies (the National Science Foundation grant CMMI-0936505, the Englekirk Advisory Board, the Charles Pankow Foundation, and the California Seismic Safety Commission), more than forty industry partners, and two oversight committees. Many individuals contributed to the overall effort of the BNCS project. We particularly thank core team members Profs. Jose Restrepo and Joel Conte, doctoral students Rodrigo Astroza, Michelle Chen, Hamed Ebrahimian, Steven Mintz (deceased), and Xiang Wang, the NEES@ UCSD and NEES@UCLA staff, Dr. Robert Englekirk, Mr. Mahmoud Faghihi, Dr. Matthew Hoehler, Prof. Ken Walsh of SDSU, and SDSU students Consuelo Aranda and Elias Espino. In addition, the input of Robert Bachman, chair of the project's Engineering Regulatory Committee, is greatly appreciated. A listing of industry project sponsors and project participants may be found on the project website: http:// bncs.ucsd.edu/index.html. The development of the study about the cladding included in the BNCS paper was particularly supported by Mark Hildebrand, Glen Underwood, and Kurt McMullin. Appendix A. Supplementary material Supplementary data to this article can be found online at https:// doi.org/10.1016/j.engstruct.2018.11.062. References [1] PCI-Precast/Prestressed Concrete Institute. PCI MNL 122. PCI Architectural precast concrete manual. Chicago, IL: PCI; 2007. [2] Pantoli E. Seismic behavior of architectural precast concrete cladding panels and connections [PhD dissertation]. San Diego: University of California; 2016. [3] American Society of Civil Engineers. ASCE 7-10 Minimum Design Loads for Buildings and Other Structures. Reston, VA: ASCE; 2010. [4] Baird A. Seismic performance of precast concrete cladding systems [Doctoral dissertation]. New Zealand: University of Canterbury; 2014. [5] EERI. The El Mayor-Cucapah, Baja California Earthquake April 4, 2010, An EERI Reconnaissance Report. Oakland, California: Earthquake Engineering Research Institute; 2010. [6] Horii S, Oka S, Inukai M, Kohno K, Sakamoto I, Seike T. A report on the damages of precast concrete curtain walls by the 1995 hyogo-ken nanbu earthquake. Tokyo, Japan: PCSA; 1995. [7] Ghosh SK, Cleland NM. Performance of precast concrete building structures. Earthquake Spectra 2012;28(S1):S349–84. [8] Miyamoto International. L’Aquila, Italy, Earthquake Field Investigation Report. West Sacramento, CA; 2009. [9] FEMA E-74. Reducing the Risks of Nonstructural Earthquake Damage: A Practical Guide; 2011. [10] Arnold C. Building envelope design guide – introduction. Whole Building Design Guide; 2009. [January, 2015]. [11] Bournas DA, Negro P, Taucer FF. Performance of industrial buildings during the Emilia earthquakes in Northern Italy and recommendations for their strengthening. Bull Earthq Eng 2013;12(5):2383–404. [12] Goodno BJ, Palsson H. Analytical studies of building cladding. J Struct Eng 1986;112(4):665–76.

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