نشریه مهندسی عمران و محیط زیست دانشگاه تبریز
سال چهل و سوم شماره 2 (پیاپی 71، تابستان 1392)

1. The use of marble rock flour residue was investigated in the laboratory for the potential use as a landfill liner material. The clayey soil, sand, and cement were used as the additives. The rock flour and the additive material were mixed, compacted, and tested for diffusion and the hydraulic conductivity. The obtained diffusion coefficients were compared with the recommended diffusion coefficient for a compacted clayey liner. The comparison showed that the mixed material containing 65% rock flour and 35% clayey soil is acceptable with regard to diffusive transport mechanism compared to other mixed materials. The hydraulic conductivity was determined by the flexible wall triaxial hydraulic conductivity apparatus and the value of 1.210-10 m/s was obtained for this mixed material. This amount of hydraulic conductivity is in the range of values recommended for the compacted clayey liners and therefore this material is acceptable in terms of advective transport. The results showed that the diffusion coefficient and the hydraulic conductivity of the compacted mixed material containing 65% marble rock flour and 35% clayey soil are in the range of the recommended values for the compacted clayey liners.2. Methodology2.1. The marbel rock flour, clayey soil, sand, and cement were used as the materials for this study. These materials were mixed in different percentages to obtain three types of mixed material. The type 1 mix included 67% rock flour, 27% sand, and 5% clayey soil. The type 2 mix included 65% rock flour and 35% clayey soil. The type 3 mix included 90% rock flour and 10% cement. The 100% pure rock flour was also tested. The material were compacted in a polyethylene tube and were tested for chloride ion diffusion coefficient. The mixed material type 2 was tested for the hydraulic conductivity.2.2. Theoretical modeling: The program POLLUTE [1] was used for the prediction of the chloride diffusion coefficient in the tested material based on the observed chloride concentrations data. The following diffusion coefficients were obtained: 5.310-10 m2/s for pure rock flour, 810-10 m2/s for type 1 mix, 5.510-10 m2/s for type 2 mix, and 710-10 m2/s for type 3 mix. 3. Results and discussion3.1. Diffusion: The comparison of the results for chloride diffusion coefficients showed that the mix material type 2 has the lowest diffusion coefficient. This material has 30% clayey soil which gives the mixture enough absorption capacity as a liner material [2]. Thus, type 2 mix material is found to be suitable as a liner material in the landfill applications compared to two other mixed material. Fig. 1 shows the observed and predicted concentration profiles for type 2 mix material.3.2. Hydraulic conductivity: The type 2 mix material was tested for the hydraulic conductivity using the flexible wall hydraulic conductivity apparatus. The hydraulic conductivity value of 1.210-10 m/s was obtained for this material.4. The type 2 mix which included 65% rock flour and 35% clayey soil materials resulted in a reasonable low diffusion coefficient of 5.510-10 m2/s which is in the range of the reported diffusion coefficients for the compacted landfill liner material. The resulted hydraulic conductivity of 1.210-10 m/s for this mix material is also acceptable and is in the range of the standard value for the compacted landfill liner material. Thus, the type 2 mix is recommended as an alternative liner material for solid waste landfill liner applications.

1. In recent decades, the use of FRP fibres in concrete structures has been considered by researchers due to their high corrosion resistance. The use of FRP composites in strengthening members of reinforced concrete structures such as beams and joints has been of great interest for civil engineers in recent years [1]. Various designing codes for using FRP material in concrete structures were presented. Many existing reinforced concrete buildings were already designed for gravity loads and lateral forces that may be much smaller than those prescribed by existing building codes [2]. Berg and others have studied on cost analysis of an FRP reinforced concrete bridge deck [3]. Their conclusions showed that construction of an FRP reinforced concrete bridge deck using conventional construction technology and labor was accomplished with a 57% savings in construction labor over nominally identical steel rebar reinforced deck2. Methodology2.1. Experimental study: In order to examine the effect of FRP stirrups distance on the cyclic behavior of concrete joints, two half-scale specimens were constructed with the same steel longitudinal bars (in columns and totally at the top and bottom of the beams) and same dimensions, 250 mm square columns and 200 by 250 mm beams (height by width). The first joint had closed space FRP stirrups considering design code ductility regulations and was introduced as FDJ (FRP Ductile Joints), and the other had wider double spaced FRP stirrups and was introduced as FNJ (FRP Non-ductile Joints). Characteristics of the joints have been shown in Fig. 1 and given in Table 1.2.2. FE modeling: ANSYS finite element program was used to examine the behaviour of these two connections (with different distances of stirrups) and the calibration or prepared models against laboratory test results. The finite element models were generated at two different groups, each group with two different stirrup spacing.Table 1. Details of experimental joint specimens with CFRP materialsSpecimen Number of rods (14 mm) Number of rods(12 mm) Stirrups Distance in the ductile region (mm) Stirrups Distance in the non-ductile region (mm) Number of stirrups in joint Bottom Top Beam Column Beam Column FDJ 8 3 3 50 50 100 150 2FNJ 8 3 3 100 100 100 150 13. The distance between stirrups has an important effect on the cyclic behaviour of connections and the amount of energy absorption and dissipation. The distance between stirrups is much influential on the behaviour of RC connections. To design the FRP stirrups cross section, equivalent cross section of steel stirrup was adopted by assuming 6 mm diameter steel stirrups and 28.3 mm2 cross section and 260 MPa yield strength. According to design code regulations of ACI 318, shear capacity of steel stirrups was calculated by Eq. (1).Where Asv, Fy, d and Ss are the sum of the areas of two steel stirrups legs, steel bar yield strength, effective height and spacing of steel stirrups, respectively. If FRP fibers are used, shear capacity of FRP stirrups is calculated by Eq. (2) based on ACI 440 and Canadian design code CSA.Where Vf, Afv, s and ffv are shear capacity of FRP stirrups, the sum of the areas of two FRP stirrups legs, spacing of stirrups, and effective tensile strength of FRP stirrups. According to ACI-440-R1, effective tensile strength in FRP stirrups is assumed to be 0.004 Ef when Ef is the modulus of elasticity of FRP stirrups. Required equivalent cross section of FRP stirrups was calculated according to Eq. (3) by assuming 0.4% strain for FRP stirrups and equality of shear capacity of FRP and steel stirrups. 4. The experimental results showed that the additional capacity of the specimen with closely spacing stirrups was 12% more than that of the companion specimen with widely spacing stirrups, and also the ductility coefficient of the former specimen was 26% higher than of that of the later specimen. The results also showed that the maximum strains gained at the connection panel stirrups was about 50% more that the maximum allowable amount given at the design code provisions. Numerical analysis showed that all specimens reinforced with FRP stirrups designed based on design code ductility regulations at closed space had averagely 8% higher loading capacity and joints with closed ductile stirrups had higher ductility coefficients up to 20% compared to joints with wide stirrup spacing.

The concrete-filled steel tube (CFT) column is a composite system that is made with concrete as inner core and thin walled steel as outer tube. The enhancement of load capacity, increase in lateral stiffness, inherent ductility, use in high rise and long span buildings and fire resistance, are some benefits of CFT columns [1].Columns are the most parts of buildings in fire that carry loads of structure. When fire happens in a structure with CFT columns, the steel tube is the first part of these columns, subjected to fire. Therefore, the steel tube expands and concrete core carries remaining loads. It's a main disadvantage of CFT columns [2]. In the present study, finite element modeling of stiffened circular CFT columns is presented. The main characteristic of the stiffened CFT column is internal longitudinal symmetric stiffeners its internal stiffeners that are longitudinal and symmetric. For predicting the load capacity of the proposed sections of CFT columns, the time – temperature curve is applied to the model.2. Methodology2.1. Fire exposure: The temperature of the column rises up (Eq. (1)) and cools down (Eq. (2)) according to the ISO-834 fire curve (Fig. 1) [3-5]. The temperature is applied to nodes of the columns.where T is the fire temperature in ◦C; t is fire exposure time in minute; ‘‘B–C–C’’ is the heating phase; ‘‘C–D’’ is the cooling phase; ‘‘C’’ is the starting point of cooling; To is the ambient temperature in ◦C; th is the fire duration time in min; Th is the maximum fire temperature in ◦C and tp is the total fire exposure time in min.2.2. Heat transfer analysis: The heat transfer process in concrete filled steel tubes includes radiation, convection and heat conduction, all of them are varying with time. There are four heat transfer processes, from fire to outer steel tube surface, from outer to inner steel tube surface, from inner steel tube surface to outer concrete surface and heat transfer into core of concrete from concrete surface [6,7].2.3. Suggested sections: One of the main disadvantages of CFT columns is the expansion of the steel tube when subjected into the fire. Therefore, the confinement of core and finally, the load capacity of CFT column will be decreased. In order to improve the performance of CFT columns, subjected to the fire, a stiffened steel section is suggested in CFT columns as shown in Fig. 2. The stiffeners enhance the confinement of columns by increasing the contact area between steel and concrete. Hence, the confinement of CFT columns is preserved in heating and cooling stage of the fire exposure. Therefore, it is expected that the load carrying capacity of CFT column to be increased.2.4. Finite element modeling: In the present study, finite element modeling of stiffened circular CFT columns is presented. These models have been analyzed using ANSYS (Ver. 11) [8]. In order to investigate into the behavior of the stiffened CFT columns under axial and thermal loadings,FE analyses should be undertaken involving the geometric and material nonlinearities. Three-dimensional solid element (Solid 65) has been used for modeling of the concrete core, shell element (Shell 43) has been used for modeling of the steel tube. Three-dimensional node-to-node contact element (Contact 178) has been used to model the contact between steel wall and concrete core. Modeling of separation, sliding and contact between two nodes during the loading process are the capabilities of this element. Fig. 3 shows the CFT column, modeled in ANSYS.2.5. Verification of finite element modeling: The material properties of concrete and steel are changing when subjected to the fire. Therefore, some researchers define material properties in ambient temperature, heating and cooling phases. In order to verify the accuracy and validity of the finite element modeling, the numerical results, obtained from nonlinear static analysis, have been compared with the experimental results of the CFT columns. Regarding the axial loading, an experimental CFT column with circular section has been used according to the specifications undertaken by Schneider [9] and in order to verify the accuracy and validity of the equations used for applying fire to the CFT column, model specifications are according to Han [10]. Consequently, it has been found that the finite element model is reliable enough to be used to undertake nonlinear analyses for comparative investigation into the behavior of the sections of CFT columns with and without stiffeners.2.6. Material heat coefficients: The thermal coefficients in CFT columns have been investigated by many researchers. In the present paper, the coefficients for concrete and steel tube, provided by Lie and Stringer [11], are adopted. The thermal properties such as conductivity coefficient (k), specific heat(c), density (ρ) and coefficient of expansion (α) have been defined in these equations. Also, they considered the water evaporation of core concrete after sbeing exposed to the fire. In the evaluation of temperature, k and α decrease, and c increases after a few minutes of fire exposure. They also considered these specifications of concrete and steel in their relations.3. A total of 54 CFT stub columns, in three types, including simple, 2 and 4 stiffened circular CFT columns, have been analyzed after subjected to fire. The specimens are divided into two groups (with the same wall thickness and with the same diameter). All of the specimens, have the same wall thickness and the D/t ratio increased in three steps. All of the specimens are subjected to real fire in ANSYS in three time durations, namely 10 minutes (680oC), 20 minutes (780oC) and 30 minutes (840oC) according to standard fire curve, shown in Fig. 2. In suggested sections the area of stiffeners, is 30% of total area of CFT column and L=3810mm. For steel, we have: Fy= 293MPa and Es= 2.01e5MPa. For concrete, we have: fc= 39.6MPa and Ec= 27800MPa.4. In the present study, finite element modeling of stiffened circular CFT columns is presented. The main characteristic of the stiffened CFT column is its internal stiffeners that are longitudinal and symmetric. These models have been analyzed using ANSYS. Having verified the finite element modeling, several different analyses have been undertaken. Based on the results of analyses, the main conclusions are as follows:-Increasing the tube diameter in simple and stiffened CFT columns has increased the load carrying capacity of columns. The increase rate of the load capacity in CFT columns during fire exposure is more considerable than ambient temperature.-Decrease in the wall thickness of CFT columns, in simple and stiffened columns, decreases the load carrying capacity of columns. The decrease rate of the load capacity in CFT columns during fire exposure is more considerable than ambient temperature.-Increasing the number of stiffeners has enhanced the load carrying capacity of CFT columns. However, increasing the number of stiffeners from 2 to 4 dose not have considerable effect on the load carrying capacity of CFT sections (only 2%).Therefore, the use of 2 stiffeners is more economic, and strongly recommended for CFT columns subjected to fire.-The residual strength index (RSI) is defined to quantify the strength of the CFT columns to standard fire. RSI is the ratio of ultimate strength corresponding to the fire duration time into the ultimate strength at ambient temperature [10]. According to the analysis results, decreasing the number of stiffeners and increasing the tube diameter, has enhanced the RSI of CFT columns.

1. The estimation of lateral earth pressures on retaining walls has been an earliest subject in geotechnical engineering. The earliest lateral earth pressure theories were suggested by Coulomb (1776) and then by Rankine (1857) and remain the basis for the present earth pressure calculation. In fact, Coulomb’s procedure can determine the static lateral total thrust on the wall rather than the distribution of earth pressures. However, the subject has been extensively expanded for better determination of lateral earth pressures on retaining structures. In the present study, a hybrid approach for estimating the total thrust and height of its application is presented. For this purpose, Kotter’s equation [1] of characteristic method is used to determine the governing equations of the failure surface. Then, on the basis of limit equilibrium approach and pseudo static methods by using global equation of forces acting on the wedge, the magnitude and position of application application point of total thrust are determined. The present study considers the inertia effect of the stable soil medium on the failure wedge. Using an analytical procedure, a closed form solution for computing the angle between the failure surface and the horizontal direction is developed. In addition, the height of the application point of total thrust has also been calculated and compared with those obtained from earlier research work. In addition, the method is expanded for layered backfill.2. The present method applies limit equilibrium approach similar to the Mononobe-Okabe (MO) method and first finds seismic reaction pressures on the failed wedge. For this purpose, the retained soil is assumed to be cohesionless, homogeneous, isotropic, semi-infinite and dry. Also for the backfill soil, no effect of strain softening or liquefaction is considered in analyses. Given the global equilibrium equations of forces,, which contain three unknowns (, and), another equation is necessary for rendering the problem statically determined. This equation can be obtained by using the condition. It is useful to choose the angle such that it maximizes the active thrust, as employed in Coulomb’s method, or uses one of Kotter’s equations on the slip surface, as done by Dewaikar and Halkude [1] and also used in the present study.By using Kotter’s equations on the slip surface, this paper can be considered as a simplified approach with respect to the method of characteristics. It assumes a slip surface of arbitrary form (planar), while in the method of characteristics, the form of the slip surface is part of the solution. Moreover, only one of the equations of Kotter is considered, but the problem is governed by a system of four differential equations in the method of characteristics.3. It is well known that in the retaining wall design for overturning stability, the application point of the total thrust is important. The main strong point of the approach of present method is the possibility of determining the position of the point of application of the total thrust. It has been shown that there is no difference for total thrusts determined from the present method and that computed from both Coulomb’s and MO methods. This position is obtained from moment equilibrium. As mentioned before, results of many experiments generally suggest that the application point of the resultant thrust be located higher than one third of the height of the wall and should be located at mid-height of the wall (Sherif and Ishibashi [2], Bolton and Steedman [3], Steedman [4]).The height of the total active thrust application predicted by the present method is similar to that obtained from tests carried out by Sherif and Ishibashi [1]. In addition, it is quite close to that given by Seed and Whitman [5]. It is noted that Al Atik [6] believes that the application point of seismic total active thrust predicted by the MO method is assumed to be at 0.33H from the wall bottom. 4. This paper has presented a theoretical solution for determining the magnitude and height of application of total active and passive thrust exerted on rigid retaining walls. The pressure distribution on the backfill failure surface has been determined using Kotter’s equations. For both active and passive seismic conditions, the angle of failure surface with the horizontal direction, total lateral thrust, soil reaction on the failed wedge and the point of application have been determined. The results of present study can be summarized as follows: The active and passive thrusts and angles of failure plane with respect to the horizontal direction are identical in the present method and both the Coulomb and MO methods. The height of the point of application of seismic total thrust determined using the developed method is in agreement with experimental results obtained in earlier research works. The developed closed-form of the solution enables the designer to determine total active\passive lateral thrusts, point of application of total thrust, and estimation of overturning moment on retaining walls.It has been found that the seismic soil reaction from the stable soil on the backfill failed wedge can not be ignored, since such ignorance would result in overestimation of seismic passive force, underestimation of seismic active force, and under-prediction of the application point of seismic active force. All these may lead to the unsafe design of retaining walls.

1. Concrete buildings reinforced by plain (smooth) bars are one of the special types of old reinforced concrete buildings. They were generally built before 1970’s; mostly in Europe, Asia and Oceania. Some of the older cases of such buildings were probably designed just for gravity loads and do not have special seismic detailing for structural members (e.g. beams, columns, joints, etc.) because the old codes did not include special seismic provisions at that time. After recent earthquakes, seismic vulnerability of old reinforced concrete buildings has became more highlighted and subsequently, demand for seismic evaluation and rehabilitation of such buildings have experienced a remarkable growth in recent decade. The first stage for proposing a rehabilitation strategy for an existing building is the seismic vulnerability assessment. In this stage a structural engineer tries to understand the behavior of the structure under seismic excitations. This behavior is controlled mostly by the weak links of the structure. Recognition of the structurally weak links under earthquake ground motions helps the engineer to define a proper rehabilitation strategy for the improvement of structural seismic performance. To achieve a better understanding of structural seismic behavior, it is essential to perform a proper analysis considering the nonlinear behavior of structural members. As columns are generally the most important structural members of a framed structure, understanding their realistic seismic behavior is very helpful in estimating structural deformations, forces and energy dissipation capacities. Furthermore, in most of old framed building structures, columns play a key role in the final behavior because of strong beam-weak column conditions. 2. Methodology2.1. Experimental study: Four cantilever half-scale column specimens were tested in this study under monotonic and cyclic loading protocols. Specimens had square sections and were conneced to a strong foundation. Concrete casting was done in two steps in vertical position. All of the specimens’ reinforcements consisted of four longitudinal plain (smooth) bars of 12 mm diameter. To investigate the effect of various types of old splice detailing practices, three types of specimens with various longitudinal reinforcement splicing were considered: specimens Without Overlap Splice (WOS specimens) which were the representative for column detailing below the floor level, Specimen with Straight Overlap Splice (SOS specimen) of 40 times of the bar diameter length representing old American practice for column detailing over the floor level and finally specimen with Hooked Overlap Splice (HOS specimen) of 20 times of the bar diameter length which was introducing old European practice for column detailing over the floor level. WOS specimens were built twice, one for monotonic test (WOS-M) and one for cyclic test (WOS-C). Transverse reinforcement details were similar in all of the specimens; i.e. plain bars with the diameter of 8 mm in 200 mm spacing intervals. The distance was adopted to be equal to the column effective depth. Full details of specimens are presented in Fig. 1. All of the specimens were tested under constant axial load equal to 15% of concrete gross section axial capacity. Loading history protocol according to ACI T1.1 [1] recommendations was applied. The history was displacement (drift) control with preliminary target drifts. Target cylindrical compressive strength for concrete specimens at the age of 28 days was 22.5 MPa.3. Results and discussion3.1. Test observations: Similar crack patterns were observed in all the tested specimens up to drift ratio of 2.20%. Three flexural cracks were formed in three elevations; i.e. base elevation (base-crack), approximately half-depth from base elevation (h/2-cracks) and approximately one-depth from the base elevation (h -cracks). In WOS-M and WOS-C specimens base-cracks were the most active cracks with larger openings; meanwhile h/2-cracks were more active than h cracks and had an inclined trend like flexure-shear cracks. In SOS-C and HOS-C specimens (-C represents cyclic loading protocol), base-cracks were similarly the most active cracks same as WOS specimens; however, h-cracks were more active than h/2-cracks and presented an inclined trend. It was visually obvious that base-cracks play a key role in total deformations of specimens and dictate a rocking like mode for all of the specimens; even in small drift ratios.3.2. Force-drift response of specimens: The response curve of WOS-M specimen was approximately linear up to drift ratio of 0.63% and then turned to nonlinear phase. WOS-C specimen had an origin-oriented hysteresis response. In the unloading phase, after fast degradation of shear strength, the curve was tending to the origin. In other words, high pinching and low residual drifts were the main characteristics of hysteresis response. SOS-C specimen same as WOS-C specimen demonstrated an origin oriented hysteresis response but in different flag-shaped appearances. Similarly, reversal curves tend to zero drifts with a negative slope in the second step of unloading phase after fast shear strength degradation and finally, near the zero drifts, the curves were tending to the origin. Consequently, high pinching and low residual drifts existed same as WOS-C specimen. The hysteresis response curves of HOS-C specimen are very similar to SOS-C specimen but with more pinching effects because of lower initial strength degradation at the first steps of unloading phase.3.3. General mode of behavior of specimens: The main characteristic of behavioral mode of specimens was rocking movement around a point near the toe. The distance between the rocking rotation point and toe (contact depth) had an approximately constant value of 50 mm after 1% drift ratio (based on experimental observations). Contact depth was spalled at drift ratio of 2.20% and then it was crushed at drift ratio of 3.5%. After the crushing, rocking continued around a point near the previous point. One may conclude here that the main governing phenomenon for specimen deformations is rocking action which is restrained by two rows of plain bars at both sides.3.4. Restrained-rocking model: As mentioned before, restrained rocking mechanism was the governing behavior mode of specimens, such as rocking of a rigid body block around its toe. However, the block was not rigid itself in this case and it was deforming in flexure and shear modes. The formation of some cracks (1 or 2) at the bottom height of specimens was evident. The cracks were active until the end of test period and contributed in total deformation of the specimens. To achieve a better understanding of load-displacement hysteresis response of the specimens, it was initially assumed that the specimens did not have longitudinal reinforcement continuity between the column and foundation. In other words, a concrete block was assumed under an axial load. The block was loaded horizontally as well. Roh and Reinhorn [2] studied the behavior of such an element and called it rocking element. Next, it is assumed that plain bars are added to the rocking element. It was obvious that before the total bond deterioration of dowel bars, yielding strength of bars at tension and compression was added to the rocking strength; thereafter, bond resistance was added. Hysteresis response curve was located approximately between two linear upper and lower bound lines which were nearly parallel to the apparent negative stiffness part of the rocking curve. The upper bound line was parallel but apart from the rocking curve negative stiffness line for a constant added strength above it, while the lower bound line was similarly parallel the rocking curve but apart for another constant value below. The upper bound line passed through maximum displacement points of hysteresis curve for cycles with peak drift ratio more than 2.2%. In other words, it had a common part with the negative stiffness slope line of backbone curve which was initiated at Negative Stiffness Initiation (NSI) point that coincided with spalling occurrence onset. The lower bound line passed through points of slope changing in unloading branch for cycles with peak drift ratio more than 2.2%. Better understanding can be achieved if we consider two upper and lower bound curves instead of upper and lower bound lines. The upper bound added strength before the NSI point can be referred to the yield properties of plain bars. After NSI point, it can be referred to the bond properties of dowel bars, especially in SOS-C and HOS-C specimens; while in WOS-C specimen, it can be referred to a combination of yield and bond properties. It’s worth noting that shear strength dropped between upper bound and lower bound curves (at the first step of unloading stage), passed through the rocking action curve. It may be assumed that the strength difference between upper bound and rocking curve arises from full tension yielding release or full tension bond deterioration. On the other hand, the distance between rocking curve and lower bound curve arises from compression yielding development or compression bonding development in the opposite direction of concrete block movement. In fact, the lower bound curve is the rocking action curve minus an opposite yield or bond resistance. The described sequences are demonstrated in Fig. 2. The main concept here, is the fact that the shear strength difference between upper and lower bounds is the summation of two strength values; one is the released strength from upper bound to rocking curve (f1 or f3) and the other one is developed from rocking curve to the lower bound (f2 or f4); totally form an origin oriented flag-shaped hysteresis response curve.4. According to the monotonic and cyclic tests that were performed on four half-scale cantilever column specimens under low axial loads, the following main conclusions are developed;•Damage pattern of concrete columns with plain bars consists of limited numbers of flexural cracks that are formed at the bottom of the column specimen, in which the base crack (crack at the interface between column and foundation) has large opening. •Hysteresis force-drift curves of plain bar specimens are origin oriented in the specimen without overlap splice and are flag-shaped in specimens with overlap splices. High pinching effects and low residual displacements are the main characteristics of all the hysteresis curves. •General mode of behavior of all specimens seems to be restrained rocking action, independent from types of splice detailing. This is proved through experimental calculated rotations which are derived from LVDT readings.•A simple theory for the explanation of hysteresis force-displacement response of specimens is proposed. The theory assumes a concrete block rocking element which is restrained with plain bars at both ends.

1. Soil-steel structures recently have been used extensively in different countries as a highway or railroad bridge. Due to the licity and ease of construction method, these structures were successful variants in the road-railway unleveled crosses’ projects. Common standard formulas for minimum depth of cover are based on highway loading pattern with significance of preventing soil tensile and wedge sliding failures. Modification of minimum depth of cover for railway bridges is needed not only in respect to different modes of soil failure but also for non-uniform soil settlement which endangers the safety of railway track. Therefore, regarding permissible track settlement and wall buckling control, new pattern of minimum depth of cover was developed for boxes and high-profile arches individually using 2D finite element analysis. However, three dimensional distribution of railroad could not be taken into account by 2D FE method. Therefore, modification of developed pattern with attitudes toward longitudinal distribution of railway load is the scope of this study. In this paper, regarding permissible track settlement, railway twist, longitudinal settlement of the structure and wall buckling control, the minimum depth of 32 structures with spans greater than 8 meters was determined by 3D finite element analysis. Then the resulted pattern was compared to those obtained from 2D analyses. Finally, the modification factors were calculated by least squares method and new formulas were established for boxes and high-profile arches individually. 2. The primary objective of the current study is to introduce a new set of minimum soil cover equations for long-span railway bridges based on the numerical interpolation of results of the 2D FE analyses. To determine the variation trend of the minimum depth of cover along with its governing parameters (geometry, length of span and the panel stiffness) through numerical analyses, the permissible settlement of the track, metal structure buckling and soil body failure criteria have been checked initially for each bridge structure for a 0.6 m depth of cover (the minimum limit of cover depth specified by CHBDC). When all of the defined criteria were not fulfilled simultaneously, the depth of soil cover above the crown was increased, and the analyses were then restarted for a new depth of cover. The minimum depth of cover in which all of the criteria were simultaneously fulfilled was chosen as the minimum depth of soil cover for a specific bridge structure. In this manner, the results of the 2D FE analyses present specific patterns for the calculation of the minimum depth of cover for box culverts and low-profile arches. In order to check the applicability of the proposed equations for minimum depth of cover in practical problems, a series of 3D finite element analyses with more realistic idealization of the railway superstructure components and the lateral slope of bridge embankment were carried out-of-plane the out of plan buckling in the steel plates. 2.1. FE modeling: The multi-purpose FEM-based software package, PLAXIS was used for the numerical modeling and analysis. Regarding to the special serviceability criteria for railway bridges, for computing the minimum depth of cover, the permissible settlement of the railway track and the buckling of the conduit walls in different sections along the longitudinal axis of the bridges in 3D analyses were controlled. For these cases, the spans from 8.07 m up to 13.46 m of box culverts and 14.13 m up to 23.40 m of low-profile arches using stiffened and non-stiffened deep corrugated panels have been considered. This study assumes that all the metal structures are buried in well-graded gravel (GW) as engineered backfill material [1]. The material nonlinearity of the soil and metal structure as well as the stage construction effects were accounted for in the numerical analyses, and the railway load model LM71 [2] was applied.3. Results and discussion3.1. The results of the 2D and 3D finite element analyses: As a result of various numerical analyses and satisfying the aforementioned criteria, the minimum depth of cover was evaluated against the span of railway box and low-profile arch bridges using various panel types (stiffness). This is due to the different structural geometry of the boxes and results in a different mechanism of behavior under the loading. In order to account for the moment of inertia, the speed and length of span which are the representative of panel stiffness, the effect of dynamic factor and geometry of metal structure, the basic form of the minimum depth of cover for the railway boxes and high-profile arches is introduced as:Where α, β, γ and µ are unknown constants. These constants were calculated separately for the boxes and low-profile arches by using the least-squares method (LSM) to determine the function of best fit.An approach similar to the 2D analysis was used for the results of the 3D FE method, which resulted in modifications to the previously developed equations. The major finding of the comparison between the results of the 2D and 3D FE analyses was that the resulting minimum depth of cover from the 3D FE analysis was smaller compared to the 2D FE results for the same conditions. This difference was due to the stiffening effects of the railway track due to the third dimension of the bridge that was considered for the analysis. Therefore, the difference between the 2D and 3D FE results can be accounted for by using the reduction factor in the (IVI/I) ratio. Consequently, Eq. (1) can be modified as:for boxes (2)for low-profile arches (3)where and were calculated by using the least-squares method for a train speed of 120 km/h. The R-squared values of Eqs. (2) and (3) are R2=0.84 and R2=0.97 for the box culverts and low-profile arches with panels (III) and (IV), respectively.3.2. Validity range of the derived equationsThe conformity of Equations (2) and (3) with 2D and 3D results was evaluated using an R-squared value that was calculated by using R2=1-SSE/SST, where SSE=  and SST=()-[(Yi)2]/n. For panels III and V, the R-squared values for Eqs. (2) and (3) were 0.67 and 0.88, respectively. Panels I and II were eliminated from the calculations because the 2D FE analysis demonstrated that these panels were not suitable for soil-steel railway bridges. The minimum depth of cover calculated from the 3D FE analysis of bridges with panel VI represented a depth of cover less than 0.6 m. A minimum of 0.6 m was maintained for railway track maintenance (complete removal of ballast layer in some operations). Therefore, the data for panel VI was not considered in the calculation of the R-squared values. Consequently, the limited amount of data for 3D FE analysis resulted in equations with less accuracy. The results of the 2D and 3D FE analysis were compared to the standard limits and are shown in Fig. 1. Regarding the maintenance operations, buckling criterion and AASHTO [3] and CHBDC [4] limits, a recommended minimum of 0.5 m and maximum of 1.5 m must be maintained for the depth of cover by using the following expressions: for boxes (120 km/h) (4) for low-profile arches (120 km/h) (5)4. The minimum depth of cover requirements given by different codes are typically based on vehicle loads, non-stiffened panels and only the geometrical shape of the metal structure to avoid the failure of soil cover above a soil-steel bridge. In this paper, the effects of spans larger than 8 meters (using stiffened panels under railway loads) are investigated using an FE analysis. For this study, 2D and 3D FE analyses of four low-profile arches and four box culverts with spans larger than 8 meters were performed to develop new patterns for the minimum depth of soil cover. Using the least-squares method to adopt the best-fit equation of the numerical data, two new sets of formulas were recommended. Based on the numerical results, the primary research findings are summarized as follows: 1) The minimum depth of cover increases exponentially along with an increase in the span of boxes and low-profile arches. 2) The efficiency of the stiffened panels in reduction of the required cover depth is more pronounced for large spans. 3) Different trends of the minimum depth of cover were determined for box bridges and low-profile arches. This difference is due to the various structural geometry of the boxes that resulted in a different mechanism of behavior under the loading. 4) The modified exponential forms of the minimum depth of cover for railway boxes and high-profile arches is applicable for a train speed of 120 km/h and exhibits relatively good conformity with 3D FE results (R2 > 0.6). 5) With respect to the permissible settlement criterion, panels with EI greater than 33062 (kN.m2/m) are found to be the only suitable panels that can be used for high speeds trains (greater than 160 km/h) for railway box and low-profile bridges.

1. Self compacting concrete (SCC) is an innovative concrete that does not require vibration for placing and compacting. This kind of concrete has been immensely used in the construction for the last decade and it is inevitable to be aware of its behavior, especially time depending deformations such as the creep and shrinkage. In concrete structures deformations due to the creep and shrinkage are several times larger than elastic deformations. Frequently, these deformations cause excessive cracking and deflections or possible failure with an inherent loss in serviceability, durability and long-time safety of concrete structures. Thus, there is an urgent need for a reliable method to predict creep and shrinkage, especially for self compacting concrete structures. In this research, three mix designs were prepared and 21 specimens were fabricated from each mix design. Then the effects of varying percentages of micro silica the on creep and shrinkage of during a period of 250 days, and also on its mechanical properties including compressive strength, tensile strength and modulus of elasticity up to the age of 28 days were investigated.2. Methodology2.1. Materials: The sand used was this study is of river type with sand value of 95% and fitness modulus of 3.1. The gravel used was broken gravel with maximum grain size of 12.5 mm and superficial specific gravity of 2700 kg/m3. Also the micro silica used was with specific gravity of 2200 kg/m3. Ordinary drinking water was used in the mix designs and since limestone powder reduces the interstices between gravels and improves the adhesion of the paste; in this study stone powder was used as filler. Moreover, in order to reach the intended self-compaction, the superplacticizer Sika Viscocrete was applied as a percentage of cement and micro silica.2.2. Testing method and curing: In this paper, three mix designs of self compacting concrete with compressive strength of 300 kg/m2 (SCC 30), 350 kg/m2 (SCC 35) and 450 kg/m2 (SCC 45) were prepared. According to Table 1, in all three types of concrete, only the amount of micro silica was altered, so that its effects on long-term deformations of the concrete including the creep and shrinkage can be studied. Using L-Box test and Slump Flow [1], the properties of fresh SCC was evaluated and after curing the specimens, compressive strength test [2], indirect tensile strength test [3], elastic modulus of elasticity until the age of 28 days [4] and the creep and shrinkage test [5] during 250 days were conducted on the specimens.Table 1. Mix designsComponents SCC 30 SCC 35 SCC 45 ((kg/m3 Cement 280 280 280(Gravel (kg/m3 647 644 638(Sand (kg/m3 971 965 957 (Water (kg/m3 167.3 171 175.4(Microsilica (kg/m3 8.4 14 22.4 (Limestone powder (kg/m3 270 270 270 (Superplacticizer (kg/m3 6.34 6.47 6.65Water-cement ratio 0.58 0.58 0.583. Results from slump flow and L-Box tests are within the allowable range in the European Code [1]. The relation between compressive and tensile strength is presented in Fig. 1. As shown in this Figure compressive and tensile strength increase with a correlation factor of R2=0.902 as the amount of micro silica is increased.According to the results and considering that elastic modulus of concrete is affected by the elastic modulus of its component, modulus of elasticity increases as micro silica and the compressive strength increase. This relation is illustrated in Fig. 2.As it can be seen from Figs. 3-5, all strains have the same trend over the time. In early ages, the strains considerably rise and the slope of the initial segment of the curve, until 50 days, is very steep. Afterwards the increase rate is reduced at old ages nearly to a constant value. 4. The addition of micro silica to SCC mixture improves its mechanical properties and as a consequence, the modulus of elasticity, compressive and tensile strength increase. The relation between indirect compressive and tensile strength of the tested specimens shows that as the compressive strength of SCC increases the tensile strength also increases though at a lower rate. Since the nature of shrinkage and creep is influenced by the same factors, the creep and shrinkage decrease in the specimens as the micro silica proportion is increased.