Shake table studies of energy dissipating segmental bridge column

Column base connections for hollow steel sections: Seismic performance and Strength models

CHAPTER 1

INTRODUCTION1.1 GENERALIn steel structures, column base connections transfer forces from the steel column to the concrete footing. For low-rise structures, such as moment resisting frames with less than three to four stories, or cantilever column systems used in mezzanines, the base connection often features a base plate similar to the one schematically illustrated in Fig. 1. As shown in the figure, these connections include various components (i.e., column, base plate, anchor rods, and concrete/grout foundation) that interact under a variety of loading conditions such as axial tension/compression, flexure, and shear. Informed by various experimental and analytical studies .the Steel Design Guide One), published by AISC, provides guidance for the design of these connections. The method presented in Steel Design Guide One represents one of the most popular approaches to characterizing the strength of base connections worldwide. Recent work, including experimental work, examines the assumptions inherent in the methods provided by Steel Design

Guide One and has determined that, by and large, the design approaches are reliable and conservative if applied to common connection configurations. However, these studies also point to limitations of the experimental/analytical data, as well as their adaptation to design approaches. Specifically, the following issues are noteworthy: The vast majority of experiments feature W-section columns subjected to major axis bending attached to base plates, and none examine the response of hollow steel section (HSS) or box column base plate connections. HSSs are commonly used as corner columns in moment frames or as cantilever column systems for mezzanines and storage racks. HSS columns require the deposition of an out-of-position curved weld at the corner of the section. Moreover, these welds are single sided welds deposited with no access to the inside of the column,with the possibility of incomplete fusion. As the columnis subjected to flexure, these welds interact with the yield line in the base plate, which also forms at the corners (or edge) of the section, possibly compromising the strength and deformation capacity (which is important in a seismic context).

The strength characterization methods presented in prior research (and most of the previous tests) are applicable to a connection configuration where four anchor rods are used (like the one shown in Fig. 1). Often, alternate configurations may be used with eight anchor rods, like the one shown in Fig. 2, to provide additional resistance. The internal stress distributions in these connections are the result of nonlinear and indeterminate interactions between the various components, such as the plate, rods, and grout. Consequently, these different configurations alter the internal stress and force distributions in the rods and the concrete and affect the pattern of development of yield lines in the base plate. With reference to the previous point, the design approaches and strength characterization methods are based on internal stress and force distributions that have been examined only indirectly due to the difficulty of measuring internal stresses in a foundation. More specifically, these methods have been validated through the overall agreement of specimen strength (as obtained from tests) with strength calculated with a method that uses these stress distributions, rather than through direct measurements of these stresses and forces inside the connection The preceding Points are especially relevant if non ductile anchors, such as post installed anchors, are used. In these situations, accurate characterization of internal forces (rather than overall connection strength) is even more critical to prevent sudden failure of the connection. Motivated by the aforementioned issues, this paper presents test results and an analysis of a series of eight tests on HSS column to base plate connections. Fig. 3 and Table 1 illustrate the test setup and the test matrix (discussed in detail subsequently). In reference to these, the test specimens (in the form of cantilever columns) were approximately prototype scale, with variations in plate size and thickness and in anchor rod layout. All specimens were subjected to cyclic lateral loading, and the anchor rods were instrumented with load cells for direct measurement of forces.

Fig 1. Schematic illustration of typical exposed base plate

connection

Fig 2. Hollow steel section column base plate with eight anchor

Rods1.2 OBJECTIVEThe objectives of the test program and subsequent analysis are to characterize the response, i.e., strength, deformation, and hysteretic characteristics, of HSS column base connections, examine the accuracy of existing methods for alternate configurations and relative to directly measured anchor force data, and refine design methods and philosophies for the use of these types of connections in structural systems. The paper begins by providing a brief overview of the design methods and other background regarding the design of column base connections. The testing program is then described. This is followed by an analysis of the test data, as well as a refined method for strength characterization, for an alternate connection configuration, such as that shown in Fig. 2. Conclusions are then presented along with recommendations for design, ongoing work (that focuses on the development of a displacement- based design method that leverages the deformation capacity of these connections), and the limitations of the study . The main objective of this paper is to examine and refine existing strength characterization methods for HSS column bases and those with alternate anchor rod layouts

CHAPTER 2 BRIEF DESCRIPTIONS ON KEY WORDS USED2.1 COLUMN BASE PLATESBase plates as one of the most important elements in structures can influence the total behaviour of structures. Behaviour of base plates as one of the connections that are used in buildings, has its own complexity. The existence of different materials such as steel and concrete, interaction between materials, existence of axial force, shear and moment are the most important problems in analysing these connections. Because of the large number of parameters involved in the behaviour of column base plates, the technical analysis of these connections has always had its special complexities. In addition to the complexities involved in the study of the interaction between steel and concrete, the existence of additional components which are normally added to the system to increase its rigidity makes the study of the system much more complicated.ABC is bridge construction that uses innovative planning, design, materials, and construction methods in a safe and cost-effective manner to reduce the onsite construction time that occurs when building new bridges or replacing and rehabilitating existing bridges. Intrinsic benefits of the ABC approach include the improvements in safety, quality, durability, social costs and environmental impacts. 2.2 SEGMENTAL BRIDGE COLUMNA segmental bridgeis abridgebuilt in short sections called segments, i.e., one piece at a time, as opposed to traditional methods that build a bridge in very large sections. The bridge is made of concretethat is either cast-in-place orprecast concrete.These bridges are very economical for long spans (over 100 meters), especially when access to the construction site is restricted. They are also chosen for their aesthetic appeal. The column constructed using segments is called segmental column.

2.3. ELASTOMERIC BEARING PAD

Elastomeric bridge bearing are a commonly used modern bridge bearing. There are several different similar types of bridge bearings that include neoprene bearing pads, neoprene bridge bearings, laminated elastomeric bearings and seismic isolators which are all generally referred to as bridge bearing pads in the construction industry. They are designed and manufactured based on standards and specifications of such organizations as British Standard, AASHTO, and European Norms En 1337. Internal structure consists of a sandwich of mild steel shims and rubber moulded as one unit. Elastomeric bearing pads compress on vertical load and accommodate horizontal rotation and provide lateral shear movement. Elastomeric bearing pads are the most economical solution used in construction of large span bridges and buildings.2.4 CARBON FIBER REINFORCED POLYMER

Katsumata et al. reported seismic retrofitting of concrete columns with carbon Fiber reinforced composites.The test results of this new technique showed that winding of carbon fibers around concrete columns greatly increased their earthquake-resistant capacity. Fiber reinforced composites, due to their high strength-to-weightand stiffness-to-weight ratios, large deformation capacity, corrosion resistance to environmental degradation, and tailorability, presentan attractive option as an alternative and extremely efficient retrofitting technique in such cases through the use of composite jackets or wraps around a deteriorated column. Another attractive advantage of FRP over steel straps as external reinforcement is its easy handling, thus minimal time and labour are required for installation

2.5 ENGINEERED CEMENTITIOUS COMPOSITES

Engineered cementitious composites (ECC) are a class of high-performance fiber reinforced cement composite with strain hardening and multiple cracking properties. J.L. Pan* F. Yuan tested a number of RC/ECC composite column joints have been tested under reversed cyclic loading to study the effect of substitution of concrete with ECC in the joint zone on the seismic behaviors of composite members. According to the test results substitution of concrete with ECC significantly increase the load capacity and ductility of the beam-column joint specimens, as well as the energy dissipation ability.2.6 SHAKE TABLE

There are several different experimental techniques that can be used to test the response of structures to verify theirseismic performance, one of which is the use of an earthquake shaking table(ashaking table, or simplyshake table). This is a device for shaking structural models or building components with a wide range ofsimulated ground motions, including reproductions of recordedearthquakestime-histories. While modern tables typically consist of a rectangular platform that is driven in up tosix degrees of freedom (DOF)byservo-hydraulicor other types of actuators Test specimens are fixed to the platform and shaken, often to the point of failure. Using video records and data from transducers, it is possible to interpret the dynamic behaviour of the specimen. Earthquake shaking tables are used extensively in seismic research, as they provide the means to excite structures in such a way that they are subjected to conditions representative of true earthquake ground motions.2.7 UNBONDED POST TENSIONING

Unbonded post-tensioned concrete differs from bonded post-tensioning by providing each individual cable permanent freedom of movement relative to the concrete. To achieve this, each individual tendon is coated with grease and covered by a plastic sheathing formed in an extrusion process. The transfer of tension to the concrete is achieved by the steel cable acting against steel anchors embedded in the perimeter of the slab. Studies have proved that one way to reduce permanent lateral displacements in bridge columns is to introduce unbonded post-tensioned tendons. CHAPTER 3STRENGTH CHARACTERIZATION APPROACHES, THE ASSUMPTIONS AND CURRENT DESIGN GUIDELINES

3.1 INTRODUCTION

The main objective of this paper is to examine and refine existing strength characterization methods for HSS column bases and those with alternate anchor rod layouts. To provide context for this, this section presents an overview of these strength characterization approaches, the assumptions that they are based on, and current design guidelines. As discussed previously, the response of base plates is controlled by nonlinear interactions between the plate, rods, and supporting concrete. Explicitly incorporating these effects into strength models is cumbersome and unsuitable for application within a design setting (e.g., Ermopoulos and Stamatopoulos 1996). As a result, the prevalent approaches to designing base connections rely on a predetermined form of stress distribution (or stress block), such as linear (triangular) or constant (rectangular) distribution on the bearing side of the connection, and then determine forces in the various components by establishing equilibrium on the entire connection; these are by far the most popular types of methods in use today; in fact, they (specifically those in Steel Design Guide One) are the sole approach used in the design of base plate connections in the United States. As per the approach presented in the Steel Design Guide One,the base connection resists the applied axial compression and moment through the mechanism illustrated in Fig. 4. If the moment is low relative to the axial load, then the load combination may be resisted exclusively through the development of the bearing stresses (i.e., the plate does not tend to uplift)the so-called low eccentricity conditionFig. 4(a). On the other hand, if the moment is large relative to the axial load, then tension must develop in the rods to prevent upliftthe high-eccentricity conditionFig. 4(b). The two conditions are separated by a critical value of load eccentricity ecrit M=P, which may be calculated using the following

formula given in Steel Design Guide One: The preceding equation assumes the development of a rectangular stress block (with a constant magnitude = fDG1max ) on the bearing side, i.e., the compression side of the connection. The symbols P, B, and N represent respectively the applied axial load, length, and width of base plate. The stress magnitude fDG1 max is the lesser of the bearing strength of the concrete or the grout (if used). In general, fDG1 max may be determined using Eq. (2) as follows:

The symbols in Eq. (2) are defined in the notation list. The term in the equation corresponding to the ratio of areas A2/A1 incorporates the effect of confinement on the bearing stresses. For the low-eccentricity condition, i.e., e < ecrit, the only possible limit state is the flexural yielding of the base plate on the bearing side of

the connection due to the upward bearing stresses. The moment in the base plate may be calculated in two steps, which are outlined in Steel Design Guide One. Briefly, these are, first, to consider the overall force and moment equilibrium of the connection to establish the bearing stress f, which (for e < ecrit) is bounded by fDG1max , and the stress block length YDG1, and, second, to calculate the plate

bending moment based on the assumption of a cantilever bending.of the flap that extends outward from the compression flange of the column, i.e., assuming that the yield failure is parallel to the column compression flange. On the other hand, if the anchor rods are engaged (almost always the case in seismic loading), then the high-eccentricity condition is active, i.e., e ecrit. In this situation, the forces in the

anchor rods TDG1 and the bearing length YDG1 may be calculated using Eqs. (3) and (4) in what follows, also presented in Steel Design Guide One: The preceding equations are derived by establishing vertical and rotational equilibrium on the connection. The symbols used in the equations are defined in the notation list at the end of this paper. In the context of this study, it is useful to note two points. First, TDG1 represents the total tension force in all anchor rods on the tension side; if only two are present, then the sharing of load between them

may be assumed to be equal, not necessarily the case if alternate rod layouts (e.g., three in a row) are used. Second, the methods outlined in Steel Design Guide One do not address the effect of an additional set of anchor rods [e.g., as shown in Fig. 2 or by the dashed lines in Fig. 4(b)]. As a result, the force TDG1 is an estimate of Touter shown in the figure. Note that the addition of the inner rods introduces an indeterminacy in the system such that the internal forces may no longer be inferred based only on the force and moment equilibrium of the connection. As discussed subsequently, the present study addresses this issue. In any case, once TDG1 and YDG1 are determined as per the foregoing process, the bending moments in the plate may bedetermined assuming an upward cantilever bending of the flap on the bearing side of the connection or the downward forces in the anchor rods on the tension side. These may then be used to appropriately detail the plate (i.e., thickness) or the rods (size, grade, and embedment, based on consideration of several failure modes; e.g., Cimellaro and Reinhorn 2011). Note that (1) this method has not been developed for HSS members or connections with alternate rod layouts; in fact, consideration of additional rows of rods requires fundamental changes to the method to address the indeterminacy produced by the additional rods; (2) even for connections similar to those implied in the design methods, i.e., those with two rows of rods, the distribution of forces among these rods is not addressed by the current methods; and (3) the internal stress and force distributions are not based on independent measurements of rod force but rather the overall connection response. The following section describes the testing program conducted to respond to these issues.

CHAPTER-4TEST SETUP, INSTRUMENTATION, ANCILLARY TESTING, AND

TEST MATRIX4.1 TEST SETUP AND INSTRUMENTATIONFigs. 3(a and b) (introduced previously) show schematic illustrations of the test setup (in elevation and plan view, respectively), whereas Fig. 3(c) shows a photograph of a specimen being tested [with reference to Table 1, i.e., the test matrix, the specimen shown in Fig. 3(c) corresponds to Test 1]. As shown in the figure, salient aspects of the test setup are as follows:

1. All specimens were cantilever columns of length 3,200 mm with cyclic lateral loading applied at the top. The length corresponds to a typical cantilever column system (such as a mezzanine), in which flexure dominates response. No axial load was applied;

2. All welds were toughness-rated complete joint penetration (CJP) welds deposited from the outside of the column;

3. The base of the column was attached to a reusable concrete footing supported by spacer beams, also shown in Fig. 3(a). The spacer beams allowed access to the underside of the concrete footing for the installation of load cells, which were installed at the lower end of all anchor rods between a nut washer assembly and the underside of the concrete footing. In this way, the tension in the rods exerted compression on the load cells. The top end of the anchor rods featured a similar nutwasher arrangement, such that the anchor rods could only

be engaged in tension during uplift of the plate;

4. Fig. 3(b) shows a plan view of the test setup. As shown in this view, the footing (designed to be reusable) had 24 holes (lined with PVC pipe) through the depth to allow for passage of anchor rods. The 24 holes were designed to accommodatethe three different base plate footprints and corresponding rod layouts;

5. Major instrumentation included load and deformation at the top of the column, i.e., at the actuator location; tension forces in all the anchor rods, measured by the load cells described previously; various displacement transducers to measure the uplift/rocking of the entire foundation; and strain gages on the top surface of the base plate to monitor local strains. The base plates were not entirely flat at the time of specimen installation, presumably because of warping induced by the column

welding process. To ensure contact between the base plate and concrete (and to restore vertical alignment of the column), the nuts on the anchor rods were tightened, inducing a prestress in each of the anchor rods; this prestress ranged from 49 kN (for the thinner plate, i.e., 19 mm) to roughly 2530 kN for the thicker (i.e., 31.75 mm) plate. The results of the experiments are interpreted (in a subsequent section) with appropriate consideration of this prestress. In heavier columns (for multistory frames), column alignment and base plate contact are often ensured using shim stacks or leveling nuts followed by grouting (e.g., Gomez et al.2010). However, for low-rise frames, especially featuring postinstalled anchors, this type of leveling/grouting procedure is unfeasible. As a result, the prestressing method used in this study is similar to construction practice (where the column is first plumbed and the nuts are then differentially turned without disturbing the

plumbing).

4.2 ANCILLARY TESTING AND MATERIAL PROPERTIESTable 1 summarizes the material properties of the various materials in the connection. The 28-day strength of the concrete in the footing (which was common for all the tests) was measured to be f0c 28.1 MPa based on standard cylinder tests. All rods were 19 mm diameter ASTM F1554 Grade 105 (i.e., Fu 723 MPa); by design, they did not yield in any of the experiments. The steel properties of the column and base plates were obtained from mill certificates. The columns were nominally A500 Grade B (Fy 317 MPa), whereas the plates were nominally A36 (i.e., Fy 248 MPa). The table shows that the column and plate yield (and

ultimate strengths) are greater (by approximately 35%) as compared to the specified values. The difference between the specified and expected strengths is common (Liu et al. 2005) and typically incorporated into capacity design by way of the Ry or Rt factors (AISC 2010). The columns were sized to remain elastic even as the connection developed its full capacity. As shown in the table,the flexural capacity of the column (based on the measured yield strength) was greater compared to the maximum moment Mtest max for all the tests, suggesting that there was no yielding in the column and that the measured peak moments correspond to the connection strength. All subsequent analyses (of connection strength/response) are based on the measured, rather than the specified, material

propert4.3 TEST MATRIX AND LOADINGTable 1 also summarizes the test matrix. The table shows that the key variables were the plate dimensions (i.e., footprint and thickness) and the anchor rod layouts, i.e., four or eight anchor rods. Tests 14 (with 305 mm square base plates) featured the four rod pattern, whereas Test 58 (with 457 or 610 mm square base plates) featured the eight-rod pattern. Thus, three plate sizes (i.e., footprints) were investigated, along with two plate thicknesses, i.e., 19 and 31.75 mm. The anchor rods were placed in a square pattern with an edge distance of g 38 mm for all the experiments. Table 1 shows that all the experiments featured square HSS columns. The dimensions and layouts of the test specimens are representative of prototypical configurations commonly used in current construction practice. For each of the parameter sets, two replicate tests were conducted. All the specimens were subjected to the ATC-SAC loading protocol (Krawinkler et al. 2000) to represent deformation histories consistent with seismic demands in moment frame buildings. The deformation history applied to the specimens (at the top) as per this protocol is expressed in terms of column drift ratio defined as the lateral displacement of the column at the application of the lateral load divided by the distance between the load and the top of the base plate (3,200 mm for all tests). In all except three of the eight tests (i.e., Tests 5, 7, and 8) failure did not occur during

the loading history. Consequently, a monotonic excursion was appended to the end of the history, in which the magnitude of the excursion was controlled by the stroke limit of the actuator (consistent with a column drift of approximately 13%).

Test setup: (a) schematic elevation; (b) plan view of concrete footing;

(c) test in progress (Test 1 shown at approximately 12% drift) CHAPTER 5

TEST RESULTS5.1 GENERALFigs. 5(ah) indicate the moment-rotation response for all the experiments.

Figs. 6(a and b) show photographs of two specimens at the conclusion of testing. The moment-rotation response shown in Fig. 5 plots the base moment (calculated asMbase Vtop Lcolumn) against the base rotation base, which is determined using Eq. (5) This provides a direct assessment of the deformation response column. The symbols in the equation are defined in the notation list at the end of the paper. A qualitative assessment of the experimental response for the various tests is now presented to facilitate the interpretation of quantitative data, which is discussed subsequently. All the tests share some features of the qualitative response. These

are first described, and then responses peculiar to specific specimens are presented. For all specimens, the initial elastic response was observed until a base rotation of approximately 0.0075 0.01 rad (0.751%); this corresponds to a column drift of approximately 11.5% [which is consistent with the expected yield drift

in moment frame systems (Krawinkler et al. 2000)]. This was followed by yielding of the base plate on the compression side (due to the upward bearing stress) and then yielding of the base plate on the tension side due to the tension in the anchor rods. The reversed cyclic yielding of the base plate produced a fairly stable hysteretic response, as shown in Figs. 5(ah). A slight pinching was also observed

for the higher amplitude cycles (more than 23% drift) as the base plate (which at this point in the loading had permanent deformations) engaged and disengaged (i.e., contacted and separated from) from the supporting concrete and the anchor rods. Note that in some tests, this so-called pinching plateau is observed at nonzero loads due to the differential prestress in the rods. This hysteretic response continued in a stable cyclic manner until either the test was concluded (due to actuator limitations) or fracture of the weld between the column and the base plate was observed (in Tests 5, 7, and 8). Fig. 6 photographically shows typical posttest deformations for two of these tests, i.e., Tests 2 and 5. Tests 1, 3, and 4, which feature the four-rod layout, show qualitatively similar deformations as Test 2. Tests 68 with the eight-rod layout, show a deformation pattern similar to that of Test 5, with the exception that Test 6 did not exhibit fracture, whereas Tests 5, 7, and 8 showed weld fracture. Other than plate yielding (which is clearly visible for the specimens shown in Fig. 6), no visible distress was observed in any of the specimens until fracture (or conclusion of the test). In the early stages, plate yielding was evidenced by flaking of the paint in the region of the yield lines, whereas in the later stages of loading (drifts greater than 34%) it was observable in the form of permanent deformations in the plate. On the bearing side, slight (and

highly localized) crushing of the concrete was observed under the corners of the compression side of the base plate. The anchor rods did not yield in any of the tests. 5.2 THE BASE PLATEThe main qualitative difference between the response of the test was in the mode of deformation of the base plate. In this regard, the response may be grouped into two categories corresponding to thetwo different types of anchor rod layouts. For Tests 14, the yield lines in the base plate on the tension side of the connection were inclined with respect to the loading direction because only two anchor rods were present in the corners. This is indicated schematically in Fig. 6(a). This is similar to the type of response observed in other testing programs (e.g., Gomez et al. 2010). Although Steel Design Guide One does not explicitly address this type of response (and methods for characterizing associated strengths), it is addressed in research by Mazzei (2012) and AISCs Steel Design Guide Ten (Fisher andWest 2003) in the context of the construction of bracing of low-rise frames. Plate yielding in Tests 58 was constrained by the presence of additional anchor rods. Consequently, yield lines formed parallel to the HSS column edges, and the mode

of plate bending was different compared to the tests with the fourrod layout. This is indicated schematically in Fig. 6(b). Also, as shown in this figure, the presence of additional rods also constrains the yielding of the base plate. Presumably, this constrained mode of deformation led to fracture (in the weld connecting the HSS column to the base plate) in three of the four tests featuring the eight-rod (and capacity) of the connection, discounting the flexibility of thelayout. In these tests, sudden fracture (not preceded by any observable ductile tearing) propagated through the tension flange (or edge of the HSS) and nearly all the way through the webs (i.e., perpendicular edges of the HSS). Fig. 6(b) indicates this fracture for Test 5. Table 1, introduced previously, indicates the deformations corresponding to fracture for Tests 5, 7, and 8.Table 1 (introduced previously) summarizes key quantitative data measured from the tests. This includes the maximum moment sustained by a connection during a test (i.e., Mmax test ), the peak base rotation max base, and the maximum drift sustained by a column, i.e., max. In addition, the table also shows a comparison of the maximum base moment observed in the experiment Mtest max to the estimates of connection strength. Two such estimates are used in the table. One of these, denoted byMDG1, is the estimated moment based on Steel Design Guide One. This estimated base moment corresponds to the initiation of flexural yielding on the tension side of the base plate. For Tests 14, this is based on the assumption of inclined yield lines as defined by Mazzei (2012) and the measured material properties summarized in Table 1. Note that the method outlined in the Steel Design Guide One also considers the limit state of plate yielding on the compression side of the connection. However, as demonstrated by Gomez et al. (2010), this limit state does not affect the overall connection strength, which requires the mobilization of other yielding modes (such as yielding of the base plate on the tension side or yielding of the anchor rods). Moreover, Steel Design Guide One does not address three rows of anchor rods (such as are present in the eight-rod layout, i.e., Tests 58). Thus, for Tests 58, MDG1 is calculated by disregarding the middle row of the anchor rods to examine the efficacy of the Steel Design Guide One in characterizing the connection strength under such conditions.

Three observations may be made based on Table 1 and the oment-rotation curves shown in Fig. 55.3 OBSERVATIONS Three observations may be made based on Table 1 and the moment-rotation curves shown in Fig. 5 .first, In all the experiments, an excellent deformation capacity and stable hysteretic response were observed. As shown in Table 1, the maximum observed rotations were in the range of 0.057 rad (Test 7) to 0.13 rad (Test 1). The corresponding column drifts were in the range of 0.070.16 rad (i.e., 16% drift). This observation is even more noteworthy when viewed in the context of prequalification standards for beamcolumn connections (e.g., FEMA 2000), which require a stable response up to drift angles of 0.06, i.e., 6%. On the other hand, it is not entirely unexpected, given that previous tests on exposed column bases (Gomez et al. 2010) demonstrated a similar deformation response. Seond , The rotation capacity for specimens with an eight-rod layout was lower compared with those with a four-rod layout. An inspection of Fig. 6(b) suggests the underlying cause: the additional anchor rods constrain the yielding of the base plate, thereby promoting the initiation of fracture. Third , The ratio Mtest max/MDG1 (based on measured, rather than specified, material properties) are in a range of 1.24 (Test 8) to 2.03 (Test 3). This indicates that the method presented in Steel Design Guide One may be conservative. This observation is also consistent with previous research (Gomez et al. 2010) and may be attributed to four factors: (1) the method uses the yield stress Fy of the base plate and disregards strain hardening,

which is evidently mobilized in the base plate given the extent of inelastic deformation; note that the measured Fu (Table 1) is in the range of 480520 MPa, roughly 50% higher than Fy used for the calculation of MDG1; (2) the yield lines in the plate are subjected to extremely large strains a high deformations; these strains may generate so-called true stresses that are even higher Fu than (since Fu corresponds to necking and localization of a tension coupon; such necking is absent during plate bending); (3) as the plate is lifted up, membrane action is mobilized, further increasing the load carrying capacity; finally, (4) referring to the prior discussion, the rods carried some prestress at the beginning of the test; this elevates the base moment required for lift-off of the plate and subsequent yielding of the base plate. For Tests 58, the Steel Design Guide One is inherently conservative (as are the previously considered factors) because it disregards the effect of the third (middle) row of anchor rods. However, it is interesting to note that the degree of conservatism for Tests 14 is larger (average Mtest max=MDG1 1.83) compared to that for Tests 58 (averageMtest max=MDG1 1.53). This may be attributed to fracture in three of these (latter) tests, which diminished the deformation capacity relative to Tests 14, thereby reducing the influence of material hardening and the associated increase in the base moment. It appears that this effect offsets the additional strength provided by the third row of anchor rods.

A large difference between test and predicted moment capacities has been observed in the context of base connections (Gomez et al. 2010) and in the context of other connections (end plate moment connections) that have similar plate deformation modes producing high strain hardening and membrane action (Adegoke and Kemp 2003). In general, when a connection is designed to remain elastic, this overstrength is not a problem (although it does indicate that plates may be somewhat oversized). However, it may become a relevant issue if either the connections are designed as a ductile (or fuse) element or the anchors are designed to develop plate yielding. In these cases, more accurate methods to characterize plate strength must be developed. Given that the strength is controlled by strain hardening and membrane action, this type of strength estimate will likely be sensitive to the expected level of deformation.5.4 THE ANCHOR RODSAs discussed previously, all the anchor rods were instrumented with load cells for direct measurement of tension forces. Figs. 7(ac) examine these measured forces (Ttest outer) relative to those estimated as per Steel Design Guide

One (TDG1). More specifically, the figures plot the evolution of the test to predicted ratio Ttest outer=TDG1 versus the deformation. Before discussing the trendsobserved in the figures, it would be useful to note a few points about the figures themselves. First, the Ttest outer values reflect the sum of all the rods in the outer row of anchor rods minus the prestress recorded at the beginning of the test, thereby providing a basis for consistent comparison with corresponding estimates

TDG1, calculated as per Eq. (4). Second, the values are plotted against the absolute value of the peak base rotation (i.e., jbasej) for each positive and negative cycle of loading. Third, since the loading is reversed-cyclic, the Ttest outer values reflect the forces in the rods that are in tension for a particular cycle; in this way, data from both positive and negative cycles may be plotted conveniently. Finally,

since the objective of Fig. 7 is to assess the efficacy of the Steel Design Guide One, only the forces in the outer rows of rods are shown. With reference to Figs. 7(ac), three points may be made: 1. For all tests, initially (i.e., for drifts less than 0.25%) the testpredicted ratios oscillate as the rods begin to engage; note that while the values oscillate, the magnitude of the forces in the rods is fairly low (i.e., less than 510% of the value at peak moment). However, as loading progresses (deformation increases), the test-predicted ratios stabilize. 2. For Tests 14, the values approach a value of 1.0, indicating that Steel Design Guide One provides a fairly accurate estimate of the rod forces. For the range of plate thickness investigated in this study, this is consistent with similar observations from finite element simulations conducted previously by the authors (Kanvinde et al. 2013). This also suggests that for the bolt edge distances and plate thicknesses used in this study, prying forces (induced by the corners of the plate bearing on the concrete) are not a significant factor. 3. For Test 56, the test-predicted ratios approach values lower than 1.0. This is not unexpected since the estimates do not consider the contribution of the middle row of anchor rods. However, it is interesting to note that the values are not significantly lower than 1.0 but in the range of 0.81.0, suggesting that Steel Design Guide One may be used conservatively to size these anchors.

5.5 RESULT SUMMARY

In summary,the method presented in Steel Design Guide One provides reasonably accurate estimates of the total force in the outer row of the anchor rods. It also provides conservative estimates of the connection strength. However, in the context of the connections with an eight-rod layout, it has some limitations. First, since it disregards the inner row of rods [Fig. 4(b)], it cannot be used to determine forces in them or to design them. Second, the method calculates the forces in a row of rods, rather than in individual rods. For a four-rod layout, only two rods are present in each row; the distribution of the force between these is trivial (i.e., the force distributes equally). Fig. 7. Comparison of estimated rod forces (as per Steel Design

Guide One) with measured forces for (a) Tests 1, 2; (b) Tests 3, 4;

(c) Tests 5, 6, 7, 8However, as discussed in the next section, the distribution is more complex for the situation where three rods are present in a row. Third, though the method provides reasonable estimates of overall strength and anchor rod forces (in the outer row) for Tests 58 in this study, it cannot be generalized to other plate sizes or thicknesses since it disregards the inner row of rods. In fact, if the contribution of the inner row is disregarded in the design method, it is difficult to justify the use of these rods, which entail additional cost. The next section proposes an approach to characterizing connection responses in these situations.

CHAPTER 6PROPOSED METHOD FOR CHARACTERIZING RESPONSE OF CONNECTIONS WITH EIGHT-ROD LAYOUTAs discussed previously, the response of the base connections is the result of nonlinear interactions whose explicit characterization is challenging even for the four-rod layout. Thus, previous approaches (including Steel Design Guide One) rely on assumed stress distributions under the plate to facilitate design. The eight-rod configuration is even more challenging to characterize, primarily because of the increased degree of static indeterminacy within the connection. The method presented in this section has the following features: (1) it does not characterize the nonlinear interactions (discussed previously) in an explicit way but rather seeks to

provide a straightforward method that may be applied in a practical setting, and (2) it augments the Steel Design Guide One method with a relatively simple extension, such that it degenerates to the Steel Design Guide One method if the middle row of anchor rods is absent. Referring to Fig. 4 shown previously, the applied force and

moment may be resisted through the development of a bearing stress block on the compression side and forces in the anchor rods on the tension side of the connection. For the low-eccentricity condition, i.e., when there is no plate uplift, the four- and eight-rod configurations may be assumed to respond in an identical manner, i.e., the eccentric force is resisted entirely by the development of the stress block in the concrete. However, when e ecrit [Eq. (2)], the anchor rods are engaged. As shown in Fig. 4(b), the following relationships may be established based on equilibrium of the entire connection:

Eq. (6) establishes the vertical force equilibrium, whereas Eq. (7) establishes the moment equilibrium. The two preceding equations contain three unknowns (i.e., Tnew outer, Tnew inner , and Ynew) and thus cannot be solved unless an additional condition is introduced. It is proposed that the tension in the inner rods should be considered as being directly proportional to the tension in the outer

rods, i.e., Tnew inner k Tnew outer . This assumption is based on a general

kinematic consideration by which the deformations in the two rows of rods are constrained by the plate. However, it is also recognized that the plate is flexible and that the plate or rods will likely yield as loading progresses, such that this direct proportion may no longer be valid. However, an explicit consideration of these effects will necessitate not only further complexity (an obstacle to implementation) but, perhaps more importantly, the introduction of additional parameters that cannot be robustly calibrated with the available data. For the tests in this study, P 0; substituting this into Eqs. (6) and (7) and solving simultaneously results in Once Tnew outer is determined, the force in the inner row of rods may be determined as Tnew inner k Tnew outer. Note that for this study, direct measurements of Tnew outer and Tnew inner are available. Based on these measurements, a value of k is determined (through trial and error) to achieve the best overall agreement with force measurements from all tests. This optimal value is determined to be k 0.5. Figs. 8(a and b) present the evolution of force in the anchor rods (in Tests 58) in a manner similar to Figs. 7(ac). Unlike Figs. 7(ac), Figs. 8(a and b) illustrate the results of the new method and, thus, also include estimated forces for the inner row of anchor rods. With reference to Figs. 8(a and b) (and Table 1 introduced previously) some observations can now be made: 1. For all four tests, the test-predicted ratios of the forces in the outer rods are either close to 1.0 or slightly lower, indicating that the method is reasonable (and perhaps slightly conservative) for estimating the forces in these rods. For all four of these tests the estimated forces in the outer row of rods is between 5 and 10% lower than the corresponding estimates from the current, i.e., Steel Design Guide One, method because the contribution of the inner row is considered. 2. The test-predicted ratios for the inner rods evolve from a value of zero and approach (or even slightly exceed) unity as the test progresses. This is because initially, only the outer row of rods is engaged (such that no forces are measured in the inner row). Upon subsequent loading, and deformation of the plate, the inner rods begin to develop force. 3. The preceding Point 2 suggests that the relationship between the forces in the inner and outer rows is nonlinear, such that the inner rods carry little (0) load at the beginning of the test, but as deformations increase, the accuracy of the method (which assumes a constant ratio between the forces in the inner and outer rows) increases. It is also interesting to note that the test predicted ratios for all the rods are lower for Tests 7 and 8 compared to Tests 5 and 6. Note that Tests 7 and 8 have a larger and thicker base plate (610 610 31.75 mm) and a larger column (HSS 254 254 12.7) compared to Test 5 and 6. This underscores that the rod forces (and the stress distributions that produce them) are sensitive to the relative

flexibility of the various interacting components such as the base plate and supporting concrete. 4. As shown by Fig. 8, the test-predicted ratios for the inner rods approach unity only as the deformations become fairly large (i.e., 34% drift). Thus, it may be argued that their influence can be neglected for lower drifts that are more consistent with design-level deformations. Nonetheless, the use of the refined

method is still desirable for two reasons. First, for the outer row of rods, the refined method (with k 0.5) provides less conservative and more accurate estimates of forces (relative to the Steel Design Guide One method) even for smaller deformations (compare test-predicted ratios for the outer rods in Figs. 7 and 8). Second, and perhaps more importantly, the proposed method provides a way to design the inner row, whereas the Steel Design Guide One method provides no such method. In this context, too, the validity of k 0.5 may be questioned, given that this characterizes rod forces well only at large displacements. It is important to consider the alternatives, which include the use of a nonlinear evolution relationship for k, which would be cumbersome to generalize and apply in a design setting, or the use of k < 0.5, which would be nonconservative if deformations greater than 34% drift were indeed encountered. Thus, the use of the refined approach (with k 0.5) appears to have little downside while providing a rational approach to designing the connection, especially the inner row of rods. 5. As shown by Table 1 (the last column, i.e., Mtest max=Mnew), the test-predicted ratios for the new method are slightly lower (i.e., less conservative) than those determined as per the Steel Design Guide One method. The moment Mnew was calculated in a manner identical to MDG1, with the exception that Tnew outer Fig. 8. Comparison of estimated rod forces (as per proposed method)

with measured forces for (a) Tests 5, 6; (b) Tests 7, 8were used instead of TDG1 to determine the magnitude of plate bending. As shown by Table 1 (and the prior discussion regarding Mtest max=MDG1), the average test-predicted ratio for the four-rod tests from the Steel Design Guide One method was higher (i.e., 1.83) than the corresponding test-predicted ratio when applied to the eight-rod tests (1.53). Thus, the premise of the proposed method that increases the estimates of base moment capacity (thereby resulting in a lower average Mtest max=Mnew 1.37) may be questioned because it suggests that the new method predicts a greater strength for the subgroup of tests (Tests 58), which

on average have a smaller test-predicted ratio as per the Steel Design Guide One method. Two points may be made in response to this concern. First, the measured Mtest max values for both the fourrod and the eight-rod tests are obtained at base rotations in the range of 0.0580.13 (drifts of 0.070.166). These are well in excess of any expected rotations and, thus, are not representative of the strengths that may be expected in design-level scenarios. Within this range, the higher test-predicted ratios of the four-rod specimens are an artifact of their greater deformation capacity relative to the eight-rod specimens rather than a consequence of the internal stress/force distribution or the mechanics. Second, the measurements of anchor rod forces indicate that the proposed method provides a better indication of the internal stress distribution. From a design standpoint, where the deformation magnitudes observed in the tests are not expected, it is prudent to use a method that reflects

the internal mechanics of the connection. Consistency with internal mechanics is even more important for the design of nonductile anchors since they are sensitive to force estimates in them, not to the overall connection strength. Thus, the use of the Steel Design GuideOne method for the eight-rod configuration has two potential problems. First, it has the potential of mischaracterizing the internal

stresses, especially the force in the inner row of rods. Second, it is influenced by the observation of higher overall connection strengths in the four-rod specimens, which in turn are only an artefact of increased deformation capacity. As an alternative argument, consider a hypothetical situation where each of the three rows of rods contains two rods (i.e., a six-rod configuration). One may imagine that this has a deformation capacity similar to the four-rod configuration since the plate will deform with less constraint and with the formation of inclined yield lines. Given

the similar deformation capacity, one may further anticipate that the additional (inner) row of rods in the six-rod configuration will result in additional strength over the four-rod configuration. This hypothetical, but consistent, comparison of the three-row versus two-row configuration (with two rods in each row) supports the use of the proposed method, which reflects the internal stress distribution in an improved way. The distribution of forces among the three rods in the outer row

also bears scrutiny. This is because when only two rods are present per row, the distribution of forces among them is trivial (i.e., forces are shared equally), an assumption that is not self-evident when three or more rods are present in a row. Fig. 9 plots the ratio of the load carried by the middle rod to the total anchor rod forces in all three rods in the outer row, i.e., Tmiddle=Ttest outer. The figure is similar to Figs. 7 and 8 such that the values are plotted against the absolute value of the peak base rotation (i.e., jbasej) for each positive and negative cycle of loading. Moreover, the forces plotted in Fig. 9 also do not include the prestress in the rods present at the beginning of the test. With reference to the figure, the following observations may be made: 1. At the beginning of the test, the Tmiddle=Ttest outer ratio is in the range 0.30.45 (suggesting that the load sharing is approximately equal among the three rods). However, as the deformations increase, the middle rod carries a progressively larger fraction of the load, reaching a peak value of roughly two-thirds (0.65) the total load in the three rods (for Test 5; the ratios are slightly lower for the other three tests). This is not surprising, given that the upward force is introduced into the base plate by the tension flange of the HSS located in the central portion of the plate width (i.e., near the middle rod rather than the corner rods). This may be alternatively explained based on plate-bending solutions (Timoshenko and Woinowsky-Krieger 1959) or by invoking a so-called tributary-width argument according to which the middle rod carries a larger share of the load. In any case, it is clear that designing the rods to carry equal forces is not a satisfactory approach. 2. At very large rotations (i.e., jbasej>0.050.06 rad), the ratio again drops to roughly 0.5 (i.e., the middle rod is still carrying half the total load). This may be attributed to a redistribution of stresses as yield lines form in the plate or to perhaps small prying forces in the corner rods.

Fig. 9. Distribution of tension forces among three outer rodsCHAPTER 7

SUMMARY, DESIGN CONSIDERATIONS, AND LIMITATIONSThis paper presents findings from eight tests on exposed column base connections to examine their seismic response and the efficacy of current design methods in ensuring safety and economy. With respect to previous work in the area, the main distinguishing features of this study were an examination of HSS rather than W-section columns, an examination of alternate anchor rod layouts featuring eight rods (i.e., three rows of rods), which is not addressed by current design methods or previous research, and a direct measurement of anchor rod forces that provide refined insights into the internal force distribution within the connection for a critical evaluation of current design methods.All eight test specimens showed excellent deformation capacity, such that column drift ratios in the range of 716% were achieved, accompanied by a stable hysteretic response and only modest degradation in strength. The primary deformation mode was inelastic bending of the base plate, with minor concrete crushing at the extremities of the plate. This deformation mode was exceptionally ductile, resulting in base rotation capacities (excluding column deformations) in a range of 0.0570.13 rad. In fact, five of the eight tests were terminated without failure (owing to limitations of test equipment). Fracture was observed in three of the tests (with the eight-rod layout). The fracture initiated at the tension flange weld between the HSS column and the base plate and was sudden with no prior observation of initiation. Presumably, the more constrained form of plate yielding in the eight-rod tests promoted fracture in these tests.

The currently used approach for designing exposed column base connections (e.g., the one in Steel Design Guide One) was evaluated against the test data. The approach was directly applicable to Tests 14 (which had the four-rod configuration, i.e., with two rows of anchor rods). For these, it was determined that the current approach was able to characterize the forces in the anchor rods with

good accuracy (suggesting that it characterizes internal force/stress distributions in connections in a reasonable manner). Moreover, with regard to connection strength (i.e., maximum moment capacity), the approach was significantly conservative; this is attributed to prestress in the rods and hardening in the plate material, which was mobilized by the high degree of inelastic deformation. This approach was also applied to Tests 58, disregarding the inner row of anchor rods, to examine its efficacy in these situations. The relatively similar degrees of conservatism for the two sets of tests (i.e., with four-rod and eight-rod configurations) also suggests that perhaps the inner row of rods does not provide significantly improved strength but may reduce deformation capacity. However, an eight-rod layout may be required to provide adequate strength if the connection is also designed for out-of-plane bending.To address the limitations of Steel Design Guide One in these (eight-rod) situations, a refinement is proposed to the current approach that explicitly incorporates the contribution of the inner row of anchor rods. The refined approach seeks to provide an improved characterization of connection response while also preserving simplicity such that the approach may be applied in a design setting.

Consideration of this additional set of rods introduces an indeterminacy in the derivations, which is resolved by establishing a linear relationship between the forces in the outer and inner rows of rods. Based on test data, the ratio of the forces in the inner rods to the outer rods is calibrated to be 0.5, such that Tnew Inner 0.5 Tnew outer. Application of the refined method to Tests 58 provides closer agreement with test data in terms of overall connection strength and anchor rod forces (in the outer row), i.e., the degree of conservatism is reduced with respect to the Steel Design Guide One method. Excepting some minor issues, the method provides a convenient way of calculating design forces. An interesting observation

(from the measured force data) is that the middle rod (in the outer row of rods) carries a disproportionate fraction (up to twothirds) of the total force in the row. This observation is important since current approaches do not address this and implicitly assume that each row contains only two anchor rods. Based on the test data and its assessment with respect to the two methods (Steel Design Guide One and the proposed refinement), this study has the following mplications for the design of these types of column base:

1. For base connections with a four-rod layout (and two rows of rods), the method presented in Steel Design Guide One is reasonable and conservative in characterizing the overall connection strength (controlled by base plate yielding) and the rod forces. These observations are based on the incremental forces introduced into the rods, not considering the prestress introduced into them during installation. It is recommended that estimated or measured prestress forces, which depend on the method of installation, be added to those determined by the Steel Design Guide One approach to designing anchors.

2. For base connections with three rows of anchors, it is recommended that the refined method be used to characterize incremental (i.e., not including prestress) forces in the anchor rods and the corresponding base moment capacity. If this is not

possible, then the Steel Design Guide One method may be used conservatively to characterize the forces in the outer row of rods, and the total force in the inner row may be determined as being half of this value (based on experimental observations). Once the force in each row is determined, special attention must be paid to the outer row if it contains three rods because experimental data suggest that the load is not shared equally among them. The central rod carries up to two-thirds of the total load in the row and should be designed with that mind. While the preceding points pertain to the calculation of design forces within the rod and the plate, perhaps the most important finding of the study is the excellent deformation capacity (and hysteretic response) of the base plates. This type of response is not anomalous but, rather, consistent with observations in previous studies that have similarly shown excellent deformation response in exposed-type base connections. As per current design practice for moment frames (e.g., SEAOC 2012), base connections are designed to be stronger than the attached column to force plastic yielding into the column rather than the base connection. Thus, base connections are often designed for amplified seismic loads that are based on the factor or the maximum expected column base moment. This necessitates detailing (such as thicker base plates or deeply embedded anchor rods, or sometimes even embedded columns) that are expensive. The assumption underlying this practice is that base connections are brittle and must be protected against force-controlled failure, while promoting yielding and plastic hinge formation in the adjacent column. The current study (and previous research on base connections) shows that this may not be the case and that base connections may be relied upon for inelastic response if detailed appropriately (notwithstanding issues such as reparability). On the other hand, some recent studies indicate that columns (especially deep columns susceptible to lateral torsional buckling) may not be as ductile in bending as previously thought. Ongoing efforts are focused on developing methodologies that effectively incorporate the deformation and energy dissipation characteristics of the connections into a design process.

Though this study provides new insights into the response of base connections and refinements to current approaches, it does have several limitations that must be considered when interpreting and generalizing the results. First, none of the tests featured axial compressive (or tensile) loading. For cantilever column systems, the axial loads are typically low (less than 10% of the concrete bearing

capacity); thus, care must be exercised in applying the findings of this study to situations where larger axial loads are present. Other loading scenarios, such as the presence of biaxial bending of the column, will limit the applicability of the resultscomputational research by Kanvinde et al. (2013) suggests that biaxial bending may compromise the strength of connectionsbut no experimental data are available to confirm this finding. Second, an obvious issue is that anchor rods were not embedded in concrete but rather passed through the concrete to enable the attachment of load cells and allows reusability of the concrete block. Third, though the prestress in the rods was measured and considered in this study, it may not be

representative of prestress in field details; appropriate adjustments should be made to design forces to account for the best estimates of prestress. Another consideration in the interpretation of results is that the evaluation of the design methods is based on limited test data, and generalization of these approaches should incorporate appropriate measures of reliability, e.g., resistance factors (-factors) developed through the analysis of previous standards, specifications, and similar test data. Fourth, in this study, columns were designed to remain elastic because the focus was on isolating the in columns. Nominally, this should not affectthe connection strength, but interactions between the column yielding and connection response cannot be ruled out. Finally, the proposed method was validated against the test data set used to develop it because no other data set is currently available. This limits the generality of the method, and future data sets may be used to refine the approach or to propose new methods. connection response, whereas in design, yielding may be expectedCHAPTER 8

CONCLUSIONIn conclusion, it is emphasized once again that the response of these connections is controlled by highly nonlinear and complex interactions between the various components. As a result, the development of a design (or strength characterization methods) that explicitly satisfies the equilibrium, compatibility (including gapping

and contact), and nonlinear constitutive response of the various components is intractable. Consequently, the methods presented in this study (as well as previous methods) are based on simplifying assumptions such as a predetermined stress distribution.This implies that caution should be applied in extrapolating the results of this study to any situation that is significantly dissimilar (in terms of geometry, plate size or thickness, loading, or other parameters) from those examined in this study.REFERENCES1. Can Akogul1 and Oguz C. Celik (2008) "Effect of Elastomeric Bearing modeling parameters on the seismis design of RC highway bridges with precast concrete girders"2. Hamid Saadatmanesh, Mohammad R. Ehsani, and Limin Jin"Repair of Earthquake-Damaged RC Columns with FRP Wraps" ACI structural journal Title no. 94-S20.

3. Hewes, J T and Priestley M.J.N(2002) "Seismic design and performance of precast concrete segmental bridge columns rep no SSRP 2001/25 Univ. of California at San Diego

4. Muhammad N. S Hadi (2007) "The behaviour of FRP wrapped HSC columns under different eccentric loads" University of Wollongong5. Olivov .K, Bilk .J (2009) "Strengthening of Concrete Columns with CFRP" Department of Concrete Structures and Bridges, Slovak University of Technology6. Pan J.L. *Yuan F. (2013) "Seismic behaviors of ecc/concrete composite beam-column joints under reversed cyclic loading" VIII International Conference on Fracture Mechanics of Concrete and Concrete Structures, Key Laboratory of Concrete and Prestressed Concrete Structures of Ministry of Education School of Civil Engineering, Southeast University, Nanjing, China7. Saiidi M, Vosooghi A, Cruz, Motaref S., Ayoub C, Kavianipour F, Haber Z,O'Brien, and David Sanders (May 2012), "Earthquake-Resistant Bridges of the Future with Advanced Materials," Keynote Speech, Ninth International Congress on Civil Engineering, 9ICCE, Isfahan, Iran,. 8. Sarira Motaref, Saiid M, David Sanders(2014). Shake Table Studies of Energy-Dissipating Segmental Bridge Columns, Journal of Bridge Engineering, Vol. 19, No. 2, February 1, 2014. ASCE

9. Yu-Chen Ou (June 2007) "Precast segmental post-tensioned concrete bridge columns for seismic regions" University of New York

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