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Shake table investigations on code non-compliant reinforced concrete frames
Alexandria Engineering Journal (2020) xxx, xxx–xxx
H O S T E D BY
Alexandria University
Alexandria Engineering Journal www.elsevier.com/locate/aej www.sciencedirect.com
ORIGINAL ARTICLE
Shake table investigations on code non-compliant reinforced concrete frames Muhammad Rizwan *, Naveed Ahmad, Akhtar Naeem Khan, Samiullah Qazi, Junaid Akbar, Muhammad Fahad Earthquake Engineering Center, Department of Civil Engineering, University of Engineering & Technology, Peshawar, Khyber Pakhtunkhwa 25120, Pakistan Received 5 November 2019; revised 26 December 2019; accepted 29 December 2019
KEYWORDS Shake table test; SMRF; ACI-318; Damage scale; Response modification factor; Beam-column joints
Abstract Shake-table tests were performed on a code conforming and four non-compliant 1:3 reduced scale two-story RC frames. The models were subjected to 1994 Northridge earthquake acceleration time history, linearly scaled to multiple levels for studying the progressive damages. The code conforming model was observed with plastic hinges at the beam-ends and base of columns on ground story. This frame incurred only slight cracks in the joint panels under extreme shaking. The code-conforming model resisted input excitation with peak horizontal acceleration of 1.0 g before attaining the incipient collapse state. The non-compliant models were observed with plastic hinges in beam/column members, and incurred extensive damages in beam-column joints under extreme shaking. The fully non-compliant frame resisted peak horizontal acceleration of 0.50 g before attaining the incipient collapse state. The response modification factor of codeconforming model is 7.54 that reduced to 2.90 in case of fully non-compliant frame, which was due to the non-compliance of constructions. This was partly due to the reduction in the overstrength of frame (which was about 20%) and partly due to the reduction in ductility of frame (which was about 50%), in comparison to the seismic response parameters given in the seismic code. The less redundant structural system and joint shear hinging are the main reasons resulting in the reduction of overstrength and ductility, respectively. Static force based performance assessment in various seismic zones revealed that the code-conforming frame performance is ‘‘OK” while noncomplaint frames performance is ‘‘NG” in both seismic zone 3 and zone 4. The present research suggests a reliability/redundancy factor to take in to account in proposing response modification factor for seismic design of structure. This will take care of the construction deficiencies found in the region. This factor can vary from 1.0 (in case of fully compliant structure) to 0.40 (in case of fully non-compliant structure). Ó 2020 Faculty of Engineering, Alexandria University. Production and hosting by Elsevier B.V. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/).
* Corresponding author. E-mail address: [email protected] (M. Rizwan). Peer review under responsibility of Faculty of Engineering, Alexandria University. https://doi.org/10.1016/j.aej.2019.12.047 1110-0168 Ó 2020 Faculty of Engineering, Alexandria University. Production and hosting by Elsevier B.V. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/). Please cite this article in press as: M. Rizwan et al., Shake table investigations on code non-compliant reinforced concrete frames, Alexandria Eng. J. (2020), https:// doi.org/10.1016/j.aej.2019.12.047
2 1. Introduction Reinforced concrete construction is on the peak in many urban areas of the world. This type of construction is particularly used for multi-storey buildings, schools, hospitals, and residential type buildings because of vast availability of constituents of concrete around the world and ease of construction. Although the design and construction of RC buildings in the developing countries are mostly based on the locally available or adopted international building codes, the proper construction execution of the specified design in the field is still a big challenge. The improper construction practices have resulted in many substandard and vulnerable building stock [1,2], with recent surveys have shown a number of construction/design deficiencies [1,3]. The commonly available construction/design defects include; substandard quality of concrete (low strength concrete), empty joints (joint panels without transverse reinforcement), beam column members having reduced flexural reinforcement along with shear reinforcement spacing larger than the code specified and practicing non-seismic hooks with among others. It is also worth to mention, majority of these RC buildings in the developing countries are located in high seismic regions and can be subjected by future large earthquakes of extreme excitations [4]. Reinforced concrete structures if not built properly can result in catastrophic failure and subsequent human and economic losses, upon subjecting to earthquake induced strong ground motions [5–8]. In the recent decades, many researchers have performed experimental investigations through shake table testing on RC frame structures. In most of these experimental studies, either full scaled and/or reduced scaled structures have been employed and the work were mostly focused on the vulnerability assessment of gravity type/old type RC buildings and/or to assess the performance of a specific retrofitting/strengthening/ isolation technique [9–29] However, experimental shake table investigations on the recently employed SMRF structures with and without the aforementioned construction deficiencies are lacking. In the current study, detail experimental shake table investigations have been performed on RC SMRF structures complying to code requirements i.e. code compliant frame and with construction/design defects i.e. code non-compliant frames. The shake table tests investigations have been performed on five one-third reduced scaled two storey RC frame specimens. Model-1 was a code compliant model (reference model) design based on the seismic building code, whereas all other models were code non-compliant specimens with the inclusion of construction/design defects systemically. The deficiencies which are included in the current study are; substandard quality of concrete (low strength concrete), empty joints (joint panels without transverse reinforcement), beam column members having reduced flexural reinforcement along with shear reinforcement spacing larger than the code specified and practicing non-seismic hooks. All the test specimens were excited with multiple excitations using the natural accelerogram of 1994 Northridge Earthquake, ranges from 5% to 130% of peak ground acceleration. The idea of using multiple scaled excitations is to force the test models from elastic stage to inelastic and finally near collapse stage. The damage mechanism of each
M. Rizwan et al. specimen is observed and reported. Acceleration and displacement response of the structure is recorded and analyzed to obtain the inter-storey drift demand, inter-storey shear and displacement profile of the structure and to develop deformation-based damage scale for seismic assessment of the considered structures. 2. Experimental program 2.1. Details of structures under investigation The current research focuses on low rise RC frame structures with and without construction/design defects, in order to quantify their seismic performance. In particular, an RC frame with two storey structure is considered, typically employed for low rise schools, hospital or apartment type public buildings. The considered structure consists of 2 by 1 bay frame, with each bay length of 18 feet (5487 mm) and storey height of 12 feet (3658 mm) for both the stories, as shown in Fig. 1. The representative frame structure was design according static force-based procedures (BCP-SP 2007/UBC 97) [30,31], considering high seismic zone of 4.0 (0.40 g design peak ground acceleration) with stiff soil type B (NEHRP classification). The modeling and design were carried out using the design software CSI ETABS considering all load combinations as per the code. For the material properties, 3000 psi (21 MPa) concrete and 60,000 psi (414 MPa) rebar yield strength was considered in the design process. The design model was detailed as per the ACI-318–14 [32] recommendations for the SMRF provisions. Fig. 1 shows the design of the prototype structure. 2.2. Construction defects under investigation As already mentioned, the improper construction and execution of the design specifications in the field is still a major challenge in many developing countries, which has caused many deficient building stock [1,2]. The most commonly available deficiencies include; substandard quality of concrete i.e. low strength concrete, empty joints i.e. joint panels without transverse reinforcement, beam column members having reduced flexural reinforcement along with shear reinforcement spacing larger than the code specified and practicing non-seismic hooks. In the current experimental campaign, these construction/design defects have been considered systematically. Model-1 is a reference test model and is based on the code design requirements. Model-2 is similar to Model-1 but provided with low strength concrete of 2000 psi (14 MPa). Model-3 is similar to Mode-1 and Model-2, but along with low strength concrete, no transverse ties have been provided in the beam-column joint panel. Similarly, Model-4 and Model-5 were similar to the Model-3 but were provided with beam and columns transverse reinforcement spacing were 100% (6 in. [152 mm] c/c) and 200% (9 in. [228 mm] c/c) larger than the previous models, respectively. In Model-5 the beam flexural reinforcement was also reduced from (3 + 3) # 6 bars to (2 + 2) # 6 bars and ties were provided with non-seismic 90° hook in both beams and columns members.
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Shake table investigations on code non-compliant reinforced concrete frames
3
Beam secon A-A Fig. 1
Prototype geometric layout and design specifications.
2.3. Preparation of one-third reduced models The seismic shake table simulator of Earthquake Engineering Center Civil Engineering Department has a table size of 5 feet by 5 feet (1.5 m 1.5 m) and can operate with loading capacity of about 5 tons only in the unidirectional excitation. Due to size and loading capacity limitations, only the interior critical frame was extracted from the prototype design building and reduced by one-third scale for seismic excitation. For scaling between the prototype and model dimension, a simple linear
Table 1
Column secon B-B
model idealization was considered [16,33], as shown in Table 1. Such simple modeling allows simplicity and also reduced the cost as well as complexities associated in scaling the stressstrain properties of the materials (concrete and reinforcing bar) in the model domain. As shown in Table 1, all the members of the extracted prototype frame i.e. beam, column and slab, and reinforcing rebar were reduced by a scale factor of SL 3. For the constituents of concrete, 3/8in down aggregate size have been employed for model preparation, in order to used scaled coarse aggregate, whereas cement and fine aggre-
Similitude conversion factors and reduced models’ dimensions.
Structural Properties
Simple Model Similitude Requirement
Prototype Frame
Test Models (Scale 1:3)
Physical Quantity
Relationship
Scale Factor
Beams: 12in 18in (304 mm 459 mm) Columns: 12in 12in (304 mm 304 mm) Slab: 6in (153 mm) Concrete Strength: 3000 psi (21 MPa) 2000 psi (14 MPa) Aggregate size: 1in Steel Strength: 60,000 psi (414 MPa) #6 Rebar (19 mm) #3 Rebar (10 mm)
Beams: 4in 6in (102 mm 153 mm) Columns: 4in 4in (102 mm 102 mm) Slab: 2in (51 mm) Concrete Strength: 3000 psi (21 MPa) 1:1.80:1.60, W/C = 0.48 2000 psi (14 MPa) 1:3.50:2.87, W/C = 0.80 Aggregate size: 3/8in Steel Strength: 60,000 psi (414 MPa) #2 Rebar (6.33 mm) #1 Rebar (3.33 mm)
Length Stress Strain Specific Mass Displacement Force Time Frequency Velocity Acceleration
SL = Lp/Lm Sf = fp/fm Sє=єp/єm Sq = qp/qm Sd = dp/dm = SL SF = Fp/Fm = S2LSf p St = tp/tm = SL (SєSq/Sf) SX = Xp/Xm = 1/St p Sv = vp/vm= (SєSq/Sf) Sa = ap/am = Sf/SLSq
3 1 1 1 3 9 3^0.5 (1/3)^0.5 1 1/3
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M. Rizwan et al.
Fig. 2
Test frames construction stages.
gate have been used with no scaling. As mention earlier two types of concrete strength have been employed in this study i.e. 3000 psi (21 MPa) and low strength test models with a 2000 psi (14 MPa) concrete compressive strength. For this purpose, ACI concrete mix design methods were followed for the
preparation of concrete constituent mix. The concrete mix proportion of cement:sand:corase aggregate, for both design specified strength of 3000 psi (21 MPa) and reduced strength of 2000 psi (14 MPa) are reported shown in Table 1. Fig. 2 shows the sequence of construction for the one-third scaled frame
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Shake table investigations on code non-compliant reinforced concrete frames
Fig. 3
Fig. 4
5
Test frames shifting to testing laboratory.
Added floor masses preparation and setup for artificial mass simulation.
models, Whereas Fig. 3 shows the test model transportation to the testing laboratory. Due to scaling and similitude requirements the reduced scale models were subjected to gravity and seismic mass less than the required. To satisfy the condition of mass simulation
for the reduced scaled test models, additional floor mass have been applied following the mass simulation model [16,33]. On each floor level extra mass of 1200 kg were applied through steel blocks mounted and fixed to the floor by means of fully secured ø inch (13 mm) steel bolts as shown in Fig. 4.
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M. Rizwan et al.
Fig. 5
Table 2
Test frame instrumentation and shaking setup.
Test model instrumentation positions and characteristics.
Channel
Position
Direction
Maximum Capacity
Parameter
Coefficient
Unit
A1 A2 A3 A4 A5 A6 D7 D8 D9
Pad level First Floor Second Floor Pad level First Floor Second Floor Pad level First Floor Second Floor
Front Front Front Back Back Back Front Front Front
±10 g
Acceleration Acceleration Acceleration Acceleration Acceleration Acceleration Displacement Displacement Displacement
492.20 501.10 510.10 508.90 490.10 502.00 1000.00 1000.30 1000.20
mv/g mv/g mv/g mv/g mv/g mv/g mv/inch mv/inch mv/inch
24 in. (610 mm)
Table 3 Earthquake Engineering Centre (EEC)’s seismic simulator (shake table) limits. Shaking simulator characteristics
Limits
Excitation direction
Single degree of freedom/unidirection 5 feet 5feet (1.5 m 1.5 m) 5 tonnes 1.1 g ±1.1 m/s ±125 mm
Physical dimension Pay load capacity Maximum acceleration Maximum velocity Maximum displacement
frames, so only external instrumentation was employed in the form of six accelerometers and three linear displacement transducers. The instruments sensitivity, maximum capacity and conversion coefficient have been reported in Table 2. On each floor level (mid position of joint panel region) and at base pad level, three accelerometers were installed on front and back side of the test models in order to record floor and base pad accelerations. For recording the displacement at each floor and pad level, three displacement transducers were attached on a fixed steel frame, which was install in-line with the model in-plane position.
3. Shake table tests on two-storey RC frames
3.2. Shake table Input loading protocols
3.1. Testing setup and model instrumentation
All the test models were tested by the Earthquake Engineering Center (EEC)’s seismic simulator in a uni-directional motion. Table 3 reports the characteristics limit values of the shake table. In order to excite the test models from elastic stage to full ultimate collapse stage, and also to be within the shaking
The test model’s setup on the shaking simulator and model instrumentation are shown in Fig. 5. The objective of experimental testing was to observe the global response of the test
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Shake table investigations on code non-compliant reinforced concrete frames
7
Time history record of 1994 Northiridge Earthquake
5% Damped Acceleration Response Spectra Fig. 6
Table 4
5% Damped Displacement Response Spectra
Input time history record.
Input protocol and testing sequence for Model-1.
Time History
Run (%)
Observed Input PGA (g)
1994 Northridge Earthquake
Self-Check 5% Run 10% Run 20% Run 30% Run 40% Run 50% Run 60% Run 70% Run 80% Run 90% Run 100% Run Self-Check 130% 130% Run
0.60 0.033 0.06 0.12 0.16 0.19 0.25 0.31 0.36 0.41 0.49 0.62 0.62 1.06
range of shaking simulator, the 1994 Northridge Earthquake natural acceleration record was selected. This selection was made after careful analysis of a number of natural acceleration records to be within the range of shake table limiting acceleration, velocity and displacement limits. Fig. 6 shows the horizontal component of 1994 Northridge Earthquake record acceleration time history, elastic accelerations and displacement spectra for 5% damping. The record has been obtained from PEER strong motion data base (090 CDMG Station 24278) and has a peak acceleration, velocity
and displacement values of 0.57 g, 20.39 in/sec (518 mm/sec) and 3.54 in (90 mm), respectively. The acceleration record time step was reduced by 1/3 to respect the input frequency requirement for the model as mentioned in Table 1. Table 4 reported the input multiple excitations and test sequence for code compliant Model-1. Each frame model was excited with incremental excitations ranges from 5% to 130% depending on the ultimate capacities of the tested frame. Input multiple excitations and test sequences for all other models are reported in Appendix A. The idea of using multiple scaled excitation i.e. from low level to high level, was to deform and force the test structures from elastic stage to inelastic and then to attain full ultimate incipient collapse stage. The EEC’s seismic simulator performs a self-adjustment motion called Self-Check, once the input time history is given to the shake table. After the simulator self-check adjustment, the test models were excited with scaled incremental excitations of the maximum acceleration records. After each run the damage mechanism of the test models was observed and documented with snapshot. The test sequence was progressively increase until the test models reached to near incipient collapse stage, after which the test was concluded. The recorded acceleration and displacement response time histories were obtained for each test run in the form of voltage values. 3.3. Recorded data processing The accelerometers and displacement transducers record data in the form of current values (voltage, mv). To get the time histories values in the form of accelerations (g) and displacements
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M. Rizwan et al. Table 5
Model-1 experimentally observed accelerations histories (prototype domain).
Run
Floor level
Acceleration History
Ground floor level
Self-Check Run
First floor level
Ground floor level 100% Run First floor level
Ground floor level 130% Run First floor level
(mm), the recorded raw data needs to be divided by the instruments conversion coefficient as reported in Table 2. Once the recorded data have been corrected for the respective instrument coefficient, the raw data was further processed for base line correction and signal filtering. This data processing correction and filtering was done to remove any noise in the actual recorded data. For this purpose, the SeismoSignal (SeismoSoft 2015) data processing software have employed with a linear category base line correction and filter type of Butterworth with frequency range of 0.10–25 Hz were considered. Once the processed data have been obtained, the displacement and acceleration histories were converted from the model domain to the prototype domain using the scaling conversion factors as mentioned in Table 1. Tables 5 and 6 shows the experimentally obtained, first and second floor level acceleration and displacement time histories for the Model-1. Both tables show the time histories records for the selected significant runs in the prototype domain. For obtaining the relative displacements of each floor relative to the base of the pad, the displacement histories of the base pad were subtracted from each floor level displacements. For each test run, the peak values of displacement were obtained and were normalized by the height of the storey to obtain the corresponding first and second floor drifts. To calculate the floor inertial forces at each floor level as well as the total base shear force at base of the prototype model, each floor accelerations were multiplied with each floor total mass. This mass included; the additional block
mass, self-weight of the slab, self-weight of the beam and half column above and below the floor level. The inertial forces at each floor level was added to obtain the total base shear force at the base of frames. 3.4. Results and discussions Table 7 shows the code compliant Model-1 maximum roof displacement, maximum drift ratio, maximum base shear force and observe damage mechanism for the selected significant runs. For all other models observed damage mechanism for the significant runs are provided in Appendix B. Fig. 7 shows all model’s comparison at the final run i.e. incipient collapse stage and joint panel damage mechanisms. The code compliant Model-1 was initially excited by shake table self-check run which has forced the structure to a drift of about 1.88% and with shaking intensity of 0.60 g. During this first run the model developed significant beam flexural cracks at the first storey level. This flexural cracking was due to the reinforcement rebar yielding and plastic hinge mechanism. During this run, on the ground storey at columns bases and on the first storey at beam ends flexural cracks have been observed. The model was then subjected to multiple excitations from 5% to 100%, with slight increase in cracking pattern. After this the test model was subjected to 130% of the maximum acceleration record during which the model experience about 1.06 g shaking intensity.
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Shake table investigations on code non-compliant reinforced concrete frames Table 6
9
Model-1 experimentally observed displacement histories (prototype domain).
Run
Floor level
Displacement History
Ground floor level
Self-Check Run First floor level
Ground floor level
100% Run First floor level
Ground floor level 130% Run
First floor level
During this run the previous damages got significantly aggravated. The test model experienced concrete crushing and core spalling at the base and top ends of the columns on the ground storey due to excessive compressive strain demand on the cover concrete. Minor spalling was also observed at the base of columns on the first storey. Additionally, the model was observed with severe diagonal cracks in the joint panel region on the ground storey and slight diagonal cracks in the joint region on the first storey, which was due to transferring moments from beam-ends to columns’ ends. This damage pattern identified to the existence of materials’ over-strength in beams that resulted in plastic section moment capacity higher than the yield moment capacity, consequently, increasing demands on the joint region. In comparison to Model-1, Model-2 deformed laterally to larger roof drift under similar input excitations. Unlike Model-1, Model-2 experienced damages in joints quite earlier and to extreme extend under significantly lower excitations. This is due to the fact of using low strength concrete in SMRFs. This reduces the steel-to- concrete bond strength and allows steel bars slip through concrete, consequently resulting in larger displacement of the model. Furthermore, the joint panels damaged under less shear demands (in transferring beam moments to columns) due to the lower principal
tensile strength of the joint panel. Since, the joint principal strength capacity primarily depends on the strength of core concrete that is related to the compressive strength of concrete. Model-3 behaved somewhat similar to Model-2 under dynamic shake table excitations. However, damage to structural members (beam and columns) and beam- column joint panels occurred at comparatively lower lateral drift demand. Unlike Model-2, the damage evolution has shown that damages in Model-3 were more limited to the joint region than the columns and beams. Joint cracks in Model-2 were spread over larger area (joint panel core and transverse beams), whereas joint cracks in Model-3 spread primarily within the joint panel core. Model-3 reached the incipient collapse state at lateral roof drift demand lower than the Model-2. All these are due to the fact that Model-3 had no confining ties in the joint panel region, besides having concrete with low strength. Model 4 behaved very much similar to Model-3 in terms of the damage pattern, however, the occurrence of damage and their severity were observed in Model-4 at lateral roof drift demand comparatively lower than the drift demand experienced by Mode-3. Model-5 also behaved very much similar to Model-3 and Model-4 (more like Model 4) in terms of the damage pattern. However, it can be seen that the model’s lateral force is relatively less than the Model-4.
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4. Seismic performance assessment of test frames 4.1. Lateral force – deformation capacity curves for prototype structure The lateral force-deformation capacity curves for the corresponding prototype structures were calculated by first transforming the model recoded data to the prototype using the conversion factors (model-to-prototype) as per the scaling requirements. The maximum roof displacement in each test run was obtained and normalized by the total height of the model to calculate the corresponding roof drift. The floor accelerations were first multiplied by the floor masses (including the additional block mass, self-weight of slab and beams on the floor and half mass of the columns above/below the floor), to calculate the floor inertial forces. The floor inertial forces were summed to calculate the base shear force. For each test run the maximum roof drift and maximum base shear force were identified and correlated to obtain the base shear force and roof drift capacity envelope curve, as shown in Fig. 8(a). For the relative comparison and calculation of the seismic response parameters of the considered structures, first the lateral force-deformation capacity curves of the structures were bi-linearized using the energy balance criterion. The maximum displacement response observed during the test were considered as the ultimate displacement capacity of the models. The idealized yield strength and yield displacement were calculated through iterative procedure, to equalize the area under the curve for the idealized elasto-plastic capacity curve response to that of area under the actual force-deformation curve, shown in Fig. 8(b). It can be observed that generally the yield stiffness, yield strength and displacement ductility of structure reduced with the inclusion of construction defects. The structure yield stiffness reduced by 40% and yield strength by 16% with the use of concrete having strength 33% less than the design specification (Model 2). The stiffness and strength reduction increased to about 50% and 31% respectively, in case the structure was not provided with confining ties in the beam-column joint panels besides the use of low strength concrete (Model 3). In case of not providing confining ties in beam-column joints and increasing the stirrups spacing by 100%, the yield stiffness reduced by about 46% and yield strength reduced by 35% (Model 4). In addition to the above, if the longitudinal reinforcement were reduced in beams (by 33%) and in columns (by 25%), the yield stiffness reduced by about 60% and yield strength reduced by about 50%. 4.2. Seismic response modification factor for prototype structures Generally, R factor for a structure can be calculated knowing the inelastic lateral force-deformation behavior of the structure. R¼
Ve Ve Vy ¼ ¼ Rl Rs Vs Vy Vs
where Ve represents the elastic force the structure will experience, if respond elastically under earthquake demand; Vy represents the idealized yield strength of the structure; Vs represents the design base shear force; Rl represents the
‘ductility factor’, structure ductility dependent factor, RS represents the ‘overstrength factor’, structure overstrength dependent factor. The overstrength factor RS is calculated directly from the lateral force-deformation capacity curve of the structure by dividing the idealized yield strength over the structure design strength, however, the ductility factor Rl is related to the structural ductility as given [34]. Short Period Intermediate Period Long Period
T < 0.20 s 0.2 s < T < 0.5 s T > 0.5 s
Rl ¼ 1:0 pffiffiffiffiffiffiffiffiffiffiffiffiffiffi Rl ¼ 2l 1 Rl ¼ l
Structure Vibration Period: qffiffiffiffi T ¼ 2p kmy
where T is the pre-yield vibration period of idealized single degree of freedom system. The weight of the considered prototype frame is 28 ton, and considering the yield stiffness obtained from the experimental idealized capacity curves, the structure vibration period is calculated using classical formula of vibration period: Model 1 (0.71 s), Model 2 (0.91 s), Model 3 (1.02 s), Model 4 (0.97 s), Model 5 (1.15 s). The code specified ultimate drift limit of 2.5% is considered as the ultimate drift capacity that corresponds to displacement capacity of about 183 mm (7.20 in.). The frame ductility l is obtained dividing the ultimate displacement capacity over the idealized yield displacement capacity of each structure model, which gives also an estimate of Rl. The response modification factor R of prototype structures is calculated by multiply the ductility dependent Rl factor with the overstrength factor RS. Table 8 reports the calculated structural parameters for all the models. The SMRF Model 1 designed and constructed to the code recommendation was observed with beam-sway mechanism, forming plastic hinges at the beam-ends and base of columns on the ground story. The response modification factor obtained for the model can be approximated as 7.5, which is however about 10% less than the UBC-97 specified 8.5 for the SMRFs. Overall, the seismic performance of Model-1 is very satisfactory and as per the code presumptions. Model 2 that employed concrete in construction with strength 33% lower than the design specified, resulted in R factor 4.5. Model 3 & Model 4 resulted in further reduced R factor of 3.5 and 4.0, respectively, which is due to the lack of confining ties in beam-column joint panels and reducing transverse reinforcement in beam-column members besides the use of low strength concrete. For Model 5, the R factor further reduced to 2.5, due to reducing longitudinal reinforcements in beam-column members in addition to the above defects. 4.3. Seismic performance assessment of code-compliant and noncompliant frames Most of the seismic codes suggest static force-based procedure for the seismic analysis and design of structures. The BCP-SP (2007), which is based on the UBC-97, included equations for calculating base shear force: V ¼ CS W where, CS ¼
CV I RT
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Shake table investigations on code non-compliant reinforced concrete frames Table 7
11
Observed damages in Model-1. Run
Self-Check– 0.60 g
100% – 0.62 g
130% – 1.06 g
Top sotrey maximum dispalcement Inch (mm)
2.42 (61.45)
5.26 (133.56)
14.69 (373.03)
Top sotrey maximum Drift (%)
0.87
1.88
5.26
MaximumBase shear Forrce kips (kN)
33.96 (151.08)
42.47 (188.90)
57.27 (254.73)
Flexure Cracks at Base of Columns
Concrete Crushing at Ground Storey Column Top
Flexure Horizontal & Vertical Cracks in Beam
Cover Spalling at Ground Storey Column Base
Diagonal Cracks in Joint Panel, First Storey
Damages Observed in Beam and Columns on the Ground Storey
Damages Observed in Beam and Columns on the Ground Storey
Damages Observed in Joint Panels on the Ground & First Storey
Diagonal Cracks in Joint Panel, Ground Storey
Observe Damage
However, it shouldn’t be greater than the maximum allowed; CS;max ¼
2:5Ca R
where, W = weight of the structure CS = base shear coefficient I = importance factor R = response modification factor Ca and CV = seismic coefficients for seismic zone The above equations were used to calculate seismic design base shear force for both the as-built and retrofitted structure for various seismic zones: 1, 2A, 2B, 3 and 4, and site soil type B ‘‘rock”. For fully non-complaint RC frame, R factor of 3.0 was considered while R factor was approximated to 7.50 for code-confirming frame. The demand on structures was calculated in terms of base shear coefficient (CS, Demand), which was compared with the design level base shear coefficient
(CS, Capacity) of the structures. CS, Capacity is the design level base shear coefficient for which the considered frames were designed. Tables 9 and 10 reports the factor of safety (FoS) for both the fully non-complaint and code compliant frames. It can be observed that the fully non-complaint RC frame can perform better in all seismic zones, however, the structure will not be able to perform better in seismic zone 3 and 4. On the other hand, the code-conforming frame can preform satisfactorily in all seismic zone with significantly higher factor of safety. 5. Conclusions and recommendations The following are concluded on the basis of experimental shake table tests performed on code-compliant and noncompliant/deficient reinforced concrete frames: The code-compliant special moment resisting frame model developed plastic hinges at the ends of first-floor beam and at the base of columns on ground floor. This model was capable to resist input excitation 1.30 times the design
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M. Rizwan et al.
Ground storey Joints Front and back Joints
First storey Joints Front and back Joints
Model-1 at 1
Model-2 at
Model-3 at
Model-4 at
Model-5 at
Final/Incipient
30% Run
90% Run
40% Run
30% Run
50% Run
collapse Run
1.06
0.59
0.73
0.44
0.49
PGA, g
14.69 (373.03)
19.56 (496.95)
13.31 (338.19)
9.63 (244.57)
10.93 (277.60)
Displacement, in (mm)
5.26
7.01
4.77
3.45
3.92
Drift, %
57.27 (254.73)
33.68 (149.83)
41.48 (184.50)
34.24 (152.33)
28.07 (124.84)
Base shear, kips (kN)
Fig. 7
Compassion of all the test model at final/incipient collapse stage.
base earthquake i.e. up to 1.0 g. Despite the less redundant structural system, the single-bay frame performed significantly well under seismic excitation representative of rare and very rare earthquakes. Shake table tests performed on deficient reinforced concrete frames have revealed that such frames exhibit mixed mechanism of column/beam plastic hinging under seismic excitation. This was followed by extreme damages in beamcolumn joints. This was due to the use of concrete having low strength and beam-column joints lacking confining ties. The incipient collapse state of such frame may reach at a
peak ground acceleration of 0.50 g, however, the joint damage state, and their reparability issues that the engineers may face following a damaging earthquake, will dictate the performance states of this structure type that may reach earlier than the incipient collapse state. Beam members experienced both horizontal and vertical flexure cracks at the ends that prompted beam fixed-end rotation. This indicates that beam longitudinal reinforcement may be subjected to bar-slip due to steel-concrete bond failure. In ACI-318 the side facing/skin reinforcements are generally recommended in deep beams having
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Shake table investigations on code non-compliant reinforced concrete frames
13
300
Base Shear Force (kN)
250
200
150 Model-1 100
Model-2 Model-3
50
Model-4 Model-5
0 0
100
200
300
400
500
Roof Displacement (mm)
(a) Experimentally Derived Capacity Curves Fig. 8
Table 8 Model Type Model Model Model Model Model
1 2 3 4 5
Table 9
(b) Idealized Elasto-Plastic Capacity Curve
Lateral force-deformation capacity curves for prototype structures.
Seismic response parameters for prototype structures. Yield Stiffness kips/inch (kN/m)
Displacement Ductility
Overstrength based RS Factor
Ductility based Rl Factor
Total R Factor Actual (Approx.)
12.38 (2167.50) 7.61 (1332.30) 6.06 (1061.93) 6.71 (1174.08) 4.77 (835.31)
1.83 1.35 1.31 1.52 1.35
4.13 3.45 2.83 2.68 2.15
1.83 1.35 1.31 1.52 1.35
7.54 4.64 3.70 4.09 2.91
(7.50) (4.50) (3.50) (4.00) (2.50)
Seismic performance of as-built RC frame in various seismic zones.
Zone
Ca
CV
Cs, Demand (R = 3.0)
Cs, Capacity
FoS
Remarks
1 2A 2B 3 4
0.08 0.15 0.2 0.3 0.4
0.08 0.15 0.2 0.3 0.4
0.06 0.11 0.14 0.22 0.29
0.20
3.45 1.84 1.38 0.92 0.69
OK OK OK N.G. N.G.
Table 10
Seismic performance of haunch retrofitted RC frame in various seismic zones.
Zone
Ca
CV
Cs, Demand (R = 7.50)
Cs, Capacity
FoS
Remarks
1 2A 2B 3 4
0.08 0.15 0.2 0.3 0.4
0.08 0.15 0.2 0.3 0.4
0.03 0.05 0.06 0.10 0.13
0.20
7.88 4.20 3.15 2.10 1.58
OK OK OK OK OK
effective depth greater than 36 in. (915 mm). However, the observed behavior also suggests the use of side facing/skin reinforcement in beam having depth equal to 15.75 in. (400 mm) to control the spread of vertical cracks at the beam-ends. The following are concluded on the basis of analytical investigation performed on the experimentally tested compliant and non-compliant frames:
The bi-linearized lateral force-displacement response of tested frame revealed that the structural lateral stiffness reduced up to 60% due to non-compliance of frame construction. The corresponding lateral load resistance reduced up to 50% and the corresponding translation displacement ductility ratio reduced up to 30%. The overstrength factor of 4.13 was calculated in case of code-compliant frame that reduced to 2.15 in case of fully non-complaint frame, exhibiting a reduction of about
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14 50%. This indicates the code-compliant frame possesses significant overstrength in comparison to the code-conforming overstrength of 3.0. The ductility factor of 1.83 was calculated in case of codecompliant frame that reduced to 1.35 in case of fully noncomplaint frame, exhibiting a reduction of about 26%. This indicates both the compliant and non-complaint frames possesses ductility lower than the ductility of codeconforming frame of 2.67. This is because of the joint mechanism of tested frame, and less redundant structural system, that hampers the ductile response of frame. The response modification factor in case of code-complaint frame is 7.54 that reduced to 2.91 in case of fully noncompliant frame, exhibiting a reduction of 60% in response modification factor of frame due to non-compliance of constructions. Realizing the fact that one-bay frame is less redundant where reduction factor of 0.80 is proposed in some international codes e.g. Mexico (Tena-Colunga and CortesBenitez, 2015) [35], the response modification factor calculated here (i.e. 7.54) can increase to 9.42 (7.45/0.8 = 9.42) in case of frame having four lines of vertical columns for carrying lateral load. In such case, the computed response modification factor is larger than the value of 8.50/8.0 given in the old/new seismic code, respectively. This indicates the beneficial role of seismic detailing given in the ACI-318 seismic provisions.
M. Rizwan et al. Appendix A. Input Protocol and test sequence for nomcompliant models
Model-2 Time History
Run (%)
Observed Input PGA (g)
1994 Northridge Earthquake
Self-Check Run 5% Run 10% Run 30% Run 40% Run 50% Run 60% Run 70% Run 80% Run 90% Run
0.30 0.048 0.29 0.22 0.37 0.38 0.48 0.55 0.88 0.59
Run (%)
Observed Input PGA (g)
Self-Check Run 5% Run 10% Run 20% Run 30% Run 40% Run
0.015 0.52 0.25 0.31 0.35 0.73
Run (%)
Observed Input PGA (g)
Self-Check Run 5% Run 10% Run 20% Run 30% Run
0.013 0.33 0.88 0.27 0.44
Run (%)
Observed Input PGA (g)
Self-Check Run 5% Run 10% Run 20% Run 30% Run 40% Run 50% Run
0.06 0.10 0.07 0.11 0.20 0.29 0.49
Model-3 Time History
1994 Northridge Earthquake
The following are concluded on the basis of static force procedure used for the seismic performance assessment of codecompliant and non-compliant/deficient reinforced concrete frames in various seismic zones: The code-conforming frame is capable to resist the design base earthquakes in highest seismic zone without exceeding the code permissible story drift limit of 2.50%. The fully non-complaint frame is capable to resist the design base earthquakes only in low to moderate seismic zones but fails to resist design base earthquakes in highest seismic zones (i.e. zone 3 and zone 4) without exceeding the code permissible story drift limit of 2.50%.
Model-4 Time History
1994 Northridge Earthquake
The present research suggests taking in to account the reliability/redundancy factor in proposing response modification factor for seismic design of structures. This factor will take care of the deficiencies encountered in the regional construction practices, and will vary from 1.0 (in case of compliance) to 0.4 (in case of fully non-compliance). Model-5
Declaration of Competing Interest
Time History
The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
1994 Northridge Earthquake
Acknowledgments The authors are grateful to the anonymous reviewers for their constructive remarks that improved the quality of manuscript.
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Shake table investigations on code non-compliant reinforced concrete frames
15
Appendix B. See Tables B1–B4.
Table B1
Observed damages in Model-2. Run
70%– 0.55 g
80% – 0.88 g
90% – 0.59 g
Top sotrey maximum dispalcement Inch (mm)
9.30 (236.26)
14.18 (360.11)
19.56 (496.95)
Top sotrey maximum Drift (%)
3.33
5.08
7.01
MaximumBase shear Forrce kips (kN)
42.39 (188.54)
40.70 (181.02)
33.68 (149.83)
Slight Diagonal Cracks in Joints on Ground Storey
Severe Diagonal Cracks in Joints on Ground Storey
Diagonal Cracks in Joint Panel, Ground Storey
Significant Flexure Cracks in First Storey Columns
Concrete Wedge Detachment from First Storey Joint
Damages Observed in Columns and Joint Panels during 70% Run
Damages Observed in Columns and Joint Panels during 80% Run
Observe Damage
Bat-Like Wedge Detachment from First Storey Joint Damages Observed in Joint Panels of Ground and First Storey during 90% Run
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M. Rizwan et al.
Table B2
Observed damages in Model-3. Run
5%– 0.52g
30% – 0.35 g
40% – 0.73 g
Top sotrey maximum dispalcement Inch (mm)
4.87 (123.69)
7.18 (182.365)
13.31 (338.19)
Top sotrey maximum Drift (%)
1.75
2.57
4.77
MaximumBase shear Forrce kips (kN)
26.51 (117.92)
30.91 (137.48)
41.48 (184.50)
Flexure Cracks in Beams and Columns, Ground Storey
Severe Bat-Like Cracks in Joints on First Storey
Slight Cracks in Joint Panel on Ground Storey
Damage Progress in Joint Panel on Ground Storey
Damages Observed in Columns, Beam
Damages Observed in Joint Panels of Ground and First Storey during 30% Run
Cover Detachment and Damage in Joint on First Storey
Observe Damage
and Joint Panels during 5% Run
Severe Damage to Joint Panel on Ground Storey Damages Observed in Joint Panels of Ground and First Storey during 40% Run
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Shake table investigations on code non-compliant reinforced concrete frames Table B3
17
Observed damages in Model-4. Run
5%– 0.33g
10% – 0.88 g
30% – 0.44 g
Top sotrey maximum dispalcement Inch (mm)
3.15 (79.97)
5.91 (150.23)
9.63 (244.57)
Top sotrey maximum Drift (%)
1.13
2.12
2.12
MaximumBase shear Forrce kips (kN)
22.87 (101.72)
30.53 (135.82)
30.53 (135.82)
Flexure Cracks in Beam and Column on Ground Storey
Severe Bat-Like Cracks in Joints on First Storey
Cracks in Joint Panel on Ground Storey
Damage Progress in Joint Panel on Ground Storey
Cover Detachment and Spalling in Joint on First Storey
Observe Damage
Damages Observed in Columns, Beam and Joint Panels during 5% Run
Damages Observed in Joint Panels of Ground and First Storey during 10% Run
Significant Damage to Joint Panel on Ground Storey Damages Observed in Joint Panels of Ground and First Storey during 30% Run
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M. Rizwan et al.
Table B4
Observed damages in Model-5. Run
5%– 0.10g
30% – 0.20 g
50% – 0.49 g
Top sotrey maximum dispalcement Inch (mm)
1.01 (25.65)
3.66 (92.95)
10.93 (277.60)
Top sotrey maximum Drift (%)
0.36
1.31
3.92
MaximumBase shear Forrce kips (kN)
8.08 (35.96)
18.73 (83.31)
28.07 (124.84)
Flexure Cracks in Beam and Column on Ground Storey
Concrete Crushing at Column Top Ends, Ground Storey
Extend of Damage to Joints on Ground Storey
Cracks in Top Ends of Columns on First Storey
Cover Concrete Detachment from Joints on First Storey
Extend of Damage to Joints on First Storey
Damages Observed in Beams and Columns during 5% Run
Damages Observed in Column Ends on Ground Storey and Joint Panels of First Storey during 30% Run
Damages Observed in Joint Panels of Ground and First Storey during 50% Run
Observe Damage
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H O S T E D BY
Alexandria University
Alexandria Engineering Journal www.elsevier.com/locate/aej www.sciencedirect.com
ORIGINAL ARTICLE
Shake table investigations on code non-compliant reinforced concrete frames Muhammad Rizwan *, Naveed Ahmad, Akhtar Naeem Khan, Samiullah Qazi, Junaid Akbar, Muhammad Fahad Earthquake Engineering Center, Department of Civil Engineering, University of Engineering & Technology, Peshawar, Khyber Pakhtunkhwa 25120, Pakistan Received 5 November 2019; revised 26 December 2019; accepted 29 December 2019
KEYWORDS Shake table test; SMRF; ACI-318; Damage scale; Response modification factor; Beam-column joints
Abstract Shake-table tests were performed on a code conforming and four non-compliant 1:3 reduced scale two-story RC frames. The models were subjected to 1994 Northridge earthquake acceleration time history, linearly scaled to multiple levels for studying the progressive damages. The code conforming model was observed with plastic hinges at the beam-ends and base of columns on ground story. This frame incurred only slight cracks in the joint panels under extreme shaking. The code-conforming model resisted input excitation with peak horizontal acceleration of 1.0 g before attaining the incipient collapse state. The non-compliant models were observed with plastic hinges in beam/column members, and incurred extensive damages in beam-column joints under extreme shaking. The fully non-compliant frame resisted peak horizontal acceleration of 0.50 g before attaining the incipient collapse state. The response modification factor of codeconforming model is 7.54 that reduced to 2.90 in case of fully non-compliant frame, which was due to the non-compliance of constructions. This was partly due to the reduction in the overstrength of frame (which was about 20%) and partly due to the reduction in ductility of frame (which was about 50%), in comparison to the seismic response parameters given in the seismic code. The less redundant structural system and joint shear hinging are the main reasons resulting in the reduction of overstrength and ductility, respectively. Static force based performance assessment in various seismic zones revealed that the code-conforming frame performance is ‘‘OK” while noncomplaint frames performance is ‘‘NG” in both seismic zone 3 and zone 4. The present research suggests a reliability/redundancy factor to take in to account in proposing response modification factor for seismic design of structure. This will take care of the construction deficiencies found in the region. This factor can vary from 1.0 (in case of fully compliant structure) to 0.40 (in case of fully non-compliant structure). Ó 2020 Faculty of Engineering, Alexandria University. Production and hosting by Elsevier B.V. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/).
* Corresponding author. E-mail address: [email protected] (M. Rizwan). Peer review under responsibility of Faculty of Engineering, Alexandria University. https://doi.org/10.1016/j.aej.2019.12.047 1110-0168 Ó 2020 Faculty of Engineering, Alexandria University. Production and hosting by Elsevier B.V. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/). Please cite this article in press as: M. Rizwan et al., Shake table investigations on code non-compliant reinforced concrete frames, Alexandria Eng. J. (2020), https:// doi.org/10.1016/j.aej.2019.12.047
2 1. Introduction Reinforced concrete construction is on the peak in many urban areas of the world. This type of construction is particularly used for multi-storey buildings, schools, hospitals, and residential type buildings because of vast availability of constituents of concrete around the world and ease of construction. Although the design and construction of RC buildings in the developing countries are mostly based on the locally available or adopted international building codes, the proper construction execution of the specified design in the field is still a big challenge. The improper construction practices have resulted in many substandard and vulnerable building stock [1,2], with recent surveys have shown a number of construction/design deficiencies [1,3]. The commonly available construction/design defects include; substandard quality of concrete (low strength concrete), empty joints (joint panels without transverse reinforcement), beam column members having reduced flexural reinforcement along with shear reinforcement spacing larger than the code specified and practicing non-seismic hooks with among others. It is also worth to mention, majority of these RC buildings in the developing countries are located in high seismic regions and can be subjected by future large earthquakes of extreme excitations [4]. Reinforced concrete structures if not built properly can result in catastrophic failure and subsequent human and economic losses, upon subjecting to earthquake induced strong ground motions [5–8]. In the recent decades, many researchers have performed experimental investigations through shake table testing on RC frame structures. In most of these experimental studies, either full scaled and/or reduced scaled structures have been employed and the work were mostly focused on the vulnerability assessment of gravity type/old type RC buildings and/or to assess the performance of a specific retrofitting/strengthening/ isolation technique [9–29] However, experimental shake table investigations on the recently employed SMRF structures with and without the aforementioned construction deficiencies are lacking. In the current study, detail experimental shake table investigations have been performed on RC SMRF structures complying to code requirements i.e. code compliant frame and with construction/design defects i.e. code non-compliant frames. The shake table tests investigations have been performed on five one-third reduced scaled two storey RC frame specimens. Model-1 was a code compliant model (reference model) design based on the seismic building code, whereas all other models were code non-compliant specimens with the inclusion of construction/design defects systemically. The deficiencies which are included in the current study are; substandard quality of concrete (low strength concrete), empty joints (joint panels without transverse reinforcement), beam column members having reduced flexural reinforcement along with shear reinforcement spacing larger than the code specified and practicing non-seismic hooks. All the test specimens were excited with multiple excitations using the natural accelerogram of 1994 Northridge Earthquake, ranges from 5% to 130% of peak ground acceleration. The idea of using multiple scaled excitations is to force the test models from elastic stage to inelastic and finally near collapse stage. The damage mechanism of each
M. Rizwan et al. specimen is observed and reported. Acceleration and displacement response of the structure is recorded and analyzed to obtain the inter-storey drift demand, inter-storey shear and displacement profile of the structure and to develop deformation-based damage scale for seismic assessment of the considered structures. 2. Experimental program 2.1. Details of structures under investigation The current research focuses on low rise RC frame structures with and without construction/design defects, in order to quantify their seismic performance. In particular, an RC frame with two storey structure is considered, typically employed for low rise schools, hospital or apartment type public buildings. The considered structure consists of 2 by 1 bay frame, with each bay length of 18 feet (5487 mm) and storey height of 12 feet (3658 mm) for both the stories, as shown in Fig. 1. The representative frame structure was design according static force-based procedures (BCP-SP 2007/UBC 97) [30,31], considering high seismic zone of 4.0 (0.40 g design peak ground acceleration) with stiff soil type B (NEHRP classification). The modeling and design were carried out using the design software CSI ETABS considering all load combinations as per the code. For the material properties, 3000 psi (21 MPa) concrete and 60,000 psi (414 MPa) rebar yield strength was considered in the design process. The design model was detailed as per the ACI-318–14 [32] recommendations for the SMRF provisions. Fig. 1 shows the design of the prototype structure. 2.2. Construction defects under investigation As already mentioned, the improper construction and execution of the design specifications in the field is still a major challenge in many developing countries, which has caused many deficient building stock [1,2]. The most commonly available deficiencies include; substandard quality of concrete i.e. low strength concrete, empty joints i.e. joint panels without transverse reinforcement, beam column members having reduced flexural reinforcement along with shear reinforcement spacing larger than the code specified and practicing non-seismic hooks. In the current experimental campaign, these construction/design defects have been considered systematically. Model-1 is a reference test model and is based on the code design requirements. Model-2 is similar to Model-1 but provided with low strength concrete of 2000 psi (14 MPa). Model-3 is similar to Mode-1 and Model-2, but along with low strength concrete, no transverse ties have been provided in the beam-column joint panel. Similarly, Model-4 and Model-5 were similar to the Model-3 but were provided with beam and columns transverse reinforcement spacing were 100% (6 in. [152 mm] c/c) and 200% (9 in. [228 mm] c/c) larger than the previous models, respectively. In Model-5 the beam flexural reinforcement was also reduced from (3 + 3) # 6 bars to (2 + 2) # 6 bars and ties were provided with non-seismic 90° hook in both beams and columns members.
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Shake table investigations on code non-compliant reinforced concrete frames
3
Beam secon A-A Fig. 1
Prototype geometric layout and design specifications.
2.3. Preparation of one-third reduced models The seismic shake table simulator of Earthquake Engineering Center Civil Engineering Department has a table size of 5 feet by 5 feet (1.5 m 1.5 m) and can operate with loading capacity of about 5 tons only in the unidirectional excitation. Due to size and loading capacity limitations, only the interior critical frame was extracted from the prototype design building and reduced by one-third scale for seismic excitation. For scaling between the prototype and model dimension, a simple linear
Table 1
Column secon B-B
model idealization was considered [16,33], as shown in Table 1. Such simple modeling allows simplicity and also reduced the cost as well as complexities associated in scaling the stressstrain properties of the materials (concrete and reinforcing bar) in the model domain. As shown in Table 1, all the members of the extracted prototype frame i.e. beam, column and slab, and reinforcing rebar were reduced by a scale factor of SL 3. For the constituents of concrete, 3/8in down aggregate size have been employed for model preparation, in order to used scaled coarse aggregate, whereas cement and fine aggre-
Similitude conversion factors and reduced models’ dimensions.
Structural Properties
Simple Model Similitude Requirement
Prototype Frame
Test Models (Scale 1:3)
Physical Quantity
Relationship
Scale Factor
Beams: 12in 18in (304 mm 459 mm) Columns: 12in 12in (304 mm 304 mm) Slab: 6in (153 mm) Concrete Strength: 3000 psi (21 MPa) 2000 psi (14 MPa) Aggregate size: 1in Steel Strength: 60,000 psi (414 MPa) #6 Rebar (19 mm) #3 Rebar (10 mm)
Beams: 4in 6in (102 mm 153 mm) Columns: 4in 4in (102 mm 102 mm) Slab: 2in (51 mm) Concrete Strength: 3000 psi (21 MPa) 1:1.80:1.60, W/C = 0.48 2000 psi (14 MPa) 1:3.50:2.87, W/C = 0.80 Aggregate size: 3/8in Steel Strength: 60,000 psi (414 MPa) #2 Rebar (6.33 mm) #1 Rebar (3.33 mm)
Length Stress Strain Specific Mass Displacement Force Time Frequency Velocity Acceleration
SL = Lp/Lm Sf = fp/fm Sє=єp/єm Sq = qp/qm Sd = dp/dm = SL SF = Fp/Fm = S2LSf p St = tp/tm = SL (SєSq/Sf) SX = Xp/Xm = 1/St p Sv = vp/vm= (SєSq/Sf) Sa = ap/am = Sf/SLSq
3 1 1 1 3 9 3^0.5 (1/3)^0.5 1 1/3
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Fig. 2
Test frames construction stages.
gate have been used with no scaling. As mention earlier two types of concrete strength have been employed in this study i.e. 3000 psi (21 MPa) and low strength test models with a 2000 psi (14 MPa) concrete compressive strength. For this purpose, ACI concrete mix design methods were followed for the
preparation of concrete constituent mix. The concrete mix proportion of cement:sand:corase aggregate, for both design specified strength of 3000 psi (21 MPa) and reduced strength of 2000 psi (14 MPa) are reported shown in Table 1. Fig. 2 shows the sequence of construction for the one-third scaled frame
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Shake table investigations on code non-compliant reinforced concrete frames
Fig. 3
Fig. 4
5
Test frames shifting to testing laboratory.
Added floor masses preparation and setup for artificial mass simulation.
models, Whereas Fig. 3 shows the test model transportation to the testing laboratory. Due to scaling and similitude requirements the reduced scale models were subjected to gravity and seismic mass less than the required. To satisfy the condition of mass simulation
for the reduced scaled test models, additional floor mass have been applied following the mass simulation model [16,33]. On each floor level extra mass of 1200 kg were applied through steel blocks mounted and fixed to the floor by means of fully secured ø inch (13 mm) steel bolts as shown in Fig. 4.
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Fig. 5
Table 2
Test frame instrumentation and shaking setup.
Test model instrumentation positions and characteristics.
Channel
Position
Direction
Maximum Capacity
Parameter
Coefficient
Unit
A1 A2 A3 A4 A5 A6 D7 D8 D9
Pad level First Floor Second Floor Pad level First Floor Second Floor Pad level First Floor Second Floor
Front Front Front Back Back Back Front Front Front
±10 g
Acceleration Acceleration Acceleration Acceleration Acceleration Acceleration Displacement Displacement Displacement
492.20 501.10 510.10 508.90 490.10 502.00 1000.00 1000.30 1000.20
mv/g mv/g mv/g mv/g mv/g mv/g mv/inch mv/inch mv/inch
24 in. (610 mm)
Table 3 Earthquake Engineering Centre (EEC)’s seismic simulator (shake table) limits. Shaking simulator characteristics
Limits
Excitation direction
Single degree of freedom/unidirection 5 feet 5feet (1.5 m 1.5 m) 5 tonnes 1.1 g ±1.1 m/s ±125 mm
Physical dimension Pay load capacity Maximum acceleration Maximum velocity Maximum displacement
frames, so only external instrumentation was employed in the form of six accelerometers and three linear displacement transducers. The instruments sensitivity, maximum capacity and conversion coefficient have been reported in Table 2. On each floor level (mid position of joint panel region) and at base pad level, three accelerometers were installed on front and back side of the test models in order to record floor and base pad accelerations. For recording the displacement at each floor and pad level, three displacement transducers were attached on a fixed steel frame, which was install in-line with the model in-plane position.
3. Shake table tests on two-storey RC frames
3.2. Shake table Input loading protocols
3.1. Testing setup and model instrumentation
All the test models were tested by the Earthquake Engineering Center (EEC)’s seismic simulator in a uni-directional motion. Table 3 reports the characteristics limit values of the shake table. In order to excite the test models from elastic stage to full ultimate collapse stage, and also to be within the shaking
The test model’s setup on the shaking simulator and model instrumentation are shown in Fig. 5. The objective of experimental testing was to observe the global response of the test
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Time history record of 1994 Northiridge Earthquake
5% Damped Acceleration Response Spectra Fig. 6
Table 4
5% Damped Displacement Response Spectra
Input time history record.
Input protocol and testing sequence for Model-1.
Time History
Run (%)
Observed Input PGA (g)
1994 Northridge Earthquake
Self-Check 5% Run 10% Run 20% Run 30% Run 40% Run 50% Run 60% Run 70% Run 80% Run 90% Run 100% Run Self-Check 130% 130% Run
0.60 0.033 0.06 0.12 0.16 0.19 0.25 0.31 0.36 0.41 0.49 0.62 0.62 1.06
range of shaking simulator, the 1994 Northridge Earthquake natural acceleration record was selected. This selection was made after careful analysis of a number of natural acceleration records to be within the range of shake table limiting acceleration, velocity and displacement limits. Fig. 6 shows the horizontal component of 1994 Northridge Earthquake record acceleration time history, elastic accelerations and displacement spectra for 5% damping. The record has been obtained from PEER strong motion data base (090 CDMG Station 24278) and has a peak acceleration, velocity
and displacement values of 0.57 g, 20.39 in/sec (518 mm/sec) and 3.54 in (90 mm), respectively. The acceleration record time step was reduced by 1/3 to respect the input frequency requirement for the model as mentioned in Table 1. Table 4 reported the input multiple excitations and test sequence for code compliant Model-1. Each frame model was excited with incremental excitations ranges from 5% to 130% depending on the ultimate capacities of the tested frame. Input multiple excitations and test sequences for all other models are reported in Appendix A. The idea of using multiple scaled excitation i.e. from low level to high level, was to deform and force the test structures from elastic stage to inelastic and then to attain full ultimate incipient collapse stage. The EEC’s seismic simulator performs a self-adjustment motion called Self-Check, once the input time history is given to the shake table. After the simulator self-check adjustment, the test models were excited with scaled incremental excitations of the maximum acceleration records. After each run the damage mechanism of the test models was observed and documented with snapshot. The test sequence was progressively increase until the test models reached to near incipient collapse stage, after which the test was concluded. The recorded acceleration and displacement response time histories were obtained for each test run in the form of voltage values. 3.3. Recorded data processing The accelerometers and displacement transducers record data in the form of current values (voltage, mv). To get the time histories values in the form of accelerations (g) and displacements
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M. Rizwan et al. Table 5
Model-1 experimentally observed accelerations histories (prototype domain).
Run
Floor level
Acceleration History
Ground floor level
Self-Check Run
First floor level
Ground floor level 100% Run First floor level
Ground floor level 130% Run First floor level
(mm), the recorded raw data needs to be divided by the instruments conversion coefficient as reported in Table 2. Once the recorded data have been corrected for the respective instrument coefficient, the raw data was further processed for base line correction and signal filtering. This data processing correction and filtering was done to remove any noise in the actual recorded data. For this purpose, the SeismoSignal (SeismoSoft 2015) data processing software have employed with a linear category base line correction and filter type of Butterworth with frequency range of 0.10–25 Hz were considered. Once the processed data have been obtained, the displacement and acceleration histories were converted from the model domain to the prototype domain using the scaling conversion factors as mentioned in Table 1. Tables 5 and 6 shows the experimentally obtained, first and second floor level acceleration and displacement time histories for the Model-1. Both tables show the time histories records for the selected significant runs in the prototype domain. For obtaining the relative displacements of each floor relative to the base of the pad, the displacement histories of the base pad were subtracted from each floor level displacements. For each test run, the peak values of displacement were obtained and were normalized by the height of the storey to obtain the corresponding first and second floor drifts. To calculate the floor inertial forces at each floor level as well as the total base shear force at base of the prototype model, each floor accelerations were multiplied with each floor total mass. This mass included; the additional block
mass, self-weight of the slab, self-weight of the beam and half column above and below the floor level. The inertial forces at each floor level was added to obtain the total base shear force at the base of frames. 3.4. Results and discussions Table 7 shows the code compliant Model-1 maximum roof displacement, maximum drift ratio, maximum base shear force and observe damage mechanism for the selected significant runs. For all other models observed damage mechanism for the significant runs are provided in Appendix B. Fig. 7 shows all model’s comparison at the final run i.e. incipient collapse stage and joint panel damage mechanisms. The code compliant Model-1 was initially excited by shake table self-check run which has forced the structure to a drift of about 1.88% and with shaking intensity of 0.60 g. During this first run the model developed significant beam flexural cracks at the first storey level. This flexural cracking was due to the reinforcement rebar yielding and plastic hinge mechanism. During this run, on the ground storey at columns bases and on the first storey at beam ends flexural cracks have been observed. The model was then subjected to multiple excitations from 5% to 100%, with slight increase in cracking pattern. After this the test model was subjected to 130% of the maximum acceleration record during which the model experience about 1.06 g shaking intensity.
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Shake table investigations on code non-compliant reinforced concrete frames Table 6
9
Model-1 experimentally observed displacement histories (prototype domain).
Run
Floor level
Displacement History
Ground floor level
Self-Check Run First floor level
Ground floor level
100% Run First floor level
Ground floor level 130% Run
First floor level
During this run the previous damages got significantly aggravated. The test model experienced concrete crushing and core spalling at the base and top ends of the columns on the ground storey due to excessive compressive strain demand on the cover concrete. Minor spalling was also observed at the base of columns on the first storey. Additionally, the model was observed with severe diagonal cracks in the joint panel region on the ground storey and slight diagonal cracks in the joint region on the first storey, which was due to transferring moments from beam-ends to columns’ ends. This damage pattern identified to the existence of materials’ over-strength in beams that resulted in plastic section moment capacity higher than the yield moment capacity, consequently, increasing demands on the joint region. In comparison to Model-1, Model-2 deformed laterally to larger roof drift under similar input excitations. Unlike Model-1, Model-2 experienced damages in joints quite earlier and to extreme extend under significantly lower excitations. This is due to the fact of using low strength concrete in SMRFs. This reduces the steel-to- concrete bond strength and allows steel bars slip through concrete, consequently resulting in larger displacement of the model. Furthermore, the joint panels damaged under less shear demands (in transferring beam moments to columns) due to the lower principal
tensile strength of the joint panel. Since, the joint principal strength capacity primarily depends on the strength of core concrete that is related to the compressive strength of concrete. Model-3 behaved somewhat similar to Model-2 under dynamic shake table excitations. However, damage to structural members (beam and columns) and beam- column joint panels occurred at comparatively lower lateral drift demand. Unlike Model-2, the damage evolution has shown that damages in Model-3 were more limited to the joint region than the columns and beams. Joint cracks in Model-2 were spread over larger area (joint panel core and transverse beams), whereas joint cracks in Model-3 spread primarily within the joint panel core. Model-3 reached the incipient collapse state at lateral roof drift demand lower than the Model-2. All these are due to the fact that Model-3 had no confining ties in the joint panel region, besides having concrete with low strength. Model 4 behaved very much similar to Model-3 in terms of the damage pattern, however, the occurrence of damage and their severity were observed in Model-4 at lateral roof drift demand comparatively lower than the drift demand experienced by Mode-3. Model-5 also behaved very much similar to Model-3 and Model-4 (more like Model 4) in terms of the damage pattern. However, it can be seen that the model’s lateral force is relatively less than the Model-4.
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4. Seismic performance assessment of test frames 4.1. Lateral force – deformation capacity curves for prototype structure The lateral force-deformation capacity curves for the corresponding prototype structures were calculated by first transforming the model recoded data to the prototype using the conversion factors (model-to-prototype) as per the scaling requirements. The maximum roof displacement in each test run was obtained and normalized by the total height of the model to calculate the corresponding roof drift. The floor accelerations were first multiplied by the floor masses (including the additional block mass, self-weight of slab and beams on the floor and half mass of the columns above/below the floor), to calculate the floor inertial forces. The floor inertial forces were summed to calculate the base shear force. For each test run the maximum roof drift and maximum base shear force were identified and correlated to obtain the base shear force and roof drift capacity envelope curve, as shown in Fig. 8(a). For the relative comparison and calculation of the seismic response parameters of the considered structures, first the lateral force-deformation capacity curves of the structures were bi-linearized using the energy balance criterion. The maximum displacement response observed during the test were considered as the ultimate displacement capacity of the models. The idealized yield strength and yield displacement were calculated through iterative procedure, to equalize the area under the curve for the idealized elasto-plastic capacity curve response to that of area under the actual force-deformation curve, shown in Fig. 8(b). It can be observed that generally the yield stiffness, yield strength and displacement ductility of structure reduced with the inclusion of construction defects. The structure yield stiffness reduced by 40% and yield strength by 16% with the use of concrete having strength 33% less than the design specification (Model 2). The stiffness and strength reduction increased to about 50% and 31% respectively, in case the structure was not provided with confining ties in the beam-column joint panels besides the use of low strength concrete (Model 3). In case of not providing confining ties in beam-column joints and increasing the stirrups spacing by 100%, the yield stiffness reduced by about 46% and yield strength reduced by 35% (Model 4). In addition to the above, if the longitudinal reinforcement were reduced in beams (by 33%) and in columns (by 25%), the yield stiffness reduced by about 60% and yield strength reduced by about 50%. 4.2. Seismic response modification factor for prototype structures Generally, R factor for a structure can be calculated knowing the inelastic lateral force-deformation behavior of the structure. R¼
Ve Ve Vy ¼ ¼ Rl Rs Vs Vy Vs
where Ve represents the elastic force the structure will experience, if respond elastically under earthquake demand; Vy represents the idealized yield strength of the structure; Vs represents the design base shear force; Rl represents the
‘ductility factor’, structure ductility dependent factor, RS represents the ‘overstrength factor’, structure overstrength dependent factor. The overstrength factor RS is calculated directly from the lateral force-deformation capacity curve of the structure by dividing the idealized yield strength over the structure design strength, however, the ductility factor Rl is related to the structural ductility as given [34]. Short Period Intermediate Period Long Period
T < 0.20 s 0.2 s < T < 0.5 s T > 0.5 s
Rl ¼ 1:0 pffiffiffiffiffiffiffiffiffiffiffiffiffiffi Rl ¼ 2l 1 Rl ¼ l
Structure Vibration Period: qffiffiffiffi T ¼ 2p kmy
where T is the pre-yield vibration period of idealized single degree of freedom system. The weight of the considered prototype frame is 28 ton, and considering the yield stiffness obtained from the experimental idealized capacity curves, the structure vibration period is calculated using classical formula of vibration period: Model 1 (0.71 s), Model 2 (0.91 s), Model 3 (1.02 s), Model 4 (0.97 s), Model 5 (1.15 s). The code specified ultimate drift limit of 2.5% is considered as the ultimate drift capacity that corresponds to displacement capacity of about 183 mm (7.20 in.). The frame ductility l is obtained dividing the ultimate displacement capacity over the idealized yield displacement capacity of each structure model, which gives also an estimate of Rl. The response modification factor R of prototype structures is calculated by multiply the ductility dependent Rl factor with the overstrength factor RS. Table 8 reports the calculated structural parameters for all the models. The SMRF Model 1 designed and constructed to the code recommendation was observed with beam-sway mechanism, forming plastic hinges at the beam-ends and base of columns on the ground story. The response modification factor obtained for the model can be approximated as 7.5, which is however about 10% less than the UBC-97 specified 8.5 for the SMRFs. Overall, the seismic performance of Model-1 is very satisfactory and as per the code presumptions. Model 2 that employed concrete in construction with strength 33% lower than the design specified, resulted in R factor 4.5. Model 3 & Model 4 resulted in further reduced R factor of 3.5 and 4.0, respectively, which is due to the lack of confining ties in beam-column joint panels and reducing transverse reinforcement in beam-column members besides the use of low strength concrete. For Model 5, the R factor further reduced to 2.5, due to reducing longitudinal reinforcements in beam-column members in addition to the above defects. 4.3. Seismic performance assessment of code-compliant and noncompliant frames Most of the seismic codes suggest static force-based procedure for the seismic analysis and design of structures. The BCP-SP (2007), which is based on the UBC-97, included equations for calculating base shear force: V ¼ CS W where, CS ¼
CV I RT
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Shake table investigations on code non-compliant reinforced concrete frames Table 7
11
Observed damages in Model-1. Run
Self-Check– 0.60 g
100% – 0.62 g
130% – 1.06 g
Top sotrey maximum dispalcement Inch (mm)
2.42 (61.45)
5.26 (133.56)
14.69 (373.03)
Top sotrey maximum Drift (%)
0.87
1.88
5.26
MaximumBase shear Forrce kips (kN)
33.96 (151.08)
42.47 (188.90)
57.27 (254.73)
Flexure Cracks at Base of Columns
Concrete Crushing at Ground Storey Column Top
Flexure Horizontal & Vertical Cracks in Beam
Cover Spalling at Ground Storey Column Base
Diagonal Cracks in Joint Panel, First Storey
Damages Observed in Beam and Columns on the Ground Storey
Damages Observed in Beam and Columns on the Ground Storey
Damages Observed in Joint Panels on the Ground & First Storey
Diagonal Cracks in Joint Panel, Ground Storey
Observe Damage
However, it shouldn’t be greater than the maximum allowed; CS;max ¼
2:5Ca R
where, W = weight of the structure CS = base shear coefficient I = importance factor R = response modification factor Ca and CV = seismic coefficients for seismic zone The above equations were used to calculate seismic design base shear force for both the as-built and retrofitted structure for various seismic zones: 1, 2A, 2B, 3 and 4, and site soil type B ‘‘rock”. For fully non-complaint RC frame, R factor of 3.0 was considered while R factor was approximated to 7.50 for code-confirming frame. The demand on structures was calculated in terms of base shear coefficient (CS, Demand), which was compared with the design level base shear coefficient
(CS, Capacity) of the structures. CS, Capacity is the design level base shear coefficient for which the considered frames were designed. Tables 9 and 10 reports the factor of safety (FoS) for both the fully non-complaint and code compliant frames. It can be observed that the fully non-complaint RC frame can perform better in all seismic zones, however, the structure will not be able to perform better in seismic zone 3 and 4. On the other hand, the code-conforming frame can preform satisfactorily in all seismic zone with significantly higher factor of safety. 5. Conclusions and recommendations The following are concluded on the basis of experimental shake table tests performed on code-compliant and noncompliant/deficient reinforced concrete frames: The code-compliant special moment resisting frame model developed plastic hinges at the ends of first-floor beam and at the base of columns on ground floor. This model was capable to resist input excitation 1.30 times the design
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Ground storey Joints Front and back Joints
First storey Joints Front and back Joints
Model-1 at 1
Model-2 at
Model-3 at
Model-4 at
Model-5 at
Final/Incipient
30% Run
90% Run
40% Run
30% Run
50% Run
collapse Run
1.06
0.59
0.73
0.44
0.49
PGA, g
14.69 (373.03)
19.56 (496.95)
13.31 (338.19)
9.63 (244.57)
10.93 (277.60)
Displacement, in (mm)
5.26
7.01
4.77
3.45
3.92
Drift, %
57.27 (254.73)
33.68 (149.83)
41.48 (184.50)
34.24 (152.33)
28.07 (124.84)
Base shear, kips (kN)
Fig. 7
Compassion of all the test model at final/incipient collapse stage.
base earthquake i.e. up to 1.0 g. Despite the less redundant structural system, the single-bay frame performed significantly well under seismic excitation representative of rare and very rare earthquakes. Shake table tests performed on deficient reinforced concrete frames have revealed that such frames exhibit mixed mechanism of column/beam plastic hinging under seismic excitation. This was followed by extreme damages in beamcolumn joints. This was due to the use of concrete having low strength and beam-column joints lacking confining ties. The incipient collapse state of such frame may reach at a
peak ground acceleration of 0.50 g, however, the joint damage state, and their reparability issues that the engineers may face following a damaging earthquake, will dictate the performance states of this structure type that may reach earlier than the incipient collapse state. Beam members experienced both horizontal and vertical flexure cracks at the ends that prompted beam fixed-end rotation. This indicates that beam longitudinal reinforcement may be subjected to bar-slip due to steel-concrete bond failure. In ACI-318 the side facing/skin reinforcements are generally recommended in deep beams having
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Shake table investigations on code non-compliant reinforced concrete frames
13
300
Base Shear Force (kN)
250
200
150 Model-1 100
Model-2 Model-3
50
Model-4 Model-5
0 0
100
200
300
400
500
Roof Displacement (mm)
(a) Experimentally Derived Capacity Curves Fig. 8
Table 8 Model Type Model Model Model Model Model
1 2 3 4 5
Table 9
(b) Idealized Elasto-Plastic Capacity Curve
Lateral force-deformation capacity curves for prototype structures.
Seismic response parameters for prototype structures. Yield Stiffness kips/inch (kN/m)
Displacement Ductility
Overstrength based RS Factor
Ductility based Rl Factor
Total R Factor Actual (Approx.)
12.38 (2167.50) 7.61 (1332.30) 6.06 (1061.93) 6.71 (1174.08) 4.77 (835.31)
1.83 1.35 1.31 1.52 1.35
4.13 3.45 2.83 2.68 2.15
1.83 1.35 1.31 1.52 1.35
7.54 4.64 3.70 4.09 2.91
(7.50) (4.50) (3.50) (4.00) (2.50)
Seismic performance of as-built RC frame in various seismic zones.
Zone
Ca
CV
Cs, Demand (R = 3.0)
Cs, Capacity
FoS
Remarks
1 2A 2B 3 4
0.08 0.15 0.2 0.3 0.4
0.08 0.15 0.2 0.3 0.4
0.06 0.11 0.14 0.22 0.29
0.20
3.45 1.84 1.38 0.92 0.69
OK OK OK N.G. N.G.
Table 10
Seismic performance of haunch retrofitted RC frame in various seismic zones.
Zone
Ca
CV
Cs, Demand (R = 7.50)
Cs, Capacity
FoS
Remarks
1 2A 2B 3 4
0.08 0.15 0.2 0.3 0.4
0.08 0.15 0.2 0.3 0.4
0.03 0.05 0.06 0.10 0.13
0.20
7.88 4.20 3.15 2.10 1.58
OK OK OK OK OK
effective depth greater than 36 in. (915 mm). However, the observed behavior also suggests the use of side facing/skin reinforcement in beam having depth equal to 15.75 in. (400 mm) to control the spread of vertical cracks at the beam-ends. The following are concluded on the basis of analytical investigation performed on the experimentally tested compliant and non-compliant frames:
The bi-linearized lateral force-displacement response of tested frame revealed that the structural lateral stiffness reduced up to 60% due to non-compliance of frame construction. The corresponding lateral load resistance reduced up to 50% and the corresponding translation displacement ductility ratio reduced up to 30%. The overstrength factor of 4.13 was calculated in case of code-compliant frame that reduced to 2.15 in case of fully non-complaint frame, exhibiting a reduction of about
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14 50%. This indicates the code-compliant frame possesses significant overstrength in comparison to the code-conforming overstrength of 3.0. The ductility factor of 1.83 was calculated in case of codecompliant frame that reduced to 1.35 in case of fully noncomplaint frame, exhibiting a reduction of about 26%. This indicates both the compliant and non-complaint frames possesses ductility lower than the ductility of codeconforming frame of 2.67. This is because of the joint mechanism of tested frame, and less redundant structural system, that hampers the ductile response of frame. The response modification factor in case of code-complaint frame is 7.54 that reduced to 2.91 in case of fully noncompliant frame, exhibiting a reduction of 60% in response modification factor of frame due to non-compliance of constructions. Realizing the fact that one-bay frame is less redundant where reduction factor of 0.80 is proposed in some international codes e.g. Mexico (Tena-Colunga and CortesBenitez, 2015) [35], the response modification factor calculated here (i.e. 7.54) can increase to 9.42 (7.45/0.8 = 9.42) in case of frame having four lines of vertical columns for carrying lateral load. In such case, the computed response modification factor is larger than the value of 8.50/8.0 given in the old/new seismic code, respectively. This indicates the beneficial role of seismic detailing given in the ACI-318 seismic provisions.
M. Rizwan et al. Appendix A. Input Protocol and test sequence for nomcompliant models
Model-2 Time History
Run (%)
Observed Input PGA (g)
1994 Northridge Earthquake
Self-Check Run 5% Run 10% Run 30% Run 40% Run 50% Run 60% Run 70% Run 80% Run 90% Run
0.30 0.048 0.29 0.22 0.37 0.38 0.48 0.55 0.88 0.59
Run (%)
Observed Input PGA (g)
Self-Check Run 5% Run 10% Run 20% Run 30% Run 40% Run
0.015 0.52 0.25 0.31 0.35 0.73
Run (%)
Observed Input PGA (g)
Self-Check Run 5% Run 10% Run 20% Run 30% Run
0.013 0.33 0.88 0.27 0.44
Run (%)
Observed Input PGA (g)
Self-Check Run 5% Run 10% Run 20% Run 30% Run 40% Run 50% Run
0.06 0.10 0.07 0.11 0.20 0.29 0.49
Model-3 Time History
1994 Northridge Earthquake
The following are concluded on the basis of static force procedure used for the seismic performance assessment of codecompliant and non-compliant/deficient reinforced concrete frames in various seismic zones: The code-conforming frame is capable to resist the design base earthquakes in highest seismic zone without exceeding the code permissible story drift limit of 2.50%. The fully non-complaint frame is capable to resist the design base earthquakes only in low to moderate seismic zones but fails to resist design base earthquakes in highest seismic zones (i.e. zone 3 and zone 4) without exceeding the code permissible story drift limit of 2.50%.
Model-4 Time History
1994 Northridge Earthquake
The present research suggests taking in to account the reliability/redundancy factor in proposing response modification factor for seismic design of structures. This factor will take care of the deficiencies encountered in the regional construction practices, and will vary from 1.0 (in case of compliance) to 0.4 (in case of fully non-compliance). Model-5
Declaration of Competing Interest
Time History
The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
1994 Northridge Earthquake
Acknowledgments The authors are grateful to the anonymous reviewers for their constructive remarks that improved the quality of manuscript.
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Shake table investigations on code non-compliant reinforced concrete frames
15
Appendix B. See Tables B1–B4.
Table B1
Observed damages in Model-2. Run
70%– 0.55 g
80% – 0.88 g
90% – 0.59 g
Top sotrey maximum dispalcement Inch (mm)
9.30 (236.26)
14.18 (360.11)
19.56 (496.95)
Top sotrey maximum Drift (%)
3.33
5.08
7.01
MaximumBase shear Forrce kips (kN)
42.39 (188.54)
40.70 (181.02)
33.68 (149.83)
Slight Diagonal Cracks in Joints on Ground Storey
Severe Diagonal Cracks in Joints on Ground Storey
Diagonal Cracks in Joint Panel, Ground Storey
Significant Flexure Cracks in First Storey Columns
Concrete Wedge Detachment from First Storey Joint
Damages Observed in Columns and Joint Panels during 70% Run
Damages Observed in Columns and Joint Panels during 80% Run
Observe Damage
Bat-Like Wedge Detachment from First Storey Joint Damages Observed in Joint Panels of Ground and First Storey during 90% Run
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16
M. Rizwan et al.
Table B2
Observed damages in Model-3. Run
5%– 0.52g
30% – 0.35 g
40% – 0.73 g
Top sotrey maximum dispalcement Inch (mm)
4.87 (123.69)
7.18 (182.365)
13.31 (338.19)
Top sotrey maximum Drift (%)
1.75
2.57
4.77
MaximumBase shear Forrce kips (kN)
26.51 (117.92)
30.91 (137.48)
41.48 (184.50)
Flexure Cracks in Beams and Columns, Ground Storey
Severe Bat-Like Cracks in Joints on First Storey
Slight Cracks in Joint Panel on Ground Storey
Damage Progress in Joint Panel on Ground Storey
Damages Observed in Columns, Beam
Damages Observed in Joint Panels of Ground and First Storey during 30% Run
Cover Detachment and Damage in Joint on First Storey
Observe Damage
and Joint Panels during 5% Run
Severe Damage to Joint Panel on Ground Storey Damages Observed in Joint Panels of Ground and First Storey during 40% Run
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Shake table investigations on code non-compliant reinforced concrete frames Table B3
17
Observed damages in Model-4. Run
5%– 0.33g
10% – 0.88 g
30% – 0.44 g
Top sotrey maximum dispalcement Inch (mm)
3.15 (79.97)
5.91 (150.23)
9.63 (244.57)
Top sotrey maximum Drift (%)
1.13
2.12
2.12
MaximumBase shear Forrce kips (kN)
22.87 (101.72)
30.53 (135.82)
30.53 (135.82)
Flexure Cracks in Beam and Column on Ground Storey
Severe Bat-Like Cracks in Joints on First Storey
Cracks in Joint Panel on Ground Storey
Damage Progress in Joint Panel on Ground Storey
Cover Detachment and Spalling in Joint on First Storey
Observe Damage
Damages Observed in Columns, Beam and Joint Panels during 5% Run
Damages Observed in Joint Panels of Ground and First Storey during 10% Run
Significant Damage to Joint Panel on Ground Storey Damages Observed in Joint Panels of Ground and First Storey during 30% Run
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18
M. Rizwan et al.
Table B4
Observed damages in Model-5. Run
5%– 0.10g
30% – 0.20 g
50% – 0.49 g
Top sotrey maximum dispalcement Inch (mm)
1.01 (25.65)
3.66 (92.95)
10.93 (277.60)
Top sotrey maximum Drift (%)
0.36
1.31
3.92
MaximumBase shear Forrce kips (kN)
8.08 (35.96)
18.73 (83.31)
28.07 (124.84)
Flexure Cracks in Beam and Column on Ground Storey
Concrete Crushing at Column Top Ends, Ground Storey
Extend of Damage to Joints on Ground Storey
Cracks in Top Ends of Columns on First Storey
Cover Concrete Detachment from Joints on First Storey
Extend of Damage to Joints on First Storey
Damages Observed in Beams and Columns during 5% Run
Damages Observed in Column Ends on Ground Storey and Joint Panels of First Storey during 30% Run
Damages Observed in Joint Panels of Ground and First Storey during 50% Run
Observe Damage
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