MASONRY RESEARCH NEWSLETTER
THE MASONRY SOCIETY
OBJECTIVE
TO PRESENT SUMMARIES
OF RECENT MASONRY
RESEARCH TO HELP
ADVANCE THE
KNOWLEDGE OF
MASONRY
EDITORS
Jennifer E. Tanner, PhD
Saadet Toker Beeson, PhD
CONTACT INFORMATION
tannerj@uwyo.edu
Saadet.Beeson@utsa.edu
FINANCIAL SUPPORT
The Masonry Society
Endowment Fund
INSIDE THIS ISSUE
Damage Detection in
1
Metamorphic Stone
Blocks Utilizing ImpactEcho Testing and
Modal Analysis
Seismic Assessment of
Out-of-Plane Loaded
Unreinforced Masonry
Walls
2
Single-Wythe Concrete 3
Masonry Walls—
Investigation of Early
Age Shrinkage Mortars
Structural Assessment of 3
Fort Sumter National
Monument
Experimental
5
Evaluation of Concrete
Masonry Prism
Compressive Strength
A Comparison of Test
Methods for Lap
Splice Performance in
Masonry Walls
6
Performance of
Concrete Masonry
Shear Walls with
Integral Confined
Concrete Boundary
Elements
7
2012
DAMAGE DETECTION IN METAMORPHIC STONE BLOCKS UTILIZING
IMPACT-ECHO TESTING AND MODAL ANALYSIS
By Alexander Charles Jording (EIT) and Ece Erdogmus, Assoc. Prof. - University of Nebraska at Lincoln
Existing buildings decay with the passage of time, and as a result,
continually require maintenance and rehabilitation. When considering existing
structures, testing is often necessary to complete structural analysis in preparation
for rehabilitation or repair. Non-destructive testing (NDT) provides a structural
assessment method applicable to a variety of materials and structures. Impactecho testing, one of the many NDT techniques can be adopted to develop a
reliable and standardized method to assess the condition of rectangular
metamorphic stones. An international research project at Antiochia ad Cragum,
near present-day Gazipaşa, Turkey was used as a case study for structural
assessment conditions. Blocks from a 3rd century Imperial Roman temple (Figure 1)
at Antiochia ad Cragum, served as the primary reference for this study.
Impact-echo testing uses impact generated stress waves and their
propagation to examine the characteristics of various materials and their
associated interaction. This NDE test method introduces of a mechanical impact,
typically with an instrumented hammer (Figure 2), to the specimen and creates an
internal stress wave. This corresponding wave is reflected by internal voids or
external surfaces. The reflected stress waves cause displacements on the surface
of the specimen. These displacements are measured and recorded with respect to
time and eventually get transformed into the frequency domain by a data
acquisition system, often a signal analyzer. From this information, the structural
integrity of the specimen can be obtained by observing predominant frequencies
displayed by the specimen in its associated amplitude spectrum (Figure 3).
Figure 1. 3rd
Century Imperial
Roman Temple
Block
Figure 2.
Instrumented
Hammer and
Accelerometer
The building stones at Antiochia ad Cragum have
large dimensional variations, degrees of damage, and
surface conditions. While many small variances were
observed between stones, the majority of the wall blocks of
the temple are of similar cross-sectional dimensions (aspect
ratios of 1.0) and of similar material characteristics. The
wall blocks having similar aspect ratios and being vital to
the restoration effort were the focus of this study.
At the University of Nebraska, a block was cast
from mortar in order to closely match the properties of
marble. Impact-echo tests were conducted on the mortar
block and compared with a theoretical eigenvalue analysis,
Figure 3. Frequency Amplitude Spectrum
previously completed by Sansalone and Streett and
explained in Impact-Echo: Nondestructive Evaluation of
Concrete and Masonry. The results from the impact-echo
testing in the lab were also correlated to a finite element model (FEM). Errors of 6% and 2.4% were
observed when comparing the mortar block results with the theoretical eigenvalue analysis and the
FEM, respectively. This correlation allowed for the development of a Real-time Impact-echo Analysis
Program (RIAP). RIAP utilizes user input variables, pertaining to the testing specimen, and outputs the
expected frequencies via eigenvalue analysis. Comparisons of the frequency peaks in RIAP with the
frequency peaks from impact-echo testing provide the necessary information for structural evaluation.
The comparisons between impact-echo testing, theoretical eigenvalue analysis, and finite element
modeling and the development of RIAP provide engineering and archaeological research teams with
reliable methods to structurally evaluate rectangular stones. Currently, RIAP, in conjunction with impactecho testing are being used in Turkey to analyze the marble wall blocks to better understand the
structural integrity of the stones and therefore the degree of restoration that may be possible. These
techniques and methods are not exclusive to marble, and with minimal alteration, can be applied to a
variety of materials including other types of masonry, concrete, and even wood.
Sansalone, M. J., & Streett, W. B. (1997). Impact-Echo: Nondestructive Evaluation of Concrete
and Masonry. Ithaca, NY: Bullbrier Press.
UPCOMING:
TMS Spring Meetings, Vancouver, CA, May 31-June 2, 2013
12th Canadian Masonry Symposium, Vancouver, CA, June 2-5, 2013
www.cms2013.ca
Page 2
2012
Masonry Research Newsletter
SEISMIC ASSESSMENT OF OUT-OF-PLANE LOADED UNREINFORCED MASONRY WALLS
By Hossein Derakhshan (PhD), Jason M. Ingham, Assoc. Prof. - University of Auckland and
Michael C. Griffith - Prof. University of Adelaide
This research involved an investigation into the behaviour of unreinforced masonry (URM) walls subjected to out-ofplane seismic forces; the study focused primarily on one-way spanning walls connected to flexible diaphragms.
Expressions
that
describe
the
intermediate-height cracking of out-of-plane
loaded one-way spanning URM walls were
formulated, and the sensitivity of crack height to
different parameters was investigated. Crack
height was identified to have significant effects
on both wall instability and lateral resistance,
thus
governing
the
force-displacement
relationship. Crack height was found to be
particularly sensitive to low values of masonry
b) Dependency of PMR ratio on the bond strength and/or to low values of applied
a) Dependency of wall instability
overburden ratio
displacement on the overburden ratio
overburden. An analytical wall behavioural
model characterised by the maximum wall
Figure 1: Dependency of actual wall behavioural data on several parameters resistance and wall instability displacement was
next developed by including the effects of finite
masonry compressive strength. Fig. 1 shows that
depending on the ratio of the applied
overburden to the wall weight (), the ratio of
the wall instability displacement (Fig. 1a) and
maximum resistance (Fig. 1b) to the equivalent
values when a rigid-body assumption is made is
significantly less than unity. The latter has been
defined as percentage of maximum resistance
(PMR) ratio.
A comprehensive quasi-static testing
a) Static airbag testing of URM walls
b) Trilinear wall behavioural model programme was developed by subjecting fullscale two-leaf and three-leaf URM walls to outFigure 2: Wall testing and behavioural model idealisation
of-plane uniform forces. The measured wall
maximum lateral resistance was compared to
the equivalent value obtained assuming infinite
masonry compressive strength and to the values
obtained from the aforementioned analytical
model. The analytical model was then
calibrated, and an idealised predictive trilinear
model was developed.
An in-field testing programme was
completed, and wall support types and
Figure 3: One-way walls tested in-situ; shades outline airbag loaded area characteristics and their effect on wall
behaviour were investigated by testing multiple
walls. Wall support types included timber
diaphragms, concrete ring beam, and grouted steel anchorage bars. The effect of
plaster and wall finishing was investigated, and conditions for ensuring that the
idealised predictive model generated acceptable results were determined.
A simplified reduced-degree-of-freedom numerical model was
developed to study the dynamic behaviour of URM walls, with critical requirements
of the model being an ability to account for the effects of flexible supports and an
ability to predict the behaviour of two-storey walls. An existing 2DOF model was
expanded to a 4DOF model to simulate the behaviour of walls of up to twostorey. The model was subjected to ground motion histories and the time
integration results were verified by using the results of shake table tests conducted
by other researchers. Computer analyses showed that diaphragm flexibility
significantly amplifies the out-of-plane loaded URM wall peak response when
compared with the wall peak response obtained assuming rigid diaphragms.
An existing New Zealand procedure for seismic assessment of out-ofplane loaded URM walls was reviewed in detail, and an alternative methodology
for wall assessment was introduced. The new methodology was based on the
Figure 4: Analytical model used to obtain average results of numerical time integration analysis using a suite of seven timehistory records defined for specific New Zealand regions and soil types.
equation of dynamic motion
Masonry Research Newsletter
Page 3
2012
SINGLE-WYTHE CONCRETE MASONRY WALLS
INVESTIGATION OF EARLY AGE SHRINKAGE OF MASONRY MORTARS
By Sarah Ebright (EIT) and Jennifer Tanner—Assoc. Prof. at University of Wyoming
With the rising popularity of single-wythe Concrete Masonry Unit (CMU) walls, concern has developed regarding the
water resistance of this system. The University of Wyoming worked with Atkinson-Noland & Associates (ANA) and the National
Concrete Masonry Association (NCMA) to perform research concerning shrinkage crack development along the bond line along
the mortar and CMU interface (Figure 1a). The research objective was to better understand the conditions that result in bond
line delaminations in CMU construction.
Preliminary tests were performed on the blocks to ensure they adhered to American Society for Testing and Materials
(ASTM) specifications. Five variables were analyzed with respect to mortar shrinkage and crack development: mortar type,
block type, temperature, relative humidity, and wind. Joint quality was also analyzed and measured by the depth of the joint.
Wall specimens were constructed and cured in different environmental conditions (Figure 1b), and shrinkage cracks were
measured and monitored. Each delamination (crack) was referenced with its environmental and material conditions, and this
information was used in the data analysis. Data was analyzed graphically and statistically. Recommendations for CMU
construction were determined from these results of 102 specimens. Cracking occurs at a significantly higher rate in specimens
constructed of Type S mortar, exposed to high temperatures (Figure 2).
Figure 1: a) Shrinkage crack
b) Masonry wall specimen
Figure 2: Total length of cracking by variable
Cracking was predominant in Type N specimens as well as those subjected to high temperatures. Another prominent
trend in the data is the high extent of cracking in the vertical head joints. Overall, vertical cracks constituted about 85% of the
total length of cracking. Furthermore, cracking was more severe on the untooled side of the specimen. It is recommended that
both sides of CMU walls be tooled, and that the masons and inspectors play very close detail to this part of the process.
Tooling both sides of the wall provides redundancy in the weather resistance of the structure.
STRUCTURAL ASSESSMENT OF FORT SUMTER NATIONAL MONUMENT
By Sezer Atamturktur, Assist. Prof.—Clemson University, Saurabh Prabhu (PhD Candidate-Clemson University) and
Rick Dorrance, Chief of Resource Management - Fort Sumter National Monument, National Park Services
Introduction: This multi-faceted study combines several experimental and numerical tasks for the structural assessment of Fort
Sumter, a 19th century North American island fortification located in Charleston, SC (Figure 1), best known as the site where the
first shots of the American Civil War were fired in 1861.
Military forts are designed with different functional purposes than typical unreinforced masonry structures: from housing heavy
ordnance, to protecting the occupants from continuous bombardment. To fulfill these specific design objectives, structural
designs of fortifications are configured differently involving for instance intentional structural discontinues between members.
As such, the existing literature on the assessment of historic masonry structures offers little guidance in developing accurate
numerical models for structural analysis.
Data Collection for Model Development: An extensive site evaluation campaign is completed for purposes of collecting data
required for the faithful representation of materials and geometry in the numerical models. In accordance with the data
collected during site evaluations, three dimensional, non-linear finite element (FE) models are developed. The material
behavior and composition is determined from laboratory tests on material samples obtained on-site. Block specimens are used
to perform tension and compression tests to obtain characteristics of the fort’s masonry. The characteristics of the tabby
concrete infill in the piers and walls are determined from tests on cored specimens.
The three-dimensional (3D), complex geometry of the fort is reproduced in its present form via high precision 3D laser
scanning (Figure 2). Owing to the speed and accuracy of recording the geometries of civil engineering systems without
contacting the structure, 3D scanning has been successfully used in hyper-accurate surveys of several historic masonry structures.
Through this approach, the true geometric features of the fort are preserved in the FE model by considering the accumulated
structural degradations (i.e. permanent deformations, material disintegration, crack formations and support movements) in
addition to initial construction imperfections. The laser scan produces a point cloud consisting of several million points which is
reduced to key points that form a wireframe model. The wireframe is then extruded to form the solid geometry of the
casemate preserving the important structural details.
Page 4
2012
Masonry Research Newsletter
Development of the Structural Finite Element Model: The 3D solid geometry of one of the fort casemates is developed in ANSYS
according to the laser scanning data. The unreinforced masonry and tabby concrete that constitute the fort’s masonry walls are
represented using SOLID65 element with capabilities of simulating brittle cracking and crushing of material in a smeared manner
employing the Willam-Warnke failure criterion. The foundations of the casemate are modeled using a series of vertical and
horizontal linear springs distributed throughout the base of the casemate and defined by a stiffness coefficient, thus, forming a
Winkler spring foundation model (Figure 3).
Similar to the foundations, the lateral restraints to the casemate due to the adjoining casemates are of an uncertain quantity.
The adjacent casemates are thus included in the model as reduced order substructures (superelements), which are represented only by
the DOFs on the interface with the casemate of interest (Figure 3). Thus, the number of DOFs to be analyzed is reduced nearly
threefold by solving only for the structure of interest and the interface DOFs. For modal analysis, component mode synthesis is
applied while for static analysis, a linear mass and stiffness matrix partitioning is applied.
Calibration of FE model: The exterior scarp wall and the interior gun casemates of the fort are constructed separately but adjacent
to each other to form a dry joint at their interface. This discontinuity is modeled using contact and target elements defined by a
friction coefficient. To estimate the friction coefficient, relative accelerations of the two adjacent surfaces due to a hammer impact
applied on the scarp wall is measured. The ratio of the maximum deformation on either side of the interface is used to calibrate the
friction coefficient in the FE model.
Despite performing on-site tests, uncertainties remain in (i) the material property values and (ii) the support conditions at the
base of the structure. To reduce the uncertainty in these input parameters, the model is calibrated through systematic comparisons of
the numerical model output against nondestructive dynamic measurements of first two natural frequencies and corresponding mode
shapes collected on site. Specifically, the elastic modulus of the piers and vaults, and the spring constant of the foundation springs are
calibrated to obtain an improved agreement between predictions and measurements.
Simulation of Loading Scenarios: The calibrated FE model can be used to simulate a variety of static and dynamic loading scenarios
including gravity, differential support settlements, wind load and earthquake. These simulations are obtained taking the remaining
uncertainties in input parameters after model calibration into account. Figure 4 presents the predictions of maximum stress for one of
the investigated support settlement scenarios.
Figure 1: Aerial view of Fort Sumter
(photo credit: www.nps.gov )
Figure 3: FE model with substructures and foundations
modeled using spring elements
Figure 2: (a) Polygonal mesh of a casemate
generated in Polyworks V11; (b) Wireframe of
casemate (Model 3) generated in Rhino v5.0.
Figure 4: Parametric analysis of maximum stress due to
support settlement
Masonry Research Newsletter
2012
Page 5
EXPERIMENTAL EVALUATION OF CONCRETE MASONRY PRISM COMPRESSIVE STRENGTH
By Tyler Witthuhn (EIT) and Jason Thompson, Vice-President for Engineering — National Concrete Masonry Association
Table 1 – Proposed Table used for Unit Strength Method for Concrete Masonry Testing laboratories, including our own
Research and Development Laboratory here at
the National Concrete Masonry Association
(NCMA), have consistently seen higher
compressive strengths then what this table has
allowed by the current design standards (Table
1). As a result, there has been considerable
debate concerning Table 2 of TMS 602-11/ACI
530.1-11/ASCE 6-11, Specifications for
Masonry Structures.
With this in mind, NCMA put together a
testing program of various strength prisms and
mortar types to determine if there was a
technical basis to update this table. Prisms
constructed for testing contained both M/S and
N type mortars mixed either by property or
proportion and tested to ASTM C270, and with
unit strengths ranging from 1900 psi to 5500
psi. For each of these variables three prisms
were tested, yielding 102 total prism tests.
Along with these new prism tests, over
40 data points were mined from previous NCMA
Laboratory research that helped broaden and
confirm the new testing. These points can be seen
in Figure 2 along with a comparison to the new
testing data.
Data points were grouped into ranges
of 500 psi and then each range was statistically
analyzed in order to find the 95% confidence
interval of that set. This number was the basis for
the new table and allowed for both a
conservative estimate while maintaining the trend
of the data. Through this method Table 1 was
formed.
Although all strengths see an increase,
the most pertinent result from the testing occurs
at the lowest unit strengths. ASTM C90 defines
the lowers allowable compressive strength for
units as 1,900 psi, which in the past has aligned
to a 1,500 psi prism strength. From the data
generated the prism strength of a 2,000 psi unit
would now be 2,000 psi, a 1/3rd increase in
overall design prism strength. This direct
correlation between unit strength and prism
strength also applies to a 2,500 psi unit, which
generates a 2,500 psi unit strength.
Data for the Type N mortar prisms was
much more scattered, leading to less consistent
averages and 95% confidence intervals.
Because of this, NCMA elected to cap the unit
strength table at 2,600 psi for Type N mortar.
Although higher strengths can be achieved with
Type N mortar, prism testing is required for
design.
The complete research report can be
found on NCMA’s website.
Net area compressive strength of concrete masonry units, lb/in.2 (MPa) 2,000 (13.8) 2,500 (17.2)
3,000 (20.7) 3,500 (24.1) 4,000 (27.6) 4,500 (31.0) 5,000 (34.5) 5,500 (37.9) Net area compressive strength of masonry, lb/in.2 (MPa) Type M or S Mortar Type N Mortar
2,000 (13.8) 2,500 (17.2)
2,700 (18.6) 2,800 (19.3) 3,100 (21.4) 3,200 (22.1) 3,600 (24.8) 4,000 (27.6) 2,000 (13.8) 2,500 (17.2)
2,600 (17.9) ‐ ‐ ‐ ‐ ‐ Figure 1: NCMA data for Type S mortar including a 95%
confidence interval
Figure 2: Summary of all data
Page 6
2012
Masonry Research Newsletter
A COMPARISON OF TEST METHODS FOR LAP SPLICE PERFORMANCE IN MASONRY WALLS
By Kawsar Ahmed (EIT) and Lisa Feldman, Assoc. Prof.—University of Saskatchewan
Current Canadian provisions for the development and splicing of deformed steel bars in masonry provided in CSA
Standard S304.1-04 are based on provisions developed for reinforced concrete construction. These provisions are likely not entirely
representative because there are several parameters that affect the bond of reinforcement in masonry that are not present in
reinforced concrete construction. The limited experimental data for the bond of deformed steel bars in masonry construction is based
primarily on direct tension tests, which do not accurately capture the behaviour of reinforcement in flexural members. Knowledge of
bond stress variability is also required for the development of reliability-based design provisions for current masonry codes.
An experimental program was conducted to evaluate the performance of lapped bars in pullout specimens and full-scale
wall splice specimens. Contact and non-contact lap splices, where the lapped bars were located in adjacent cells, with No. 15 (0.60
in) bars with 300 mm (12 in.) lap splice lengths, were tested. Eight replicates for each splice arrangement and specimen type, for a
total of 32 specimens, were tested to allow for any statistical differences between mean splice capacities to be identified. Figure 1
shows a double pullout test that was adopted to test splice capacity under direct applied tension. Figure 2 shows a full-scale wall
splice specimen that was tested under four-point loading for a more realistic evaluation of splice capacity.
The contact lap spliced bars in the double pullout specimens developed, as a minimum, the theoretical yield load of the
reinforcement (Figure 3a). In contrast, the mean splice resistance developed by the non-contact splices in the double pullout
specimens was only 46% of the theoretically predicted yield load of the reinforcement (Figure 3b). Similar results were obtained for
wall splice specimens (Figure 4), with non-contact splices developing 78% of the theoretical yield load (Figure 4b). Further statistical
analysis using the student “t” test showed that the capacity of non-contact splices is significantly different from that of contact lap
splices.
Figure 1: Experimental test setup – double pullout specimens
(a)
(b)
Figure 3: Splice resistance versus displacement of the two splices
in representative double pullout specimens with: (a) contact lap
splices, and (b) non-contact lap splices
Figure 2: : Experimental test setup – wall splice specimens
(a)
(b)
Figure 4: Load versus midspan deflection for representative
wall splice specimens with: (a) contact lap splices, and (b) noncontact lap splices
Even though double pullout specimens and wall splice specimens with contact lap splices produced reasonably similar splice
capacity (pullout tests were about 10% conservative), significantly higher strain levels in the wall splice specimens were recorded,
indicating that the difference would be more decisive for a shorter lap length that would result in a bond failure. Non-contact
lapped bars in the double pullout specimens developed 41% lower splice capacity than that developed in the wall splice specimens.
The difference in failure modes was even more significant between the two specimen types with this splice arrangement. Specimen
splitting was observed in double pullout specimens due to development of the resulting in-plane moment in contrast to typical bond
failure by reinforcement pullout in the wall splice specimens. The experimental program therefore suggests that wall splice specimens
are more realistic and appropriate for the evaluation of splice capacity in reinforced masonry.
Masonry Research Newsletter
Page 7
2012
PERFORMANCE OF CONCRETE MASONRY SHEAR WALLS WITH
INTEGRAL CONFINED CONCRETE BOUNDARY ELEMENTS By Will Cyrier (EIT and Graduate Assistant) and David McLean – Professor, Washington State University
Benson Shing - Professor, University of California at San Diego
and Richard Klingner – Professor, University of Texas at Austin
This research project investigated the behavior of masonry walls incorporating
integral confined concrete boundary elements at each end under lateral loading. This
project was funded by the National Institute of Standards and Technology (NIST) as part of
a joint study between researchers at the University of California at San Diego, the
University of Texas at Austin and Washington State University to develop improved
performance-based design provisions and methodologies for reinforced concrete masonry
shear walls. Detailed discussion and results of this project are provided by Cyrier (2012).
The 2011 Building Code Requirements and Specifications for Masonry Structures (MSJC,
2011) provides design guidelines for boundary elements in masonry walls. However, these
guidelines cover primarily geometric issues. No guidance is given in the MSJC Code for
effective confinement techniques for application to masonry. As a result, the MSJC Code
requires that testing be performed to verify that the provided detailing is capable of
developing a strain capacity in the boundary elements that is in excess of the imposed
strains. In contrast, the ACI 318-11 Building Code Requirements for Structural Concrete and
Commentary (ACI, 2011) provides prescriptive detailing requirements for specially confined
boundary elements in structural concrete walls. The MSJC Commentary states “it is hoped
that reasonably extensive tests will be conducted in the near future, leading to the
development of prescriptive detailing requirements for specially confined boundary
elements of intermediate as well as special reinforced masonry shear walls” (MSJC 2011).
In this project, four, fully grouted, concrete masonry shear walls with integral confined
concrete boundary elements were designed according to the provisions of the 2011 MSJC
and the 2011 ACI-318 codes. The walls were subjected to a prescribed cyclic, in-plane
lateral displacement sequence. Performance measures evaluated include peak load
capacities; drifts at various limit states; decoupled drift components from shear, flexure and
sliding; displacement and curvature ductilities; plastic hinge lengths; total energy
dissipation; and equivalent viscous damping values. The effects of incorporating the
confined concrete boundary elements, axial compressive stress, boundary element
geometry, and size of transverse hoops in the boundary elements were evaluated to
determine their influence on wall performance. Test results in this research were compared
to results from tests on two similar masonry walls without boundary elements performed in
an earlier study by Kapoi (2012).
The configuration of a typical shear wall specimen with the integral reinforced
concrete boundary elements is given in Figure 1. Two of the walls with boundary elements
(BE-1 and BE-2) had rectangular cross sections, while the other two walls with boundary
elements (BE-3 and BE-4) had flanged sections, as shown in Fig. 2. The comparable
masonry walls (C7 and C8) had the same section and reinforcement as BE-1 and BE-2 but
without boundary elements were subjected to the same loading conditions.
Figure 3 compares the load-vs.-displacement hysteretic curves for walls with
boundary elements (BE-1 and BE-2) to the similar masonry walls (C7 and C8). It is evident
that the boundary elements led to a significantly more ductile behavior.
Very significant improvements in performance were achieved in the masonry walls
with integral confined concrete boundary elements when compared to similar masonry walls
without confinement. Masonry walls with integral confined concrete boundary elements
Figure 1: Typical wall specimen with
boundary elements
Figure 2: Wall Sections: Rectangular
Boundary Elements for Specimens BE-1
and BE-2 (top) and Boundary Elements
with Return for Specimens BE-3 and BE-4
(bottom)
increased displacement ductility
values by 50% and total energy
dissipation was 2-3 times greater
compared to results from similar
masonry walls without boundary
elements. Boundary elements with
returns provided added stability at
large displacements, allowing for
significant increases in energy
dissipation. Walls with rectangular
boundary elements failed when the
boundary element core buckled out-of
-plane. Walls with flanged boundary
Figure 3: Comparison of Load-Displacement Hysteretic Curves for Walls BE-1 and C7 (left)
elements failed due to low-cycle
and Walls BE-2 and C8 (right)
fatigue fracture of the longitudinal
reinforcing bars.
References: Cyrier, W. (2012). “Performance of Concrete Masonry Shear Walls with Integral Confined Concrete Boundary
Elements,” M.S. Thesis, Department of Civil and Environmental Engineering, Washington State University, Pullman, WA
Kapoi, C. (2012). “Experimental Performance of Concrete Masonry Shear Walls Under In-Plane Loading,” M.S. Thesis,
Department of Civil and Environmental Engineering, Washington State University, Pullman, WA.