strength and stability of slender post-tensioned concrete masonry walls under ...€¦ · ·...

University of Minnesota Post-tensioned Masonry Research Program Department of Civil Engineering

STRENGTH AND STABILITY OF SLENDER, POST-TENSIONED CONCRETE MASONRY WALLS

UNDER TRANSVERSE LOADING

by

J. R. Bean, C. R. Drake, and A. E. Schultz

Department of Civil Engineering University of Minnesota Minneapolis, MN 55455

A Final Report to

The National Concrete Masonry Association Foundation 13750 Sunrise Valley Drive Herndon, VA 20171-4662

July 2004

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TABLE OF CONTENTS

TABLE OF FIGURES.................................................................................................................. 4

TABLE OF TABLES.................................................................................................................... 5

1. INTRODUCTION ................................................................................................................... 6

1. INTRODUCTION ................................................................................................................... 6 1.1 ELEMENTS OF PRESTRESSED MASONRY................................................................................. 6 1.2 ADVANTAGES OF PRESTRESSED MASONRY SYSTEMS ............................................................ 8 1.3 APPLICATIONS OF PRESTRESSED MASONRY........................................................................... 9

2. BACKGROUND INFORMATION ........................................................................................ 9 2.1 GENERAL BEHAVIOR ........................................................................................................... 10 2.2 OBJECTIVES ......................................................................................................................... 11

3. EXPERIMENTAL PROGRAM............................................................................................ 12 3.1 VARIABLES .......................................................................................................................... 12 3.2 PRIMARY ELEMENTS OF SPECIMENS .................................................................................... 13 3.3 CONSTRUCTION OF SPECIMENS ............................................................................................ 16 3.4 LOAD FRAME ....................................................................................................................... 19

3.4.1 Axial Loading System................................................................................................... 20 3.4.2 Lateral Loading System ............................................................................................... 22

3.5 INSTRUMENTATION .............................................................................................................. 23 3.5.1 Load Measurement....................................................................................................... 24 3.5.2 Displacement Measurement......................................................................................... 24 3.5.3 Rotation Measurement ................................................................................................. 25 3.5.4 Data Collection............................................................................................................ 26

3.6 LOADING PROCEDURE ......................................................................................................... 26 3.7 MODIFICATIONS TO TEST SETUP AND INSTRUMENTATION ................................................... 26

4. MATERIAL PROPERTIES................................................................................................. 28 4.1 UNIT PROPERTIES ................................................................................................................ 28 4.2 MORTAR PROPERTIES .......................................................................................................... 29 4.3 PRISM PROPERTIES............................................................................................................... 29

4.3.1 Masonry Compressive Strength, f’m............................................................................. 30 4.3.2 Modulus of Elasticity, Em............................................................................................. 31 4.3.3 Flexural Tensile Strength Normal to Bed Joints, f’tn .................................................. 32 4.3.4 Measured Material Properties..................................................................................... 33

5. SPECIMEN BEHAVIOR...................................................................................................... 33 5.1 VISUAL OBSERVATIONS....................................................................................................... 34 5.2 MEASURED LOADS .............................................................................................................. 35

5.2.1 Distribution of Lateral Load........................................................................................ 36 5.2.2 Moment-Displacement Behavior ................................................................................. 36 5.2.3 Tendon Stress Histories ............................................................................................... 38

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5.3 MEASURED DISPLACEMENTS ............................................................................................... 39 5.3.1 Displacement Profiles.................................................................................................. 39 5.3.2 Tendon Displacement Relative to the Wall.................................................................. 41 5.3.3 Axial Strain-Displacement Relations........................................................................... 42

5.4 MEASURED ROTATION......................................................................................................... 43 5.5 MEMBER STABILITY ............................................................................................................ 45

6. ASSESSMENT OF WALL BEHAVIOR ............................................................................. 45 6.1 INITIAL CRACKING............................................................................................................... 45 6.2 JOINT OPENING .................................................................................................................... 47 6.3 MOMENT STRENGTH............................................................................................................ 50 6.4 CRUSHING............................................................................................................................ 50

7. CONCLUSIONS AND RECOMMENDATIONS................................................................ 52

Acknowledgements ..................................................................................................................... 53

References.................................................................................................................................... 53

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TABLE OF FIGURES FIGURE 3.1 STEEL PLATE............................................................................................................... 14 FIGURE 3.2 CONCRETE BOND BEAM.............................................................................................. 14 FIGURE 3.3 EXTERNAL TENDON RESTRAINTS................................................................................ 15 FIGURE 3.4 SPECIMEN DETAILS ..................................................................................................... 16 FIGURE 3.5 PHOTOGRAPH OF LAYING FIRST COURSE OF A CLAY BRICK WALL ............................ 17 FIGURE 3.6 PHOTOGRAPH OF CONCRETE BLOCK WALL CONSTRUCTION....................................... 17 FIGURE 3.7 PHOTOGRAPH OF CONCRETE AND CLAY WALL SPECIMENS ........................................ 18 FIGURE 3.8 SCHEMATIC OF LOAD FRAME ...................................................................................... 19 FIGURE 3.9 VERTICALLY ORIENTED, SIMPLY SUPPORTED SPECIMENS .......................................... 20 FIGURE 3.10 SCHEMATIC OF LOAD FRAME PLAN VIEW................................................................. 21 FIGURE 3.11 SCHEMATIC OF WHIFFLETREE SYSTEM ..................................................................... 23 FIGURE 3.12 SCHEMATIC OF LVDT AND TILTMETER DIAGRAM ................................................... 25 FIGURE 4.1 DIMENSIONAL PROPERTIES OF THE CONCRETE BLOCK UNITS..................................... 28 FIGURE 4.2 DIMENSIONAL PROPERTIES OF THE CLAY BRICK UNITS.............................................. 29 FIGURE 5.1 PHOTOGRAPH OF WALL PC6-150-R DURING TESTING ............................................... 35 FIGURE 5.2 PROFILE OF PC6-150-R DURING TESTING .................................................................. 35 FIGURE 5.3 MOMENT DISTRIBUTIONS FOR WALL PC2-75-U......................................................... 36 FIGURE 5.4 EXPERIMENTAL MOMENT - DISPLACEMENT BEHAVIOR OF CMU WALLS .................. 37 FIGURE 5.5 EXPERIMENTAL MOMENT - DISPLACEMENT BEHAVIOR OF CLAY BRICK WALLS........ 37 FIGURE 5.6 EXPERIMENTAL TENDON STRESS HISTORIES FOR THE CMU WALLS .......................... 38 FIGURE 5.7 EXPERIMENTAL TENDON STRESS HISTORIES FOR THE CLAY BRICK WALLS ............... 39 FIGURE 5.8 DISPLACEMENT PROFILES OF THE CONCRETE BLOCK WALLS AT NOMINAL CAPACITY

............................................................................................................................................... 40 FIGURE 5.9 DISPLACEMENT PROFILES OF THE CLAY BRICK WALLS AT NOMINAL CAPACITY ....... 40 FIGURE 5.10 TENDON DISPLACEMENT RELATIVE TO WALL FOR THE CONCRETE BLOCK SPECIMENS

............................................................................................................................................... 41 FIGURE 5.11 TENDON DISPLACEMENT RELATIVE TO WALL FOR THE CLAY BRICK SPECIMENS..... 42 FIGURE 5.12 TENSION AND COMPRESSION SURFACE STRAIN FOR THE CONCRETE BLOCK WALLS 43 FIGURE 5.13 TENSION AND COMPRESSION SURFACE STRAIN FOR THE CLAY BRICK WALLS ......... 43 FIGURE 5.14 ROTATION-DISPLACEMENT RELATIONSHIP FOR THE CONCRETE BLOCK WALLS ...... 44 FIGURE 5.15 ROTATION-DISPLACEMENT RELATIONSHIP FOR THE CLAY BRICK WALLS................ 44 FIGURE 6.1 INITIAL CRACKING DETERMINED BY VERTICAL STRAINS ........................................... 46 FIGURE 6.2 DISPLACEMENT AND STRAIN DISTRIBUTIONS ............................................................. 48 FIGURE 6.3 JOINT OPENING DETERMINED BY VERTICAL STRAINS................................................. 49 FIGURE 6.4 CRUSHING STRAIN OF THE CONCRETE BLOCK WALLS ................................................ 51 FIGURE 6.5 CRUSHING STRAIN OF THE CLAY BRICK WALLS ......................................................... 51

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TABLE OF TABLES TABLE 3.1 SPECIMEN TEST MATRIX .............................................................................................. 13 TABLE 4.1 MASONRY COMPRESSIVE STRENGTH, F’M .................................................................... 31 TABLE 4.2 MODULUS OF ELASTICITY, EM ...................................................................................... 32 TABLE 4.3 MODULUS OF RUPTURE, F’TN ........................................................................................ 33 TABLE 4.4 MEASURED MATERIAL PROPERTIES ............................................................................. 33 TABLE 5.1 LOCATION OF JOINT OPENING ...................................................................................... 34 TABLE 6.1 INITIAL CRACKING ....................................................................................................... 47 TABLE 6.2 JOINT OPENING............................................................................................................. 49 TABLE 6.3 MOMENT STRENGTH .................................................................................................... 50

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1. INTRODUCTION

Since the advent of high-rise frame buildings, masonry has had overwhelming competition in the construction industry, primarily from steel and concrete. In addition, timber has proven to be an alternative competitor ([2], [27]). The cause of masonry’s reduced use in the recent past has been attributed to the fact that there is less of a knowledge base for structural masonry than there is for its primary competitors. Additionally, many engineers do not have experience designing with this material ([2], [27]). In response to the need to further understand the behavior of masonry, a renaissance began in Europe during the 1940’s, led by Professor Paul Haller in Zurich, Switzerland ([2], [7]). Through research, improved materials, innovative construction ideas and refined design procedures, the first masonry building code was developed in Switzerland in 1943 [2]. Following these developments in Europe, the first masonry code was established in the United States in 1966 [2].

During the mid 1960’s, research, development, and application of a new masonry construction process, prestressed masonry was taking place in England, Australia and New Zealand ([2], [7], [8], [9], [10], [11], [27], [30]). Prestressed masonry was used to design and build walls, even though it took about two decades before any guidance was given in the form of prestressed masonry design code provisions. In 1985, the first provisions for the design of prestressed masonry were released by the British Standards Institute ([4], [8], [9], [27], [29]). Since then, structural engineers and builders in the United Kingdom have implemented prestressed masonry primarily into designing diaphragm and fin walls ([11], [22], [27]). Following these developments in the United Kingdom, the Masonry Standards Joint Committee (MSJC) released its 1999 code provisions in the United States with guidelines for designing prestressed masonry ([4], [8], [13], [29], [31]).

The potential for prestressed masonry to become a viable building system in contemporary construction is great. However, much must be understood about its properties before architects, engineers, and builders can incorporate it safely and economically in building construction. This section will address the following aspects of prestressed masonry:

1) description of the elements of prestressed masonry,

2) analysis of the advantages of the system, and 3) possible applications for prestressed masonry.

1.1 Elements of Prestressed Masonry

The primary elements in a prestressed masonry system are units laid using standard procedures, high strength prestressing steel run through the cavities in the units, and anchorages for the prestressing at the bottom and top of the walls. Either brick or block units, laid on a bed of mortar, are used for this system, but there is also a prestressed masonry system being developed using mortarless masonry units ([6], [17], [19]).

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Most common applications of post-tensioned masonry (PTM) have incorporated high strength threaded rods joined together with couplers along the height of the wall, with other systems using continuous, flexible, high-strength strand installed inside a duct ([8], [13], [17], [20]). Corrosion protection for prestressing steel should not depend on masonry cover alone ([8], [20]). Suggested corrosion protection for tendons includes the use of a bitumen paint coating, heavyweight paper wrap, waterproof plastic tape, or a grease filled duct, while stainless steel and galvanized tendons are acknowledged as alternative methods of protection in the 2002 MSJC specifications commentary ([20], [27,] [29]). There are two common construction methods for placing the tendons, restrained and unrestrained ([3], [13]). These two methods, in turn, affect the manner in which the tendons interact with the surrounding masonry. In the restrained condition, the transverse motion between the tendon and surrounding masonry is prevented over the full height of the wall, and precautions can be taken to allow the tendon to be bonded or unbonded along its length ([2], [28]). The most obvious technique for achieving the restrained condition is grouting, but even though there are structural benefits to using a fully grouted cavity, it is anticipated that the most successful prestressed masonry systems will be ungrouted because of the ease of construction and overall economy ([28], [29]). Besides full grouting along the length of the tendon, a tendon can also be restrained intermittently along its height by maintaining the location of the tendon relative to the masonry by placing the tendon in contact with the masonry, or through the use of grout plugs or mechanical devices ([2], [8], [13], [18], [28], [31]). It should be noted that only acceptable restraint device currently recognized by the MSJC code provisions is the grout plug. At this time, the use of three restraints located at the quarter points of the walls is considered to be sufficient for providing full lateral restraint, but additional restraints may be necessary to distribute the lateral tendon force along the height of the walls when loaded out-of-plane ([13], [20], [28]). In the case of unrestrained tendons, nothing is placed between the top and bottom anchorage location, so the tendon can move freely within the masonry cavity [2].

Anchorage is needed both at the bottom and top of the wall, and it is often preferable to place the anchors in concrete elements [9]. At the bottom of the tendon, anchorage to the foundation or floor slab can be accomplished through the use of cast-in-place anchors, floor slab anchors, and adhesive or epoxy anchors ([4], [13], [29]). At the top, the anchorage is incorporated into a bond beam. The bond beam is typically a precast concrete element, but it can also be a masonry bond beam, which distributes the tendon force to the masonry below ([9], [10], [13], [27]). The bond beam may contain openings at the tendon locations that allow the bearing plate and anchorage device to be recessed so that they do not protrude above the top of the wall [27]. The top anchorage can be as simple as a nut, washer, and plate in the case of a threaded rod tendon, or a self-locking device when a strand tendon is used [27]. As was the case for tendons, care must be taken to prevent corrosion at the anchorages. This protection is done by providing flashing and weep holes at the base of the wall so water does not collect at the bottom anchor, and by filling any recesses for the top anchors in the bond beam with grease, mortar, or grout ([20], [27]). Once the wall has been constructed using the necessary elements and the masonry has achieved its specified strength, the tendons can be stressed ([8], [13], [27]). Since effective prestressed masonry design depends on applying a specified amount of precompression to the wall through tendon stressing, it is vital to achieve the required prestressing force in each tendon. Various

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methods can be used to obtain the desired prestressing force, including tightening the tendon with a torque wrench, hydraulic jack, or impact wrench, and the tendon force can be monitored using load cells, pressure meters, or direct tension indicating (DTI) washers ([13], [15], [27]). In addition to the aforementioned procedures, it is also advisable to verify tendon stress by measuring tendon elongation to ensure that the tendon is being stressed over its entire length ([5], [13]).

1.2 Advantages of Prestressed Masonry Systems In the information presented thus far, several primary advantages of a prestressed masonry system have been mentioned, but it is important to analyze these more thoroughly. In the design of many masonry buildings, the masonry is only considered to be architectural hence the reason for incorporating a steel or concrete frame. Prestressed masonry, like other masonry systems, offers a high degree of architectural desirability. Yet, the structural capacity of prestressed masonry offers the opportunity to lighten, and possibly eliminate, the traditional steel and concrete structures that are often relied upon in many masonry buildings [27]. Prestressed masonry provides several economic advantages in the building process. Economy, material availability, relative simplicity of the process, and ease of construction provide advantages for achieving low construction costs ([9], [27], [10]). In addition, full grouting is not necessary for prestressed masonry; which results in material and labor cost savings when compared to conventionally reinforced masonry ([8], [32]). Selective grouting of bond beams and grout plugs can be achieved with much greater economy and efficiency than full grouting. Also, when compared with conventional masonry construction, the number and spacing of bars is reduced, in turn facilitating the laying of the units [17]. Because of these aspects of prestressed masonry, there is the potential for a shorter construction schedule depending upon the manner in which the prestressing operation is scheduled relative to other construction tasks[17]. Other economic advantages of a prestressed system include lowering building weight if grouting is eliminated or minimized, prestressing serving as temporary wall bracing, and minimal interruption to a building function when used for repairing existing structures ([8], [29], [31]). Another positive feature of prestressed masonry is that, because it has high durability, maintenance costs are minimized ([2], [9]). Additionally, good insulating properties can reduce energy costs ([2], [27]). At this time it should be noted that controlling the cost of prestressing hardware and minimizing the amount of restraint necessary to provide structural integrity are keys in keeping this construction procedure economical ([28], [32]).

In terms of loadbearing capacity, prestressed masonry also has many positive aspects. Since the masonry is designed to remain crack-free under service loads and because it is used more efficiently, both thinner elements, as well as taller walls, become feasible ([8], [17], [27], [29]). Because of the applied compressive stress, the shear strength of the system is enhanced to the point that shear reinforcement may be eliminated in many cases [17]. Flexural behavior is also enhanced and ductile response is observed through large displacement capacity and the tendency for the structure to return to its initial state after loading ([14], [17]). This self-righting nature of prestressed masonry proves to be of primary significance when seismic resistance is considered [23].

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Structural advantage can also be obtained through anchorage at the top and bottom of the wall. These anchorages can provide fixity at the base and can also provide effective connections between walls and diaphragms ([9], [31]). The final potential advantage that should be considered is the high level of fire resistance of a well-designed and built prestressed masonry wall ([2], [27]). 1.3 Applications of Prestressed Masonry Now that the elements of a prestressed masonry system and its strengths have been discussed, the way in which this system has been employed, as well as how engineers and builders may be able to apply this technology in the future, can be examined. As mentioned initially, prestressed masonry exhibits the most benefits when applied to situations in which axial loads are low and out-of-plane lateral loads, such as wind, earth pressure, and ground acceleration due to earthquakes, are significant. Because of this, most applications at this time have been for low-rise wall construction. These low-rise walls are common to residential construction, commercial buildings, institutional buildings, retaining walls, and sound walls ([2], [10], [12], [13], [22], [27], [29]). Since these are all common structures, there should be broad opportunities to readily implement this system. Construction of other non-building structures, such as bridge abutments, water tanks, and grain silos, has taken place, and these applications show the versatility of prestressed masonry ([2], [9], [10], [22], [27]). Specific elements in newly built structures can also benefit from the use of a prestressing system, including veneers, lintels, columns and partition walls ([13], [15]). In addition to new construction, prestressed masonry technology is proving its worth when implemented in the restoration, retrofit, and repair of existing unreinforced structures, as well as conventionally reinforced masonry structures. When used in upgrading of structures, the original character of the building can be maintained while the structure is strengthened to the point of accomplishing an effective seismic retrofit ([2], [9], [10], [17], [21], [22], [27], [31]). Additional applications that are being considered include prefabricated wall and veneer panels, which can be manufactured off-site and delivered to the project ready to install, floor panels, beams, and infills for large frames ([10], [13], [29]). 2. BACKGROUND INFORMATION Previous research at the University of Minnesota addressed the stability of unreinforced masonry (URM) walls subjected to compression and flexure ([2], [24], [26]). Initially, an elastic stability solution was developed which was then used to conduct a parametric study to explore the impact of this issue. The work revealed that a broad range of URM walls have a higher potential for buckling instability failure than for any other failure mode currently addressed in design codes ([24], [26]). This finding revealed the need to study the problem further; therefore, an experimental investigation of the stability of slender URM walls loaded axially, as well as laterally, was initiated. The experimental work was found to validate the elastic stability solution ([2], [26]). Buckling failures were found to occur in all experimental cases with noticeable

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second order effects (i.e., P-∆ effects) in which the external axial load acted eccentrically when out-of-plane deflections were imposed on the walls ([2], [26]). Following this finding, a series of related question were posed in regard to the stability of post-tensioned masonry (PTM) walls [2]: Does prestressing force act as a stationary external axial load or as a follower load? If the former, does it mobilize member instability (i.e., buckling)? Does tendon restraint affect the stable or unstable response of PTM walls to lateral loading? These questions are of great importance to the development of prestressed masonry because its most likely application in the U.S. is for tall walls in single-story commercial buildings (e.g., warehouses, discount stores, etc.). To begin answering this question, a literature review of experimental testing of PTM walls subjected to lateral loads was undertaken [2]. Nine experimental programs conducted between 1987 and 1998 were discovered ([1], [7], [11], [12], [15], [16], [21], [22], [30]). Data on a total of 84 PTM specimens, constructed of clay bricks, concrete blocks, and calcium-silicate units, was available from various tests. The compiled experiments had been conducted almost entirely on stocky walls and explored variables that govern the behavior of prestressed masonry, specifically: level of prestress, type of tendon restraint, and axial load interaction. By considering these variables, researchers have been able to refine and validate current understanding of how certain PTM walls can be expected to behave under service and ultimate load conditions [2]. 2.1 General Behavior Prestressed masonry walls subjected to lateral loads have been shown to display two distinct phases of behavior. The first phase is defined by approximately linear elastic response to loading until the wall cracks ([3], [8]). The cracking at this point typically occurs at the bed joints when both the compressive prestressing force and the bond strength between the mortar and unit are overcome. Post-cracking behavior, the second phase, features nonlinear load-displacement response until failure ([3], [8]). Ultimate strength and failure of the specimen is typically controlled by how well the reinforced section can maintain its integrity and resist the external forces under increasing compressive strains. While the above behavior was found to be true in past studies, variables such as the initial level of prestress, the level of tendon restraint, and axial load interaction contributed to behavior both before and after cracking occurred. A variable that still remained to be addressed, however, is wall slenderness. From the tests on stocky walls, the following relationships were observed. First, the level of prestress primarily impacts the range through which walls are able to maintain elastic behavior [8]. Higher levels of prestress result in higher cracking loads and vice versa. How the level of prestress affects the cracking load has been found to be similar for walls using different restraint conditions [8]. Also, prestressing force affects post-cracking behavior, but here results vary depending upon the restraint condition of the tendon. For walls with restrained tendons, a specimen with a lower level of prestress will typically experience more deformation before reaching its ultimate strength, but will display a capacity similar to a specimen with more

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prestress [8]. In a wall that uses unrestrained tendons, however, the ultimate strength can be significantly diminished when lower levels of prestress are applied [8]. This discrepancy points out the importance of understanding why walls with restrained and unrestrained tendons behave differently. In walls with restrained tendons, because they maintain their location within the cross-section, the effective depth of the tendon stays constant. Also, since there is no additional eccentricity of the prestressing force due to member deformation, the lack of eccentricity does not contribute to second order effects (i.e., P-∆ effects), which can lead to instability ([2], [10], [11]). On the other hand, unrestrained tendons are able to displace within the cross-section and are susceptible to instability and buckling due to the eccentric positioning of the tendon, which causes P-∆ effects ([2], [10], [31]). Overall, the use of tendon restraints is desirable and produces masonry walls that display better post-cracking behavior with improved strength and ductility as compared to walls built using unrestrained tendons ([27], [28]) Instability is an issue that arises when axial loads are applied externally to slender walls and, in the case of prestressed masonry, it may be worsened when unrestrained tendons are used. As a wall with an externally applied axial load deflects, the axial load acts eccentrically and causes second order effects, similar to those a prestressed wall using unrestrained tendons experiences, and this can lead to buckling ([10], [11], [17]). Since walls with considerably high slenderness ratios are currently being built, there is concern that these designs may be unsafe, especially in seismic regions and under high winds [17]. In order to further understand the extent to which instability is a concern for slender post-tensioned masonry walls, an experimental study was needed. Determining the limits for which prestressed masonry is applicable is very important; through research these limits are being defined while at the same time behavior is being revealed. A primary desirable aspect of behavior is the flexural stiffness prestressed masonry walls possess under service conditions prior to cracking ([8], [22], [27]). Additionally, when the cracking moment is reached, lateral loads are higher and cracking is less extensive than it would be if the wall were built using other masonry construction methods ([8], [27]). Also, after a prestressed wall has experienced excessive loading up to its nominal strength, the post-tensioning provides a restoring effect, which closes the cracks and limits residual deformations ([8], [9], [17], [27]). Finally, when ultimate strength is required, post-tensioning serves to increase wall capacity ([14], [22], [27]). 2.2 Objectives The objectives of this project are to investigate and demonstrate by experimentation the strength, toughness, and stability of slender PTM walls subjected to out-of-plane lateral loads. The experimental program at the University of Minnesota focused on the impact of masonry type (concrete block or clay brick), tendon restraint condition (restrained or unrestrained), and levels of effective prestress on the lateral load behavior of these walls. An overall goal of the research is to propel an innovative, cost-effective, and reliable loadbearing masonry wall system to the forefront of engineering practice by means of improved code provisions.

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3. EXPERIMENTAL PROGRAM 3.1 Variables Because of the questions concerning the stability of slender post-tensioned masonry walls, very slender walls were selected for this investigation. The twelve specimens, selected for this study, had a target height to thickness ratio (h/t) between 35 and 40. While the slenderness of each specimen was held constant, other aspects of the walls were varied, including the masonry type, tendon restraint conditions, and effective level of prestress. Regarding masonry type, since MSJC code provisions address the nominal capacity of both concrete and clay post-tensioned masonry walls through a single set of equations, the question arises as to whether or not the behavior of these walls is sufficiently similar to be accurately predictable for both types of masonry [34]. To address this point, the specimens used in this experiment were built using both types of masonry. Of the twelve post-tensioned masonry walls constructed, half were built using concrete block units while the other half were built using clay brick units. Tendon restraint condition was the next variable incorporated into the specimen test matrix. Current MSJC provisions allow for the use of unbonded tendons that are either unrestrained or restrained in the design of post-tensioned masonry walls, but the provisions address the issue of nominal strength only when the tendons are restrained. To concentrate on this issue, half of the specimens used unrestrained tendons, three concrete and three clay specimens, while the other half used restrained tendons, again three specimens of both types of masonry. The final variable included in this portion of the study was the effective level of prestress that acts on the net cross-sectional area of the wall. The analysis of previous experiments indicated that current MSJC provisions could underestimate moment capacities when low magnitudes of prestress are used [34]. Three different levels of prestress were selected for this experimental program, 35 psi, 75 psi, and 150 psi. Each level of prestress was applied to an unrestrained and a restrained wall of each type of masonry. By selecting this range of prestress, it is possible to explore the low magnitudes of prestress as desired while also addressing effective levels of prestress that are more applicable in the design of post-tensioned masonry walls. To differentiate the variables for each specimen in this study, the designations shown in Table 3.1 were used. First, the wall was denoted as being a post-tensioned concrete block wall (PC) or a post-tensioned clay brick (PB) wall. The subsequent digit defines the number of the wall in the test series. The value listed in the middle of the designation denotes the effective stress, in psi, placed on the net area of masonry, prior to testing, through the post-tensioning rods (35 psi, 75 psi, or 150 psi). The last letter of the designation classifies the tendon restraint condition as unrestrained (U) or restrained (R).

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Table 3.1 Specimen Test Matrix

Prestressing Level Restraint Condition 35 psi 75 psi 150 psi Unrestrained Restrained

PC1-35-U X X PC2-75-U X X

PC3-150-U X X PC4-35-R X X PC5-75-R X X

Con

cret

e B

lock

W

alls

PC6-150-R X X PB1-35-U X X PB2-75-U X X

PB3-150-U X X PB4-35-R X X PB5-75-R X X C

lay

Bric

k W

alls

PB6-150-R X X 3.2 Primary Elements of Specimens The slender post-tensioned masonry walls designed for this experimental program incorporated the primary elements that are used in this type of construction, concrete and clay wall sections, with additional features to facilitate testing: header and footer beams, prestressing tendons, external tendon restraints, and lateral loading sleeves. The header and footer beams were constructed using a steel plate and a concrete bond beam. The steel plate, shown in Figure 3.1, was incorporated into the design to allow the specimens to be bolted to the top and bottom pin of the load frame, to position the post-tensioning rods at the centerline of the wall at the top and bottom, and to provide a bearing surface for the post-tensioning rods. The concrete bond beam, shown in Figure 3.2, which had a height that was equivalent to a single course of concrete blocks, as well as three courses of clay bricks, served to effectively transfer the post-tensioning force from the tendons to the masonry cross-section. The bond beam design incorporated ducts at each end that allowed the post-tensioning rods to pass through allowing the rods to be unbonded along the full height of the wall. Each bond beam was attached to a steel plate through the use of Hilti RM 700 EP sand-epoxy mortar.

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Figure 3.1 Steel Plate

The prestressing tendons used for all of the specimens were ASTM B7 threaded rods, which have nominal yield strength of 105 ksi and a nominal tensile strength of 125 ksi. Different diameters of threaded rod were used depending upon the effective level of prestress applied to the wall. The diameters were 1/4", 3/8”, and 1/2" for prestress levels of 35 psi, 75 psi, and 150 psi respectively. The net cross-sectional area of the rod was 0.0395 in2 for each 1/4" diameter rod, 0.0860 in2 for each 3/8” diameter rod, and 0.1562 in2 for each 1/2" diameter rod. Two external post-tensioning rods, one on either side, were used to apply the prestressing force to every specimen.

Figure 3.2 Concrete Bond Beam

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Figure 3.3 External Tendon Restraints

External tendon restraints were incorporated into the design of the specimens to simulate the inside surface of the masonry cavities since an unrestrained tendon condition would be difficult to simulate within a four-inch nominal unit. The external restraints were scaled to represent the dimensions of an eight-inch unit. Additionally, this restraint system allowed for the observation of tendons during out-of-plane loading. As shown in Figure 3.3, the restraints were designed to wrap around the edge of the wall, with six restraints being placed along each edge of the wall at five primary elevations (two restraints were used mid-height, with one just above and one just below the central bed joint). In the specimens that incorporated unrestrained tendons, the restraint cavity was left open so that the tendon was able to displace relative to the wall until it came into contact with the inside edge of the restraint. Alternately, specimens with restrained tendons used a wooden block placed within the restraint cavity to maintain the position of the tendon within the restraint. A hole that was 1/16” larger than the diameter of the threaded rod used for each specimen was drilled in the center of each block to allow for the installation of the tendons, while still requiring the tendon to displace in conjunction with the wall during flexural testing. Metal sleeves were located in select bed joints of each wall specimen to enable connection to the whiffletree system, which generated the lateral loading during testing. Copper tubes were used for the sleeves, which enabled placement of ¼” threaded rods through the wall to attach loading channels to both faces of the specimen. The loading channels were attached at four levels along the height of the wall using two threaded rods per level. The placement of the sleeves was such that both the concrete block walls and clay brick walls were loading at the same elevations. These elements were incorporated into the construction of the concrete and clay wall sections as seen in Figure 3.4. Four-inch units were used to generate the required slenderness for the wall specimens within the height limitation of the laboratory. The units were laid using Type S Portland cement-lime mortar to build masonry sections that had an average width of 31.63”. The concrete sections were face-shell bedded and had a net cross-sectional area of 82.6 in2, while the clay sections were fully bedded and had a net cross-sectional area of 78.4 in2. The masonry sections were built to have an average height of 112.63” (14 courses of concrete units, or 42

15


courses of clay units) and a measured width of 31.63”. Since the masonry section of each wall was between a header and a footer beam, the resulting average height from the top plate to the bottom plate was 130.75” with a standard deviation of 0.18”. Additionally, once the specimens were bolted in place and ready to be tested, the resulting clear height between the supports, pin-to-pin height, was 139.25”. Therefore, the final height to thickness ratio (h/t) for these specimens was 38.4 for the concrete block and 40.5 for the clay brick walls.

Figure 3.4 Specimen Details

3.3 Construction of Specimens To facilitate the construction of the concrete and clay wall sections, wooden backdrops were constructed for each specimen. Nominal 2 x 4 lumber and 3/4" plywood was used to construct the backdrops, which were built on top of bases made using nominal 4 x 4 lumber. The width of each backdrop was 31.63” and served to define the width of the wall while the height of the backdrop was 110.5”, just less than the height of the masonry wall sections. After plumbing the backdrop and leveling the base, the first step in the construction of the specimens was to position a footer beam on each base. The footer beam served as the base of the

16


wall where masonry units would be laid by professional masons from the Minnesota Bricklayers and Allied Craft Workers Local Union #1. The first course of each specimen was laid, as illustrated in Figure 3.5, on Friday, May 30th, 2003. This course was laid using Hilti RM 700 EP sand-epoxy mortar to provide a good bond between the concrete bond beam and the masonry units. On Monday, June 2nd, 2003, the masons laid seven of fourteen courses on the six concrete block walls, 14 courses on three of the clay brick walls, and 21 of 42 courses on the remaining three clay brick specimens. The following day, Tuesday, June 3rd, 2003, the masons finished constructing the masonry sections for each specimen, which is demonstrated in Figure 3.6.

Figure 3.5 Photograph of Laying First Course of a Clay Brick Wall

Figure 3.6 Photograph of Concrete Block Wall Construction

Some details of this work should be noted. First, the copper sleeves for the lateral loading system were placed in the bed joints while the units were being laid. Only the exposed face was tooled as the wall was constructed. The masonry sections were braced to the backdrop at three points along the height of the wall for support. Once the concrete block and clay brick sections had cured for several days, the header beams were placed. As with the footer beams, this bond was made using sand-epoxy mortar. The beams were set on this bed of mortar while shims were used to level each one. After the header beam was in place, post-tensioning rods were run through the header and footer beam and then

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were tensioned slightly. This stabilized the beam while the mortar set and also placed the specimen in slight compression.

While the specimens were still braced against their backdrops, the tendon restraints were fabricated and attached along both edges of each wall. To facilitate connecting each restraint to the wall, a hole was drilled through the restraint as well as through the wall, which allowed a 1/4” threaded rod to be used to clamp the restraint in place. Additionally, each restraint was adhered to the wall using Sikadur 32 Hi-Mod LPL epoxy. Fully constructed specimens are shown in Figure 3.7.

Figure 3.7 Photograph of Concrete and Clay Wall Specimens

After the primary elements of each specimen were in place, the walls were moved to a storage area. The walls were stored four deep, and as each wall was moved into storage it was braced at three points along the height of the masonry section, using one of the three remaining backdrops for support. In addition, the tops of the walls were tied together to provide further bracing. Once the walls were in this storage area and they had cured adequately, they were prestressed. The initial level of prestress was measured using two load cells on each of the post-tensioning

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rods. The first walls to be stressed were the six concrete block walls. This was done between December 15th and December 18th, 2003. Later, on April 22nd, 2004, the six clay brick walls were prestressed. 3.4 Load Frame To test the slender post-tensioned masonry walls constructed for this experimental investigation, the load frame shown in Figure 3.8 was assembled. This load frame consisted of two primary systems, an axial loading system and a lateral loading system. The axial loading system was used to accommodate the specimen to be tested in the simply supported, vertical orientation pictured in Figure 3.9, and to allow vertical deflection of the walls while applying a small, but finite level of vertical load (i.e., load control with a load signal = 100 lbs). The lateral loading system, on the other hand, facilitated the application of a lateral moment distribution that simulates a uniformly distributed lateral load.

Figure 3.8 Schematic of Load Frame

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Figure 3.9 Vertically Oriented, Simply Supported Specimens

3.4.1 Axial Loading System For the axial loading system, a four-column frame, shown in Figure 3.8 and 3.10, was used. The columns were 24’ tall W12 x 120 steel sections bolted through the strong floor and oriented such that the weak axis of each column resisted the lateral load applied to the walls. At the top of these columns, two 34” deep beams were connected between the east set and the west set of columns. Joining these beams together was the top actuator support beam, a W14 x 159 section. A single 77- kip MTS actuator was oriented vertically in the center of the four-column frame. The top of the actuator was attached to the W14 x 159 section, which acted as a spacer between the actuator and the actuator support beam so the proper elevation was achieved for the top of wall connection, before being bolted to the actuator support beam. The orientation of the actuator was such that the pin could swivel in the east-west (i.e., out-of-plane) direction. To facilitate the installation of the vertically oriented, simply-supported specimens into this load frame, as presented in Figures 3.8 and 3.9, double webbed I-beams attached to pins were used both bottom and top. At the bottom, the double webbed I-beam was bolted to a foundation poured on and connected to the strong floor in the center of the frame. At the top, the double webbed I-beam was bolted directly to the 77-kip MTS actuator. This section was then braced laterally to prevent transverse motion. Both the top and the bottom pins were bolted to these sections.

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Figure 3.10 Schematic of Load Frame Plan View

The pins used in this setup are illustrated in Figures 3.8 and 3.9. They were Link-Belt Company high capacity pillow block spherical roller bearings that incorporate a 2” high-strength, hardened steel shaft. The pins have a capacity of 44.4 kips and are self-aligning to 5 degrees of in-plane shaft rotation. The two sets of pins were welded to 1” steel plates both top and bottom to allow for a bolted connection to the double webbed I-beams, as well as the connection to the wall specimens. Another key component of the load frame, as viewed in Figure 3.8, was a shelf piece consisting of a W18 x 71 section bolted between the west set of columns at an elevation that coincides with the top double webbed I-beam. This shelf piece served as the attachment point for restraint brackets which allowed for two ¾” x 4” aluminum braces to be connected to the top double webbed I-beam. One end of each brace was connected to the double webbed I-beam with two bolts, while the other end was connected to a vertically slotted hole in the restraint bracket using a single bolt. The single bolt was then fastened with two nuts so that it could act as a pin connection. Additionally, a restraint block was connected to each brace, made of the same material as the brace, to bear against the restraint bracket in order to control rotation. This restraint detail functioned as a simply supported end condition, allowing the top of the wall to move vertically while resisting lateral displacement. Rigidity of this frame was gained through the use of 6” x 6” tube sections with a wall thickness of 3/8”. The tube sections were bolted between each set of east-west columns at the mid-height of the shear beams. When viewed from above, as is the case in Figure 3.10, the tube sections in conjunction with the shear beams effectively created a closed section.

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The final aspect of the four-column frame that should be mentioned was the implementation of crane rails with attached hoist, which were used to move specimens in and out of the load frame. Two W8 x 31 sections served as the rails and were oriented in the east-west direction within the frame. The rails were set at an elevation between the shear beams and the shelf piece. A single CM Series 622 hoist was attached to each of the crane rails, with a capacity of two tons for each hoist.

3.4.2 Lateral Loading System The primary structural component of the lateral loading system was a 20’ tall W12 x 106 column shown in Figure 3.8. The column was attached to the strong floor and braced mid-height through the use of a W18 x 65 section. The brace created a 45-degree angle with the column and was attached directly to the column at top and to a transfer section attached to the strong floor at the bottom. Both the column and the brace were oriented so the strong axis of the section resisted the lateral load delivered to the specimen. Figure 3.8 depicts the location of the single 35-kip MTS actuator, which was used to control the lateral displacement of the specimen during testing. This actuator was positioned horizontally and was directly connected to the main column. A W12 x 65 stub column, bolted to the strong floor, supported the body of the actuator. The lateral load generated by the actuator was delivered to the wall through the use of the whiffletree system shown in Figure 3.8 and detailed in Figure 3.11. This system comprised spreader beams, threaded rods, and cylindrical washers. A vertically oriented T6 x 6 x 1/4 spreader beam was connected to the actuator. Two T4 x 4 x 3/16 spreader beams, also positioned vertically, connected to each end of the larger section through 1” threaded rods. Next, two horizontally positioned T2 x 2 x 3/16 spreader beams were connected at each end of the two preceding spreader beams using ½” threaded rods. Load cells were then connected to both ends of these four spreader beams using ¼” threaded rods, which further connected directly to the wall. For the sections of ¼” threaded rod spanning between the load cells and the wall, a tube section with a ¼” inner diameter and 1/8” wall thickness was used to sleeve the rod. This tube section prevented buckling of the ¼” rods if the load cells went into compression when the walls became unstable. This arrangement allowed the walls to be loaded at four levels along the vertical plane through eight distinct points. At each of the four loading levels, C4 x 5.4 sections were placed on both faces of the wall to allow the point loads to be distributed across the full width of the specimen. Additionally, ¼” neoprene pads were placed between the wall and the channels to prevent a concentrated load from being placed in any location. The channels on each face were connected using cables following the perimeter, which were fastened to the top and the bottom plate of each specimen during testing. The cables were kept loose so as not to interfere with the behavior of the specimen during testing, but were incorporated to serve as a containment system during any unexpected failures.

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Support for the whiffletree system came from two elements, support rails and wire rope. First, the main branch, the T6 x 6 x 1/4 support beam, rode on two C5 x 9 support rails, which were clamped around the stub column and supported by two sets of C3 x 5 sections oriented perpendicularly to the support rails. C3 x 5 sections were also clamped around the T6 x 6 x 1/4 support beam and served as the glides that made contact with the rails. To reduce the friction that is inherent in this type of system, two 1/8” Teflon sheets were placed between the support rails and the glides so that the sliding surface was Teflon on Teflon. The second support system for the whiffletree came from wire rope suspended from the east shear beam in the four-column load frame and attached to the two T4 x 4 x 3/16 spreader beams. This support system allowed the whiffletree to act as a series of simply supported beams as designed, while closely achieving horizontal force equilibrium between the horizontal actuator and the load cells attached to the specimen.

Figure 3.11 Schematic of Whiffletree System

3.5 Instrumentation When the instrumentation plan was developed, the goal was to measure loads and specimen response using available instruments or instruments fabricated to be compatible with the available data acquisition system. This was done by using internal load cells and LVDTs (Linear Variable Displacement Transducers) within each actuator, load cells within the whiffletree system, load cells on the post-tensioning rods, horizontal and vertical LVDTs at various locations along the height of the wall, horizontal LVDTs connected to the tendons, and tilt meters. Through the use of these instruments, which are further described below, it was possible

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to measure applied loads, stress in the tendons, wall deflection, expansion and contraction of the wall at mid-height, and rotation at the top and bottom of the wall. 3.5.1 Load Measurement The axial load in each specimen came from a combination of post-tensioning force and force applied using the vertical actuator. Both sources were monitored during testing, so the total axial load in the specimen was known. The post-tensioning force was measured during the application of post-tensioning and the experimental testing of the wall specimen by using two load cells placed at discrete locations, shown in Figure 3.4, on each post-tensioning rod. The load cells were fabricated using couplers that were compatible with the post-tensioning rods. Strain gages were glued to opposite faces of the hexagonal couplers, wires were soldered connecting the gages, and a one-MΩ resistor was soldered into the circuitry to create a null balance Wheatstone bridge circuit. These load cells were also coated with a protective covering to prevent them from being damaged. Depending on the size of the coupler each load cell was calibrated in the laboratory’s MTS-200 kip machine to the level reported here: 1.6 kips for ¼” couplers, 4.4 kips for 3/8” couplers, and 8.8 kips for ½” couplers. Axial load delivered to the specimen through the MTS 77-kip actuator was measured using an internal load cell. Lateral load was applied to the specimen at eight points, yet all loads came from a single source, the horizontal actuator. To facilitate a static analysis, lateral loads were measured at each of the eight load points by using load cells connected between the wall and the whiffletree system. These load cells were fabricated out of aluminum round stock with three strain gages attached. A sealed steel tube, which served as a protective covering, enclosed the aluminum rod and strain gages. Each of these load cells had been calibrated to 2 kips. Total applied lateral load could be checked based on the load applied by the MTS 35-kip actuator, which was measured using an internal load cell. 3.5.2 Displacement Measurement Lateral displacement of the specimen was measured using a combination of 1”, 2”, and 5” LVDTs placed along the wall centerline at the seven different elevations, labeled H-C1 through H-C7. Four 1” LVDTs were used to measure displacement at the top and bottom plate as well as at the highest and lowest loading channels. Two 2” LVDTs were used to measure displacement at the loading channels just above and below the mid-height of the wall. A single 5” LVDT measured mid-height displacement. In order to monitor tendon movement relative to the wall, 1” LVDTs were attached to each tendon at mid-height. The placement of these LVDTs is defined in Figure 3.12 by points T-N4 and T-S4. Vertical deformation along the mid-height of the specimen was monitored using LVDTs positioned vertically on both the compression face and the tension face, as depicted in Figure 3.12. These LVDTs were used to monitor cracking, as well as to ascertain axial strain. Three 0.5” LVDTs, denoted V-C1, V-C2, and V-C3, were placed on the compression face 3.5” to either side of the wall centerline. An 8” gage length was used for each LVDT, which resulted in deformation being monitored 12” above and below the mid-height of the wall. On the tension

24


face, a total of seven LVDTs, labeled V-T1 to V-T7, were used. The three 0.5” LVDTs were positioned so that they mirrored those attached to the compression face. Similar to the other LVDTs, the four additional 0.1” LVDTs were offset 3.5” from the wall centerline and had a gage length of 8”. These LVDTs were positioned so that two were placed above the 0.5” LVDTs and two below the 0.5” LVDTs. This arrangement allowed for 28” above and below mid-height to be monitored during testing. It should be noted that an extra 0.1” LVDT was installed on the compression face during testing if joint opening occurred in a region that was not monitored by the initial configuration on the compression side. Additionally, an internal LVDT within the vertical MTS 77-kip actuator, as well as one within the horizontal MTS 35-kip actuator were used to measure displacements.

Figure 3.12 Schematic of LVDT and Tiltmeter Diagram

3.5.3 Rotation Measurement Tiltmeters were attached to both the top plate and the bottom plate of the wall to measure rotation. The tiltmeters were oriented as seen in Figure 3.12. Both tiltmeter units were Applied Mechanics Model 800, which measure an effective rotation of +/- 0.5 degrees.

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3.5.4 Data Collection Because of the number of channels needed to record the data from all of the instrumentation, two 16-bit data acquisition systems were used, MegaDAC 3008AC and National Instruments SCXI-1000. While the majority of the data was read into the MegaDAC system, the 0.1” vertical LVDTs, as well as the tilt meters, fed into the National Instruments system. During testing, both systems were synchronized, and they sampled and recorded data at a frequency of 2 Hz. 3.6 Loading Procedure Prior to the commencement of a test, the specimen was installed in the four-column load frame and oriented so that the finished side of the wall would be subjected to flexural tension during testing. As for the unfinished flexural compression side, all excess mortar was removed before the loading channels were connected to the wall and the instrumentation was set up. Additionally, all temporary bracing was removed so the specimen was simply supported. To begin, the vertical actuator was set to maintain an axial load of 100 pounds on the specimen throughout the test. The purpose of this load was two-fold. First, a small ‘seating’ load was desired throughout the tests to ensure bearing of all components the in the vertical direction. Second, this setup enabled the wall to expand and contract vertically during loading without any change to this small vertical load. The vertical stress on the net area of masonry corresponding to this load did not exceed 2 psi. Once the vertical load had been established, all stress in the post-tensioning rods was released and then tensioned again to the desired level of effective stress to minimize losses. After the proper amount of stress was placed on the specimen through the post-tensioning rods, the specimen was connected to the whiffletree and adjustments were made so that lateral load would be applied evenly along the height of the wall during testing. The lateral load was controlled by the 35-kip horizontal actuator, which applied a monotonically increasing lateral displacement. The loading rate used was 0.05 in/min for specimens with effective stress levels of 35 psi and 75 psi and 0.1 in/min for the 150 psi specimens. Each test was continued beyond the nominal capacity of the wall in order to determine how the specimen behaved as capacity was exhausted. Once the loading stage had concluded, the horizontal actuator was again operated in displacement control to allow the specimen to be unloaded in a controlled manner. 3.7 Modifications to Test Setup and Instrumentation The first two walls tested were PC2-75-U and PC5-75-R. Following both of these tests, modifications were made to the original experimental program. Following the first test, PC2-75-U, two minor modifications were made to the load frame, with one further load frame modifications after the second test, PC5-75-R. Additionally, after testing PC5-75-R, additions were made to the instrumentation plan.

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The first modification to the load frame was in regard to a difference in the horizontal equilibrium between the load measured in the actuator and the load cells in the whiffletree system. This difference was found to be caused by friction between the whiffletree support rail and the whiffletree glides. To reduce this discrepancy, Teflon sheets were placed between the support rail and the glides so the sliding surface was Teflon on Teflon. Additionally, after the first test, it was observed that as the load in the middle branches of the whiffletree went into compression the 1/4” rods connecting the load cells to the wall had a tendency to buckle as the compression force increased. To alleviate this problem, tube sections with a 1/4” inner diameter and a 1/8” wall thickness were incorporated into the last branch of the whiffletree. The tubes were used as sleeves for the 1/4” rod, and they effectively increased the range of testing since the lateral loading system can carry additional levels of compressive force. With these two modifications made following the initial test, wall PC5-75-R was tested. After this test, one additional modification was made to the load frame, as well as an addition to the instrumentation plan. The alteration made to the load frame addressed rotation of the top pin due to the weight of the original 1” x 4” steel braces that cantilevered between the double webbed I-beam and the restraint bracket attached to the shelf piece. To control this rotation, the steel braces were replaced with 3/4” x 4” aluminum braces, which also incorporated restraint blocks that beared against the restraint brackets. To verify that this detail did not change the restraint condition of the top pin, this wall was retested before another specimen was tested. The new detail was found to be acceptable. Regarding modifications to the instrumentation plan, two alterations were made. First, during the testing of Wall PC5-75-R, joint opening occurred in a location where there were no vertical LVDTs. Originally, only three 0.5” LVDTs were being used to monitor the 12” above and below the mid-height of the wall on both the tension and compression face. To increase this range, the five 0.1” LVDTs were added into the instrumentation plan to monitor the expansion and contraction along the faces of the wall. Four of the LVDTs were used on the tension face, while one was made available to be placed on the compression face if joint opening were to occur in a location other than where the 0.5” LVDTs were located. The second change concerned instrumentation of the tendons, which was not incorporated into Wall PC5-75-R. After this test, it was determined to be prudent to measure the tendon displacement relative to the wall during the remaining tests, so tendon location could be confidently defined. It should be noted that all of the information presented prior to this section describes the modified test setup. All of the walls except PC2-75-U and PC5-75-R were tested using the modified test setup. These modifications were the only ones that were made during the experimental program and were done to refine portions of the load frame and the instrumentation plan used for testing.

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4. MATERIAL PROPERTIES Material sampling and testing was done on the concrete and clay units, the mortar cubes, and the prisms constructed from each type of masonry to determine the properties of the materials. The units were tested after they had been maintained in the structures laboratory throughout the construction process. The mortar cubes and the masonry prisms had been made while the walls were being constructed to provide a representative sample of the materials used. All of the tests on the masonry units and the mortar cubes were conducted using equipment available in the Structures Laboratory at the Department of Civil Engineering at the University of Minnesota, including a Forney machine used for compression testing. While the prism flexural bond tests were conducted using a bond wrench frame in the structural laboratory, the facilities at Twin Cities Testing were used for all prism compression testing. 4.1 Unit Properties From September to December of 2003, the properties of both the clay brick units and the concrete masonry units used for the construction of the walls were determined. Following ASTM C 140-02 (Standard Test Methods for Sampling and Testing Concrete Masonry Units and Related Units) the percent solid, density, absorption, moisture content, and compressive strength were determined for the concrete units. ASTM C 67-02a (Standard Test Methods for Sampling and Testing Brick and Structural Clay Tile) was used to determine the percent solid, weight per unit area, absorption properties, and compressive strength of the clay units. Once testing of the units was completed ASTM C 90-02 (Standard Specification for Loadbearing Concrete Masonry Units) and ASTM C 62-01 (Standard Specification for Building Brick [Solid Masonry Units Made from Clay or Shale]) were used to classify the concrete units and clay units respectively. For the concrete masonry units, the average result of tests conducted on five units was used to determine all reported properties except for the compressive strength, when the average result from three units was used. The units were found to be 70.61 percent solid and to have a density of 132.15 lb/ft3. The measured water absorption for the units was 7.19 lb/ft3 while the in-situ moisture content of the units was 11.56 percent. The compressive strength of the units was determined to be 3,520 psi. At the conclusion of these tests, the units were accordingly classified as hollow, normal weight concrete masonry units. Shown in Figure 4.1, the average dimensions of the units, as measured for five blocks, were 7.65 inch (std. dev. = 0.02 inch) by 15.6 inch (std. dev. = 0.02 inch) by 3.64 inch (std. dev. = 0.008 inch) for the height, length, and width, respectively. tf = 1.23”

L = 15.6”

w = 3.64”

Figure 4.1 Dimensional Properties of the Concrete Block Units

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Reported results for the clay brick units are based on the average results of five units. The units were determined to be 75.56 percent solid and have a weight per unit area of 25.63 lb/ft2. Regarding the absorption properties of the units, the initial rate of absorption (IRA), the twenty-four hour submersion (C), the five hour boiling (B), and the resulting saturation coefficient (C/B) were determined. The IRA test showed the units to absorb 2.17 grams per minute per 30 in2. C and B were found to be 2.8 percent and 3.2 respectively, resulting in a C/B value of 0.85. The final property that was measured for the clay units was the compressive strength, which was found to be 12,850 psi. These results were used to properly classify the clay bricks as solid, grade MW units. The average dimensional properties of the clay brick units, as measured for ten units, were a height of 2.28 inch (std. dev = 0.01 inch), length of 7.53 inch (std. dev. = 0.02 inch), and width of 3.44 inch (std. dev. = 0.04 inch), as depicted in Figure 4.2.

L = 7.53”

w = 3.44”

Figure 4.2 Dimensional Properties of the Clay Brick Units

4.2 Mortar Properties ASTM C 270 (Standard Specification for Mortar for Unit Masonry) was used to determine the proportions used for the mortar mix. A Type S mortar with the following material proportions by volume was used: 1 part Portland cement, 1/3 part lime, and 3-1/2 parts sand. Water was added as specified by the mason. Nine batches of mortar were mixed during the construction of the walls and mortar cubes were made for five of the batches. Cubes and compression tests were conducted following ASTM C109/C 109M-02 (Standard Test Methods for Compressive Strength of Hydraulic Cement Mortars [Using 2-in. Cube Specimens]). Testing of the compressive strength was done at 3 days, 7 days, and 28 days for each set of cubes. The 28-day compressive strength for four of the batches was consistent and produced an average value of 1,430 psi. One of the batches, which had been noted as being much wetter than typical, only reached a 28-day strength of 900 psi. 4.3 Prism Properties Since the properties of masonry units differ significantly from those of the mortar used to bond them, it is necessary to determine the properties of the unit and mortar assemblies, or prisms. Prism properties that were relevant during this study were masonry compressive strength, modulus of elasticity, and flexural tensile strength. The MSJC code was used to predict each of

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these properties while material testing was also conducted to determine the measured behavior of the specimens [20]. 4.3.1 Masonry Compressive Strength, f’m From the MSJC code, the compressive strength of masonry can be predicted based on the compressive strength of units and the type of mortar used in construction. For the concrete masonry units, with a unit compressive strength of 3,520 psi and using Type S mortar, the predicted masonry compressive strength is 2,380 psi. For the clay brick units, having a unit compressive strength of 12,850 psi and using Type S mortar, a masonry compressive strength of 3,950 psi is predicted. To determine the actual compressive strength of the block and brick prisms, two sets of tests were conducted. The first set of prisms was tested to determine the 28-day strength of both types of masonry assemblages. Eleven concrete block prisms (two units tall) as well as eleven clay brick prisms (five units tall) were tested at this time. The second set of tests was conducted on April 26, 2004 during the time the wall specimens were being tested. Six prisms of each masonry type were tested at this time and reconfirmed the strengths measured during the first set of tests. Seventeen block prisms and seventeen brick prisms were tested in compression. Before the average compression strength of the prisms was calculated, data that fell outside of plus or minus two standard deviations of the full data set was excluded. This resulted in one of the original data points being excluded from each set of prisms, concrete block and clay brick. For the block prisms, the excluded data point was below the average minus two standard deviations, while for the brick prisms, the outlier was above the average plus two standard deviations. The average strength measured for the concrete block prisms was 1,860 psi having a standard deviation of 380 psi. As for the clay brick prisms, the average compression strength was 3,920 psi with a standard deviation of 880 psi. Additionally, following ASTM C 1314-02 (Standard Test Method for Compressive Strength of Masonry Prisms), a height to thickness correction factor was applied to the average masonry prism compressive strength. The concrete prisms were two units tall and had a height to thickness ratio of 4.3, which corresponds with applying a correction factor of 1.15. The clay brick prisms were constructed of five units with a resulting height to thickness ratio of 3.8 corresponding to a correction factor of 1.12. Through the application of these correction factors, the resulting compression strength of the concrete block and clay brick prisms was 2,140 psi and 4,390 psi respectively. These values agree well with the strengths of 2,380 and 3,950 psi predicted using the MSJC code. However, since prism tests were done for research purposes (i.e., prisms had sufficient height-to-thickness ratios to develop an unconfined failure mode representative of the wall specimens), this correction is not considered representative of the material properties of the test specimens used in this experimental program [8]. Thus, the uncorrected compression strengths are used in the remainder of this report. A comparison of the predicted as well as measured masonry compressive strength values is shown in Table 4.1.

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Table 4.1 Masonry Compressive Strength, f’m

Number of Specimens

(each)

Compressive Strength, f’m

(psi)

MSJC (Predicted) N/A 2,380

UMN (Measured/Uncorrected) 16 1,860 (380)

Con

cret

e B

lock

UMN (Measured/Corrected) 16 2,140 (440)

MSJC (Predicted) N/A 3,950

UMN (Measured/Uncorrected) 16 3,920 (880)

Cla

y B

rick

UMN (Measured/Corrected) 16 4,390 (980)

Note: Standard deviations are shown in parenthesis.

4.3.2 Modulus of Elasticity, Em The MSJC code defines the modulus of elasticity as a function of the masonry compressive strength, 900f’m for concrete masonry and 700f’m for clay masonry. Therefore, the predicted modulus of elasticity based on the final concrete prism compressive strength (f’m = 1,860 psi) was 1,674,000 psi. Similarly, considering the final clay prism compressive strength (f’m = 3,920 psi) the predicted modulus of elasticity was 2,744,000 psi. In addition to the modulus of elasticity predicted by MSJC provisions, this property was measured by collecting stress and strain data during compression tests conducted on both concrete and clay masonry prisms. Results from three tests done on each type of masonry were used to determine the chord modulus of elasticity (taken between 5 percent and 33 percent of the maximum compressive strength of the prism). Measured values showed the average modulus of elasticity of the concrete specimens to be 1,262,900 psi while on the clay specimens it was 3,824,000 psi. Table 4.2 is included to allow for comparison of the modulus of elasticity results for both the concrete and clay specimens.

31


Table 4.2 Modulus of Elasticity, Em

Number of Specimens

(each)

Modulus of Elasticity, Em

(psi) MSJC (Predicted) N/A 1,674,000

Con

c.

Blo

ck

UMN (Measured) 3 1,262,900

MSJC (Predicted) N/A 2,744,000

Cla

y B

rick

UMN (Measured) 3 3,824,000

4.3.3 Flexural Tensile Strength Normal to Bed Joints, f’tn Depending upon the direction of flexural tension stress, mortar type, and masonry type the MSJC code provides a means of predicting the modulus of rupture, or flexural bond strength, of masonry assemblages. For this experimental program, with out-of-plane lateral loads being applied to the specimens, the flexural tensile stress acted normal to the bed joints and the units were laid in running bond using a Type S Portland cement-lime mortar for both concrete and clay PTM walls. In regard to masonry type, the MSJC predicted flexural bond strength depends upon whether the units are solid or hollow and not upon whether they are concrete or clay. For this project, the hollow concrete unit assemblages were predicted to have a flexural bond strength of 63 psi, and the clay unit assemblages, which were solid, were projected to have a flexural bond strength of 100 psi. In addition to the predicted values, the flexural bond strength of both concrete block assemblages and clay brick assemblages was tested using a bond wrench frame following ASTM C 1072-00a (Standard Test Method for Measurement of Masonry Flexural Bond Strength). Prisms of both types of masonry were made from three different batches of mortar. Two prisms of each type were made from each batch of mortar. Both types of flexural bond prisms were four units tall, which allowed for three joints to be tested in each prism, resulting in six concrete masonry and six clay masonry joints being tested from each batch of mortar sampled. Therefore, a total of eighteen joints were tested for both types of masonry to evaluate the flexural bond strength of the specimens. During testing, there were two joints in each set of tests, concrete block and clay brick, which results were not obtained due to mechanical difficulties. As with the evaluation of the compression strength data, data outlying plus or minus two standard deviations of the full data set was to be excluded before the average flexural bond strength was calculated. However, no data points fell outside of the range. The average flexural bond strength of the concrete prisms was 120 psi and had a standard deviation of 40 psi, while the clay prisms had an average flexural bond strength of 75 psi and a standard deviation of 30 psi. Again, Table 4.3 directly shows the flexural tensile strength normal to the bed joints, f’tn, values presented above.

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Table 4.3 Modulus of Rupture, f’tn Number of

Specimens (each)

Modulus of Rupture, f’tn

(psi) MSJC (Predicted) N/A 63

Con

c.

Blo

ck

UMN (Measured) 16 120 (40)

MSJC (Predicted) N/A 100

Cla

y B

rick

UMN (Measured) 16 75 (30)

Note: Standard deviations are shown in parenthesis. 4.3.4 Measured Material Properties By knowing the material properties for both the concrete and clay prisms, it was possible to apply this information in the process of predicting the strength the post-tensioned masonry walls, as well as during the analysis of the measured behavior of the specimens. Measured material properties gathered using ASTM testing procedures were considered to be most representative of the masonry specimens used in this experimental program. A summary of the measured values is given in Table 4.4. These values were used for all calculations unless stated otherwise.

Table 4.4 Measured Material Properties

Number of

Specimens (each)

Measured Result (psi)

Compressive Strength, f’m 16 1,860 (380)

Modulus of Elasticity, Em 3 1,262,900

Con

cret

e

Blo

ck

Flexural Tensile Strength, f’tn 16 120 (40)

Compressive Strength, f’m 16 3,920 (880)

Modulus of Elasticity, Em 3 3,824,000

Cla

y

Bric

k

Flexural Tensile Strength, f’tn 16 75 (30)

Note: Standard deviations are shown in parenthesis. 5. SPECIMEN BEHAVIOR This section describes the experimental behavior of twelve post-tensioned masonry wall specimens. There were six concrete block walls, PC1-35-U, PC2-75-U, PC3-150-U, PC4-35-R, PC5-150-R, and PC6-150-R, as well as six clay brick walls, PB1-35-U, PB2-75-U, PB3-150-U,

33


PB4-35-R, PB5-150-R, and PB6-150-R. The testing consisted of applying a monotonically increasing out-of-plane lateral load. 5.1 Visual Observations For all walls, the lateral displacement of the horizontal actuator was increased monotonically as loading was applied. As the lateral loading continued, a horizontal crack opened in a single bed joint at the masonry-mortar interface at or near mid-height, as shown in Figure 5.1. As testing progressed, the crack width at the opening joint became very pronounced and the lateral displacement increased quickly with subsequent loading. Lateral displacements at the end of loading were significant, on the order of 4 to 5% of wall height, as seen in Figure 5.2. Additionally, crushing of the mortar on the compression face of the opening joint was observed after the walls reached maximum moment capacity. The location of the joint that opened for each wall is shown in Table 5.1, as well as the distance from the open joint to the top and bottom pin, respectively.

Table 5.1 Location of Joint Opening

Wall Specimen

Bed Joint Number of Opened Joint (Numbered Sequentially

from Bottom)

Distance of Crack from the Top Pin/Bottom Pin,

Respectively (in) PC1-35-U 7 65.4 / 65.4 PC2-75-U 6 73.4 / 57.4 PC3-150-U 6 73.4 / 57.4 PC4-35-R 5 81.4 / 49.4 PC5-75-R 9 49.4 / 81.4 PC6-150-R 6 73.4 / 57.4 PB1-35-U 19 70.7 / 60.1 PB2-75-U 19 70.7 / 60.1 PB3-150-U 26 52.1 / 78.7 PB4-35-R 28 46.8 / 84.0 PB5-75-R 16 78.7 / 52.1 PB6-150-R 26 52.1 / 78.7

Tendon behavior during loading was dependant upon tendon restraint condition. The unrestrained tendons were seen to move relative to the wall until contact was made with the inside of the restraint, while the restrained tendon movement corresponded with that of the wall except for small movements associated mostly with the tolerance in the restraint devices. Additionally, for the restrained cases, very small amounts of compression of the wooden block inserts (on the order of 1/64" to 1/32") was observed during last stages of testing.

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Figure 5.1 Photograph of Wall PC6-150-R During Testing

Figure 5.2 Profile of PC6-150-R During Testing

5.2 Measured Loads Loads during each test were measured using laboratory fabricated and calibrated load cells that were directly connected to the specimen, as well as load cells that were internal to the horizontal and vertical actuators. The reported results are those obtained from the load cells directly connected to the specimens, while the loads applied by the actuators were used to validate the

35


reported loads. Additionally, the load in the vertical prestressing tendons was recorded through use of four vertical load cells, two on each rod. 5.2.1 Distribution of Lateral Load An important aspect of this experiment was the distribution of lateral load, and how well the lateral loading system simulated uniformly distributed lateral loading. The horizontal load cells and whiffletree system had been utilized in a previous research project at the University of Minnesota ([2], [26]). In the previous research, the lateral loading system provided an acceptable degree of uniformity in lateral loading, and, therefore the system was used in the current project. The ability of the lateral loading system to simulate a uniformly distributed lateral load was again validated through its use in this experimental program. Analysis of the test data demonstrated that the horizontal load cell readings were approximately uniform until the peak capacity of the wall. Following peak capacity, the capacity in each branch of the whiffletree system became independent of one another due to the location of the opening joint. This behavior is shown Figure 5.3 for Wall PC2-75-U.

Figure 5.3 Moment Distributions for Wall PC2-75-U

5.2.2 Moment-Displacement Behavior A plot of the mid-height moment capacity-displacement response of the specimens can be seen in Figure 5.4 and 5.5 for the concrete block and clay brick walls, respectively. It was observed that the behavior of all the specimens prior to cracking was linear-elastic. The post-cracking behavior of the wall specimens differed based on the restraint condition of the tendons. It was also noted that the maximum moment was located at some finite (small) distance from mid-height. To maintain consistency when comparing the specimens, the mid-height moment capacity will be reported unless otherwise noted.

36


The concrete block walls experienced slightly lower moment capacities than the brick walls. However, for the lower levels of effective prestress, the post-peak behavior of the block walls was more desirable (i.e., less steep unloading branches). As shown in Figures 5.3 and 5.4, specimens with restrained tendons displayed more controlled behavior after reaching nominal strength. Additionally, it was observed that higher levels of effective prestress were seen to result in higher capacities.

Figure 5.4 Experimental Moment - Displacement Behavior of CMU Walls

Figure 5.5 Experimental Moment - Displacement Behavior of Clay Brick Walls

37


Joint opening controlled the behavior of ten of the twelve walls, while one wall developed a compression failure mechanism, and one wall exhibited both joint opening and a compression failure mechanism at the instant of maximum moment capacity.

5.2.3 Tendon Stress Histories Analysis of the behavior of the post-tensioning tendons during loading is also necessary to develop a complete characterization of wall response to lateral loads. Tendon stresses increased during lateral loading, and it was determined that the initiation of tendon stress increase corresponded directly to initial cracking of the specimen. The measure tendon stress histories for the concrete block and clay brick specimens are presented in Figures 5.6 and 5.7, respectively. Prior to cracking, the unrestrained tendons, which were placed at wall mid-depth, coincided with the neutral axis, and elongation was not imposed on the tendon. However, the neutral axis shift associated with wall cracking resulted in the tendon being located in the tension region of the wall section. Since the neutral axis continued to shift with increasing lateral displacement, the tendons experienced increases in tension stress with loading. The rate of stress increase in the restrained cases was larger than that of the unrestrained cases. After cracking, the stress in the restrained tendons was larger because the tendons were restricted to the center of the wall, which caused the tendon location to be further from the shifted neutral axis than in the unrestrained tendons.

Figure 5.6 Experimental Tendon Stress Histories for the CMU Walls

38


Figure 5.7 Experimental Tendon Stress Histories for the Clay Brick Walls

5.3 Measured Displacements Numerous Linear Variable Displacement Transducers (LVDTs) were used as the primary means to measure the displacement of the specimens. Seven horizontal LVDTs were attached to the compression side of the wall specimen to generate a displacement profile of the wall during experimental testing. Additionally, horizontal LVDTs were used to measure the displacement of each prestressing rods relative to the wall. Vertical LVDTs were also positioned to the tension and compression face of each wall to measure vertical displacements between gage points along the wall height. These measurements were used to calculate surface strains and neutral axis depth during testing. 5.3.1 Displacement Profiles Using data from the seven horizontal LVDTs positioned along the height of the wall, displacement profiles at nominal capacity were obtained for the concrete block walls as shown in Figure 5.8, while the clay brick wall profiles are depicted in Figure 5.9. It was determined that when cracking controlled the wall behavior, the displacement corresponding to nominal capacity was significantly less than when the walls failed in compression. Additionally, wall specimens with higher levels of effective stress sustained higher displacements.

39


Figure 5.8 Displacement Profiles of the Concrete Block Walls at Nominal Capacity

Figure 5.9 Displacement Profiles of the Clay Brick Walls at Nominal Capacity

40


5.3.2 Tendon Displacement Relative to the Wall Two horizontal LVDTs, one connected to each tendon at mid-height, were used to determine the position of the tendons relative to the wall during the lateral load tests. The movement of the tendon with respect to the wall is plotted in Figure 5.10 and Figure 5.11 for the concrete block walls. As was demonstrated, the movement of the tendon in the unrestrained restraint cases initiated when testing began. This behavior indicated that the tendon moved relative to the centerline of the wall until the tendon come into contact with the inside surface of the restraint. Once contact was made, the tendon moved concurrently with the wall. The movement of the tendon in the restrained walls generally coincided with the movement of the wall specimen. Some differential movement with the restrained tendon originated from two sources. First, the hole for the tendon in the wooden block insert was slightly oversized by 1/16" to allow for the tendons to be installed, which resulted in some displacement until contact with the edge of the hole was made. Additionally, a limited amount of compression of the wooden block insert occurred during testing. However, the majority of the measured relative tendon motion was due to tolerance. Tendon relative displacement measurements were not taken during the testing of Wall PC5-75-R.

Figure 5.10 Tendon Displacement Relative to Wall for the Concrete Block Specimens

41


Figure 5.11 Tendon Displacement Relative to Wall for the Clay Brick Specimens

5.3.3 Axial Strain-Displacement Relations Before the start of each test, ten LVDTs were positioned vertically on the faces of the wall, seven on the tension side and three on the compression side of the wall. Additionally, an extra LVDT was made available and installed on the compression face of an opening joint if it was not monitored by the three original LVDTs on the compression face. The data collected from these instruments was used to determine where cracking and crushing of the joints occurred on the tension and compression faces. Shown in Figure 5.12 and Figure 5.13, tension and compression strains of the opening joint are plotted for the concrete block and clay brick walls, respectively. Initial cracking of the wall specimen was determined to be when the rate of change of the LVDT began to increase. It is noted that the magnitude of the strains are large, due to the extensive level of testing, which demonstrates that crushing did occur on the compression face of the opening joint (i.e., values exceeded 0.0025 for block and 0.0035 for brick). It was not possible to measure the displacements needed to calculate compression strains in the opening joints of all of the walls. But, all of the walls demonstrated evidence of compression failure by crushing and powdering of the mortar in the compression side of the opening joint.

42


Figure 5.12 Tension and Compression Surface Strain for the Concrete Block Walls

Figure 5.13 Tension and Compression Surface Strain for the Clay Brick Walls

5.4 Measured Rotation Two tiltmeters were utilized during experimental testing, one on the top pin and the other on the bottom pin. The pin rotations experienced during testing are shown in Figure 5.14 and 5.15. The rotation history at each end was linear relative to the mid-height displacement. In some

43


instances, there were small rotations seen of the top pin due to the seating of the restraint bracket in the initial wall tests. In these tests, the interaction between the top pin, vertical actuator swivel, and the single pin connection used on each brace caused the top pin to rotate opposite of the direction caused by the lateral loading due to initial misalignment of the top restraint detail. This restraint detail was modified, described in Section 3.7, and the contradictory rotation was not seen in further tests. It is also noted that the top tiltmeter was out of range for Wall PC2-75-U.

Figure 5.14 Rotation-Displacement Relationship for the Concrete Block Walls

Figure 5.15 Rotation-Displacement Relationship for the Clay Brick Walls

44


5.5 Member Stability None of the specimens exhibited failure by member instability, that is, there were no buckling failures. Buckling failures in masonry are characterized by sudden (i.e., nearly instantaneous) losses in capacity with little or no prior material distress except joint opening. Buckling was suppressed in these walls for several reasons. First, there was very little external axial load, approximately 100 lbs, which was maintained constant during the entire test to ensure that the elements of the top pin support remained seated. Additionally, the post-tensioning force in the restrained tendons could not act as a destabilizing axial force because the tendon position was maintained relative to the masonry. Moreover, as the wall deflected laterally, they introduced a horizontal component of force in the tendon that opposed that motion, thus serving as a restraint against the displacements associated with buckling. For the unrestrained tendons, the post-tensioning force behaved similar to the restrained cases after the tendon came into contact with the masonry, which occurred at displacements smaller than 1 inch, shown in Figure 5.10 and 5.11. The descending branch (i.e., negatively-sloped portions) of the moment-displacement response (Figures 5.4 and 5.5) of the walls is associated with crushing of the mortar in the joints that opened, and crushing of the masonry at these joints is responsible for the observed loss of capacity of all of the walls. 6. ASSESSMENT OF WALL BEHAVIOR Through the analysis of the experimental data, wall behavior was defined on the basis of four main events: 1) initial cracking, 2) joint opening, 3) moment strength, and 4) crushing. The method for defining these events is described below. 6.1 Initial Cracking

Initial cracking is the instant at which the first joint to initiate the cracking process develops a tensile strain at the extreme tension face which overcomes the combination of the bond strength between the mortar and the units and the vertical compression stress from post-tensioning, weight, and vertical load. The primary method used to define initial cracking was through analysis of the data from the vertical LVDTs. By examining the surface strains at joints near the mid-height of the walls, the initial cracking point was determined. Surface strains on the tension and compression faces of the walls were computed assuming linear, elastic behavior and uncracked sections. Plots of the wall surface strains were used to identify first cracking to be when the rate of change of strain on the compression face no longer matched that on the tension face, as shown in Figure 6.1 for Wall PC3-150-U. The strain at this instant was approximately equal to the ideal cracking strain computed as follows

Equation 6-1 m

rcr E

f=.ε

45


where

εcr = initial cracking strain (in/in) fr = bond strength (psi) Em = modulus of elasticity (psi).

Figure 6.1 Initial Cracking Determined by Vertical Strains

Additionally, initial cracking was supported by two other methods. First, the same initial cracking point was also observed as the point when the vertical load cells on the tendon began to increase, which is validated in Figure 5.6 and 5.7, due to the shift of the neutral axis that occurred due to cracking. The second method, when the horizontal load cells experienced a change in slope, was also seen to illustrate the same initial cracking point because stiffness of the wall was reduced at this point. The displacements corresponding to the initial cracking, as well as the corresponding experimental cracking moment, are tabulated in Table 6.1. Further analysis of behavior at these points indicates that cracking involved multiple joints. Cracking of each joint was determined by looking at the corresponding measured strains and comparing them to levels of cracking predicted using linear-elastic analysis and measured material properties.

46


Table 6.1 Initial Cracking

Wall Specimen

Displacement Corresponding to

Initial Cracking (in)

Experimental Cracking

Moment (lb-in) PC1-35-U 0.09 5140 PC2-75-U 0.18 8340 PC3-150-U 0.32 9570 PC4-35-R 0.08 6250 PC5-75-R 0.26 11200 PC6-150-R 0.14 14300 PB1-35-U 0.05 6100 PB2-75-U 0.10 8240 PB3-150-U 0.13 9180 PB4-35-R 0.09 6300 PB5-75-R 0.07 9180 PB6-150-R 0.18 14900

6.2 Joint Opening While initial cracking was measured in multiple joints, these crack tips developed into near full-depth joint cracks in only a single joint per wall, and these joints were observed to open continuously through the duration of the test for each specimen. Again, the surface strains from the vertical LVDT data were the primary method for determining when joint opening occurred. These surface strains were computed as follows using Figure 6.2

Equation 6-2 ..LG

tt

∆=γ

Equation 6-3 ..LG

cc

∆=γ

Equation 6-4 q

ct γγφ

−=

Equation 6-5 xcc *φγε +=Equation 6-6 xtt *φγε −=

47


where εc = scaled surface strain on the compression face (in/in) εt = scaled surface strain on the tension face (in/in)

φ = curvature (in) γc = gross strain on the compression face (in/in) γ∆

t = gross strain on the tension face (in/in) c = displacement measured on the compression face (in)

∆t = displacement measured on the tension face (in) G.L. = gage length of the LVDT (in) d = wall width (in) q = distance from the LVDT on the tension face to the LVDT on the compression

face (in) x = distance between the LVDT and the face of the wall specimen (in)

γc

∆

∆

t

q d x x

γ ε

εφ

φ

Strain Distribution

c

t t

d x x q

c

Displacement Distribution

Figure 6.2 Displacement and Strain Distributions

By analyzing surface strains on the tension face of non-opening joints as compared to the opening joint, the peak value of the linear portion following initial cracking was defined as joint opening. This event was defined as joint opening since the tension strains relaxed once a single joint opened, as shown in Figure 6.3 for Wall PB2-75-U.

48


Figure 6.3 Joint Opening Determined by Vertical Strains

The joint opening event occurred simultaneously with a drop in load in the horizontal load cells for eleven of twelve walls. Wall PC6-150-R was the only exception. However, it should be noted that this wall was the only wall that experienced ultimate strength at a point that did not correspond with joint opening. The mid-height displacement values corresponding to joint opening are listed in Table 6.2

Table 6.2 Joint Opening

Wall

Specimen Displacement

Corresponding to Joint Opening (in)

Experimental Moment Capacity Corresponding to Joint Opening (lb-in)

PC1-35-U 0.20 5850 PC2-75-U 0.58 8740 PC3-150-U 0.64 13200 PC4-35-R 0.16 6800 PC5-75-R 0.51 13100 PC6-150-R 0.86 21100 PB1-35-U 0.13 8690 PB2-75-U 0.21 14200 PB3-150-U 0.36 18000 PB4-35-R 0.41 9190 PB5-75-R 0.49 13900 PB6-150-R 0.65 22300

49


6.3 Moment Strength As mentioned previously, moment strength was controlled by two different factors: either joint opening or compression failure mechanism in which the compression strain capacity of the extreme compression fiber was exhausted. Ten of the twelve walls experienced their highest moment strength corresponding to joint opening. Since Wall PC5-75-R was able to re-establish capacity after joint opening, it was considered in both categories. Wall PC6-150-R was the only wall to demonstrate a peak capacity controlled by the compression failure. The mid-height displacements corresponding to moment strength are tabulated in Table 6.3 depending on the failure type.

Table 6.3 Moment Strength

Wall

Specimen Displacement

Corresponding to Joint Opening (in)

Displacement Corresponding to

Compression Failure (in)

Experimental Moment Strength

(lb-in) PC1-35-U 0.15 5980 PC2-75-U 0.36 9470 PC3-150-U 0.52 13900 PC4-35-R 0.14 6800 PC5-75-R 0.50 1.16 12800* PC6-150-R 1.41 22000* PB1-35-U 0.10 8990 PB2-75-U 0.21 14500 PB3-150-U 0.31 18700 PB4-35-R 0.38 9190 PB5-75-R 0.45 13900 PB6-150-R 0.64 22300

*Plastic moment capacity corresponding to a fully developed compression zone 6.4 Crushing The surface strain data obtained from the vertical LVDTs on the compression face of the wall specimen, were used to monitor the strain demands in the mortar joints spanned by the instrumentation. Crushing strains, as defined in the MSJC strength design provisions [20], were considered to be 0.0025 for concrete and 0.0035 for clay. It was determined that crushing strains developed on the compression side of the opening joint, as depicted in Figures 6.4 and 6.5 for the concrete block and the clay brick walls, respectively. Visual crushing of the compression face of the opening joint was present on all walls during the latter half of the experimental testing.

50


Figure 6.4 Crushing Strain of the Concrete Block Walls

Figure 6.5 Crushing Strain of the Clay Brick Walls

51


7. CONCLUSIONS AND RECOMMENDATIONS The objectives of this project were to investigate by experiment the lateral load behavior of slender PTM walls. All of the walls had a slenderness ratio, h/t, equal to between 38 and 40.5, depending on the type of masonry, and the experimental variables included masonry type (concrete block or clay brick), tendon restraint condition (restrained or unrestrained), and magnitude of effective prestress (35, 75 and 150 psi on the net area of masonry). All walls exhibited linear-elastic response to the distributed lateral loading up to initial cracking, followed by nonlinear behavior. The nature of the post-cracking response differed depending on the failure mode of the walls. Joint opening controlled the behavior of ten of the twelve walls (Wall PC1-35-U, Wall PC2-75-U, Wall PC3-150-U, Wall PC4-35-R, Wall PB1-35-U, Wall PB2-75-U, Wall PB3-150-U, Wall PB4-35-R, Wall PB5-75-R, and Wall PB6-150-R), one wall developed a compression failure mechanism (Wall PC6-150-R), and one wall exhibited both joint opening and a compression failure mechanism at the instant of maximum moment capacity (Wall PC5-75-R). All walls underwent large displacements before losing their load carrying capacity, with maximum mid-height displacements being 3 to 5% of total wall height (i.e., pin-to-pin). The nature of wall response to lateral load, as indicated by the shape of the moment-displacement curves, depended upon the magnitude of prestress, the restraint condition of the tendons, and the type of masonry. Generally, walls with larger magnitudes of prestress performed better than did those with smaller prestress magnitudes, and walls with restrained tendons exhibited better response than did those with unrestrained tendons. The block walls experienced slightly lower moment capacities than did the brick walls, but the post-peak behavior of the block walls was more controlled (i.e., less steep unloading branches) at the levels of effective prestress selected for this project. In regard to the restraint conditions, specimens with restrained tendons demonstrated more controlled behavior after reaching nominal strength. Higher levels of effective prestress were seen to result in higher capacities, and more desirable behavior was observed in the restrained walls. The results of this experimental project add to the knowledge base of post-tensioned masonry walls, and further analysis of these results need to be addressed to develop a reliable design procedures for slender walls. Through this experimental program, it is becoming apparent that there is a relationship between wall slenderness influences the potential for mortar crushing, and this issue requires further study. In addition, it appears that restrictions may be necessary in building code provisions to limit 1) very low levels of effective prestress and 2) usage of unrestrained tendons in very slender walls. However, there are large potential gains to be made in terms of cost-effective construction as the design of slender post-tensioned masonry is moved into mainstream practice.

52


Acknowledgements This research was conducted with financial support from the National Concrete Masonry Association (NCMA) Foundation through the Paul and Helen Lenchuk Fellowship program, the Brick Industry Association (BIA), and the International Masonry Institute (IMI). The project was initiated with support from the US National Science Foundation (NSF) through grant CMS-9904110, and the University of Minnesota. Masonry materials were donated by the Anchor Block Company and the Minnesota Brick & Tile Company, and the labor to build the masonry wall panels was donated by mason instructors from the Bricklayers and Allied Craft Workers Local Union 1, Minneapolis, Minnesota. Other materials were donated by the Sika Corporation. This support is most gratefully acknowledged. References

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13. International Masonry Institute, “Masonry Construction Guide: Prestressed Masonry,” Section 8-12, Annapolis, Maryland.

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strength and stability of slender post-tensioned concrete masonry walls under ...€¦ · ·...

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