Estimation of Compressive Load Bearing Capacity of Helical Piles Using Torque Method and Induced Settlements By Muhammad Umair Shabbir Khan A THESIS SUBMITTED IN PARTIAL FULLFILMENT OF THE REQUIREMENT FOR THE DEGREE OF MASTER OF APPLIED SCIENCE in THE COLLEGE OF GRADUATE STUDIES (Civil Engineering) THE UNIVERSITY OF BRITISH COLUMBIA (Okanagan) September 2016 © Muhammad Khan, 2016 The undersigned certify that they have read, and recommend to the College of Graduate Studies for acceptance, a thesis entitled: Estimation of Compressive Load Bearing Capacity of Helical Piles Using Torque Method and Induced Settlements Submitted by Muhammad Khan in partial fulfillment of the requirement of the degree of Master of Applied Science . Dr. Sumi Siddiqua, Faculty of Applied Science/School of Engineering Supervisor, Assistant Professor (please print name and faculty/school above the line) Dr. Shahria Alam, Faculty of Applied Science/School of Engineering Supervisory Committee Member, Associate Professor (please print name and faculty/school in the line above) Dr. Zheng Liu, Faculty of Applied Science/School of Engineering Supervisory Committee Member, Associate Professor (please print name and faculty/school in the line above) Dr. Liwei Wang, Faculty of Applied Science/School of Engineering University Examiner, Assistant Professor (please print name and faculty/school in the line above) August 26, 2016 (Date Submitted to Grad Studies) ii Abstract Helical piles are deep foundations that have a helix at the end. The traditional approaches to determine the load capacity such as loading tests and in situ tests (i.e. SPT, CPT and LCPC) are not economically feasible for the small scaled constructions, for which helical piles are generally recommended. In order to estimate the ultimate load that helical piles can carry, torque method is thus mostly used. Torque method does not account for the possible settlements induced at calculated loads. Settlement induced is the main load capacity governing factor for deep foundations, as they are considered failed when a settlement more than the permissible amount is attained. The possibility that the piles might fail well before the calculated load is achieved because of excessive settlements make the results of torque method dubious. This research attempts to investigate the torque method for the settlements and for its precision. For this purpose, seven RS2875.203 helical piles were installed and their ultimate compressive loads are calculated using the torque method. On seventh pile, static axial compression test was conducted. The settlements at torque method’s ultimate loads are determined from the load movement curve of compression test. Results underscore that the settlements at torque method’s ultimate and allowable loads are within the permissible amount. The load movement curve of compression test is interpreted using different failure criteria to calculate the failure load. Results show that 10% failure criterion is the most suitable criterion to interpret the load movement curve of RS2875.203 helical piles. Additionally, different bearing equations are used to compute the ultimate compressive loads of helical piles. Result suggest that the loads calculated using torque method and bearing equations correlate well with each other. iii Preface This thesis is the original work done by the author under the supervision of Dr. Sumi Siddiqua. A research article comprising of this research has been submitted in scientific journal: The contents of this research has been submitted to a journal: “Estimation of Compressive Load of Helical Piles Using Torque Method and Induced Settlements”. iv Table of contents Abstract .......................................................................................................................................... ii Preface ........................................................................................................................................... iii Table of contents .......................................................................................................................... iv List of tables.................................................................................................................................. ix List of figures ................................................................................................................................. x List of abbreviation and notations............................................................................................. xii Acknowledgements .................................................................................................................... xiv Chapter 1 Introduction ............................................................................................................. 1 1.1 Objectives ......................................................................................................................... 4 1.2 Test program .................................................................................................................... 4 Chapter 2 Literature review .................................................................................................... 6 2.1 Failure modes for helical piles ......................................................................................... 7 2.1.1 Individual plate bearing model ................................................................................. 7 2.1.2 Cylindrical shear plane bearing model ..................................................................... 8 2.2 Design of helical piles .................................................................................................... 10 v 2.2.1 Direct methods ........................................................................................................ 10 2.2.2 Indirect method ....................................................................................................... 12 2.3 Torque method ............................................................................................................... 13 2.4 Bearing equations ........................................................................................................... 16 2.4.1 Load bearing equation for shallow foundations...................................................... 18 2.4.2 Load bearing equations for deep foundations ......................................................... 21 2.4.2.1 Bearing factors for pile foundations ................................................................ 22 2.4.2.2 Meyerhof method ............................................................................................ 24 2.4.2.3 Vesic (1977) method ....................................................................................... 26 2.4.2.4 Load bearing equation for helical piles ........................................................... 27 2.5 Pile load tests .................................................................................................................. 30 Chapter 3 Installation of helical piles and site investigation ............................................... 32 3.1 Configuration and material properties of helical piles ................................................... 32 3.2 Soil investigation ............................................................................................................ 37 3.2.1 Soil classification .................................................................................................... 37 3.2.2 Direct shear test....................................................................................................... 39 vi 3.2.3 Water contents of soil ............................................................................................. 41 3.2.4 Unit weight of soil .................................................................................................. 42 3.2.5 Maximum dry unit weight and optimum moisture content .................................... 42 Chapter 4 Static axial compression test ................................................................................ 44 4.1 Introduction .................................................................................................................... 44 4.2 Types of static axial compression test ............................................................................ 45 4.3 Components and requirements of loading assembly ...................................................... 47 4.4 Interpretation of static axial compression test ................................................................ 49 4.4.1 Davisson offset limit load criterion......................................................................... 50 4.4.2 10%, 8% and 5% failure criterions ......................................................................... 51 4.4.3 L1-L2 failure criterion ............................................................................................ 53 4.4.4 Chin failure criterion ............................................................................................... 56 4.4.5 Brinch-Hansen 80% and 90% failure criterions ..................................................... 57 4.4.6 De Beer failure criterion ......................................................................................... 58 4.4.7 Decourt’s extrapolation criterion ............................................................................ 59 4.5 Static axial compression test on test pile ........................................................................ 59 vii 4.5.1 Test assembly .......................................................................................................... 59 4.5.2 Hydraulic jack and calibration ................................................................................ 62 4.5.3 Static axial compression test ................................................................................... 64 4.6 Interpretation of test pile load movement curve ............................................................. 68 4.6.1 Davisson offset limit load ....................................................................................... 68 4.6.2 10%, 8% and 5% failure loads ................................................................................ 70 4.6.3 L1-L2 failure load ................................................................................................... 72 4.6.4 Chin failure load ..................................................................................................... 74 4.6.5 Brinch Hansen failure loads .................................................................................... 75 4.6.6 De Beer failure load ................................................................................................ 76 4.6.7 Decourt’s failure load ............................................................................................. 78 Chapter 5 Estimation of ultimate loads from torque method and bearing equations ...... 80 5.1 Ultimate loads estimated from the torque method and the corresponding settlements .. 81 5.2 Torque method and compression test interpretation criterions ...................................... 86 5.3 Estimation of ultimate loads using bearing equations .................................................... 90 5.4 Torque method and bearing equations ........................................................................... 92 viii 5.5 Numerical modelling of RS2875.203............................................................................. 96 Chapter 6 Conclusions .......................................................................................................... 100 References .................................................................................................................................. 103 ix List of tables Table 1:Bearing capacity factors for piles .................................................................................... 22 Table 2:Settlements at ultimate and allowable loads corresponding to the torque averaged over last 2.5ft..................................................................................................... 84 Table 3: Calculated ultimate loads of helical piles at 7ft .............................................................. 91 Table 4: Bearing factor from A.B CHANCE & Hubbell Inc at 7ft .............................................. 92 Table 5: Ultimate loads of installed helical piles calculated from torque method at 7ft depth ........................................................................................................................... 92 Table 6 : Ratios between ultimate loads calculated from bearing equations and torque method .............................................................................................................. 94 Table 7: Ratios for pile no 4 ......................................................................................................... 95 Table 8: Ratios between ultimate load calculated from bearing equations and average ultimate load of helical piles from the torque method ........................................... 95 Table 9: Simulated results of RS2875.203 helical piles using HelixPile ..................................... 99 x List of figures Figure 1. Typical single helix helical pile ....................................................................................... 1 Figure 2. Failure plane of a typical footing. .................................................................................. 20 Figure 3. Nq values for circular deep foundations. ....................................................................... 23 Figure 4. Meyerhof Bearing capacity factors and critical depth ................................................... 26 Figure 5. Installation torques of installed helical piles ................................................................. 36 Figure 6. Liquid limit test apparatus ............................................................................................. 37 Figure 7. Particle size distribution curve of airport soil ................................................................ 39 Figure 8. Shearing of cylindrical sample ...................................................................................... 40 Figure 9. Shear and normal stress of airport soil .......................................................................... 41 Figure 10. Compaction curve for airport soil ................................................................................ 43 Figure 11. Standard loading assembly .......................................................................................... 48 Figure 12. Regions in load movement curve ................................................................................ 54 Figure 13. Loading assembly for compression test conducted ..................................................... 61 Figure 14. Mounted dial gauges on reference beams ................................................................... 62 xi Figure 15. Hydraulic jack system ................................................................................................. 63 Figure 16. Calibration chart of AME ram ..................................................................................... 64 Figure 17. Load movement curve of the test pile ......................................................................... 67 Figure 18. Davisson offset limit load ............................................................................................ 69 Figure 19. 8% failure load ............................................................................................................ 72 Figure 20. L1-L2 failure plot ........................................................................................................ 73 Figure 21. Chin failure plot ........................................................................................................... 74 Figure 22. Brinch-Hansen failure plot .......................................................................................... 76 Figure 23. De Beer failure plot ..................................................................................................... 77 Figure 24. Decourt’s extrapolation ............................................................................................... 78 Figure 25. Ultimate loads calculated from torque method .......................................................... 83 Figure 26. Comparison of calculated loads for torque averaged over last 2.5ft and failure loads for KT =10/ft ................................................................................................... 88 Figure 27. Comparison of calculated loads for torque averaged over final 2.5ft and failure loads for KT=9/ft ...................................................................................................... 90 Figure 28. Comparison of loads computed from torque method and bearing equations .............. 94 Figure 29. Model of single helix helical pile ................................................................................ 98 xii List of abbreviation and notations A Cross sectional area of footing As Shaft area Ah Helix area B Width of footing C Cohesion CPT Cone penetration test D Average helix diameter D1 Diameter of uppermost helix fs Unit skin resistance H Depth of uppermost helix or installation depth Heff Effective depth of pile KT Torque coefficient Nc, Nq, Ny, Nσ Bearing capacity factors Qb Tip resistance capacity QP Ultimate end bearing load Qs Skin resistance capacity QT Total ultimate capacity qb Unit tip resistance QUlt Ultimate capacity of pile xiii q Overburden pressure S spacing between two helixes SPT Standard proctor test T Installation torque KP Passive earth pressure coefficient y Unit weight of soil ϕ Angle of internal friction of soil σ0 mean normal ground effective stress at pile toe xiv Acknowledgements This research has been completed in collaboration with Mitacs Accelerate program and Team Foundations. The support of Mitacs Accelerate program is commendable who provided sufficient funds to conduct the testing. I would like to thank Team foundation for their technical support in conducting the tests and installation of helical piles. I would like to acknowledge my parents for their motivation whenever I needed. They stood besides me whenever I was going through hard time thousands of kilometers away from my homeland. I would like to acknowledge my supervisor Dr. Sumi Siddiqua who demonstrated continued support to me. Her sympathetic behaviour towards me provided me with a stress less environment for research. Her knowledge helped me unfold mysteries of research. I would like to thank my committee members who provided valuable suggestions and lauded my efforts to keep me motivated. I would particularly like to mention the unconditional support of Jonah Schwab who helped me pointing out the editing mistakes of thesis. 1 Chapter 1 Introduction Helical piles are a type of deep foundations that are also referred to as helical piers or screw piles or anchors. They have been used in construction industry for many years for a range of applications to resist axial compression, axial tension or lateral loads. The helical piles consist of several sections of a central pile shaft which can be circular or square and have varying dimensions. These are attached to each other by means of welds, bolts or screwing threads as the pile is placed in the ground. This shaft has either single or multiple helixes, welded at specified spacing to make the whole pile as seen in figure (1). Figure 1. Typical single helix helical pile (Reprinted from screwpiles, in Wikipedia. n.d., Retrieved May1, 2016, from https://en.wikipedia.org/wiki/Screwpiles#/media/File:Screwpilediagram.gif. Copyright 2007 by Steve Lewenhoff. Reprinted with permission) These helixes can have constant or varying diameters but all must maintain the same pitch. Same pitch lengths make installation much easier and minimize the soil disturbance. 2 Placement of helical piles is done using hydraulic rotary heads that are commonly mounted on truck rigs, track hoes and an array of other equipment, to apply a turning motion to the pile. The combination of the rotary motion and the weight applied on pile drives helical piles into ground. This advancement should be one pitch of a helix per revolution and every helix should have the same pitch. The spacing of the helixes is also very important; spacing should be such that each helix follow the same path. All of these considerations help to minimize the disturbance of the soil as the helical pile is being advanced (Tappenden, 2007). There are many advantages of helical piles over other conventional deep foundations. The most evident benefit is the cost savings due to the rapid and simple installation, since the piling process only requires two-person team and is very fast. Normally, installation of a helical pile takes about 20 minutes depending upon the penetration depth and soil conditions. Their installation also has the benefits of having little noise or vibration, minimal disruption to the surface and can be completed using light weight equipment. Another major benefit is that piles can be loaded to its full capacity instantly after placement and can be removed and re-used (Tappenden, 2007). They can also be used under unique conditions such as high ground water table because they do not require dewatering for installation, which helps to reduce the time and cost of the installation process (Sakr, 2011). They are considered to perform better than other steel piles, and can be estimated to provide loading capacity 3 to 5 times higher than same sized conventional steel deep foundations (Sakr, 2011). However, helical plates can be damaged during the installation process in soils that contain gravel and that are hard. Bearing capacity/ultimate loads of any deep foundation can be determined using various techniques which include direct methods, indirect methods and pile load tests. Direct methods 3 provide bearing capacity of footings directly from in situ testing. These methods include standard penetration test and cone penetration test. Indirect methods are empirical correlations that use different parameters such as soil strength characteristics and dynamic load settlement data to compute bearing capacity of footings. But for helical piles, a different indirect method called “torque method” can also be used. Today, most widely used method to estimate the load bearing capacity of these piles is torque method. In this methods, load capacity is calculated by multiplying the installation torque of helical piles by a suitable torque coefficient which depends on shape and size of shaft (Society, 2006). This empirical method only considers the installation torque for load capacity estimation and ignores soil characteristics and settlements at calculated load capacities. Helical piles like any other deep foundation are structurally designed to be strong enough to carry anticipated loads. Their structural load capacity is significantly higher than the anticipated loads on them. Also, soil strength is always less than the steel which helical pile is made of. Thus, failure of helical piles is governed by the geotechnical strength of soil and pile sustain significant settlements when geotechnical failure occurs. Hence, the bearing resistance of deep foundations is governed by the amount of settlements in pile head. As discussed earlier that torque method does not give settlement response of helical piles, the settlements at calculated load bearing capacities might be significantly higher than the allowable settlements rendering its results invalid. In order to have reliable results which agree well with settlements in pile head and soil strength, empirical torque method need to be studied further. Pile load tests give a complete behavior of footing in terms of settlements at any load. These settlements are plotted against applied loads which is referred as load movement curve. 4 The settlement at any load can be determined using this curve and ultimate loads can be fixed. There are various criterions that tend to interpret this load movement curve. 1.1 Objectives Helical piles are primarily used for housing and small industrial purposes. Geotechnical investigations, in situ direct test methods and pile loading tests are costly and are not feasible for such small projects. To save time and money, torque method is mostly used to estimate loading capacity of helical piles. Torque method do not anticipate settlements in pile. In practice, there is a significant likelihood that the settlements at calculated loads will be more than what is permissible, and these settlements will go unidentified by the torque method. As failure of deep foundations is governed by excessive settlements, torque method need to be investigated primarily regarding settlements at calculated loads. Therefore, following objectives are studied in this research: 1. Identify and assess a suitable torque coefficient for RS2875.203 helical piles 2. Investigate torque method by incorporating settlements at calculated loads 3. Explore and understand various interpretation criterions of load movement curve for RS2875.203 helical piles 4. Compare ultimate loads calculated from torque method and bearing equation 1.2 Test program In total, seven helical piles were installed at Kelowna international airport and their installation torque was noted. Each installed piles were RS2875.203 type helical pile. These piles have single 5 helix of 10inch diameter. The shaft dimeter is 2-7/8 inches. Based on these torque values, ultimate loads on each pier was calculated. On seventh pile, static axial compression test was conducted according to recommendations of ASTM D1143. Interpretation of this compression test is done based on several criterions and resultant ultimate loads are compared with those calculated from torque method. Additionally, soil samples were collected from the testing site and were analyzed in laboratory to get its strength characteristics. These characteristics have been used in bearing equations to anticipate ultimate loads of each helical pier. 6 Chapter 2 Literature review This chapter provide relevant background information about helical piers/piles which is necessary to understand the tests conducted and further analysis. As discussed above, helical piles can have single and multiple helixes welded around a single shaft. There are various shafts sizes used to fabricate the helical piles. The helix sizes can vary from 10 to 14 inches. Normally, there are maximum three helixes on a single pile. Although helical piles are a subgroup of deep foundations, but they can be categorized further into shallow, deep and transition helical piles based on their embedment ratio. Narasimha Rao, Prasad, & Veeresh (1993) categorized helical piles in shallow, transition and deep helical piles based on their embedment depth. Relative embedment is ratio between the depth to the uppermost helix (H) divided by the diameter of the uppermost helix (D1). A pile with a relative embedment of less than or equal 2 is considered a shallow pile. Piles with H/D1 from 2- 4 are categorized as transition piles. The piles having embedment ratio equal to or greater than 4 can be considered as deep piles. In cohesionless soils, the critical embedment ratio after which deep failure conditions prevail depend on soil friction angle (Meyerhof & Adams, 1968). For a relative embedment of less than 5 it is considered to be shallow failure conditions and greater that 5 is considered deep failure conditions (Meyerhof & Adams, 1968). The load carrying capacity of helical piles depend on the failure modes. The total load carrying capacity of a helical pile is dependent on the sum of the resistance from the end bearing of 7 helixes, friction incurred from the shaft of the pile and the cylindrical shear plane resistance due to the helixes (Zhang, 1999) as shown in equation 1. 𝑄𝑇𝑜𝑡𝑎𝑙 = 𝑄𝐵𝑒𝑎𝑟𝑖𝑛𝑔 + 𝑄𝑆ℎ𝑎𝑓𝑡 + 𝑄𝐻𝑒𝑙𝑖𝑥 (1) 2.1 Failure modes for helical piles The combination of these three terms given in equation (1) depends on the failure mode of the soil. The failure modes of helical piles have been split in two different types, cylindrical shear model and the individual plate bearing model. For single helix pile, only individual plate bearing model is developed. Whereas, for multihelix helical pile, both failure models can develop. Type of failure mode that will occur in multihelix pile depends on its inter-helix spacing ratio (S/D). The inter-helix spacing ratio is the spacing between two helixes on a helical piles shaft (S) divided by the average helix diameter (D) on a multi helical pile. The failure modes are further discussed below. 2.1.1 Individual plate bearing model Individual plate bearing failure is a helical pile mode of failure model. In this model, each helix acts as an anchor independently of the others, and failure of the soil occurs above and below the helix plates in tension and compression respectively. In this type of failure, the total bearing capacity of the helical pile is the sum of the individual bearing capacities of each helix. This failure mode is applicable for all single helix helical piles. The multihelix helical piles that do not 8 develop cylindrical shear plane (discussed below in section 2.2.2), their helix act individually and this failure model becomes dominant. 2.1.2 Cylindrical shear plane bearing model In multihelix helical piles, another phenomenon called cylindrical shear plane is developed which provides with additional load carrying potential. The cylindrical shear plane is a plane of resistance developed between the cylindrical volume of soil contained within the circumference of two adjacent helixes and soil outside this contained area. This plane is only developed with specific inter-helix spacing ratio. There is contradiction over what values of this inter-helix spacing ratio should be in order to have cylindrical shear plane resistance in helical piles. Mitsch & Clemence (1985) explained the formation of cylindrical shear plane and suggested that installation disturbances increase the possibility of this plane’s formation. During installation of helical piles, soil surrounding the pile is stressed and thus gets densified. At the same time soil is also sheared by the helixes rending it weaker around helixes. However, soil outside the circumference of helixes remains strong due to induced lateral stresses. This difference of soil’s strength causes the formation of cylindrical shear plane. Vesic (1971) also attributed formation of this plane to weaker soil zones around pile created by installation disturbances. Tappenden & Sego (2007) proposed that cylindrical shear plane is applicable when inter helix spacing ratio is less than 3. Whereas, Narasimha Rao & Prasad (1993) suggested that S/D must be less than or equal to 1.5 in order for the cylindrical shear plane to be formed. Their results were based on tested helical piles in clays for uplift loads. They tested the helical piles 9 with S/D ≤ 1.5 and computed the load carrying capacities taking into account the formation of cylindrical shear plane. Their calculated loading capacities were in a good agreement with those found from the field tests and lead to the conclusion that at inter helix spacing ratio of more than 1.5, the cylindrical shear resistance plane cease to exist. Similarly, Narasimha Rao, Prasad, & Dinakara Shetty (1991) indicated that for helical piles with spacing ration of more than 1.5 installed in cohesionless soils, load carrying capacity estimated from individual plate mode shows a good agreement with experimental capacities. Their pull out test showed that for spacing ratio of 1-1.5, the failure surface was very close to cylindrical shear model. Based on their results, Narasimha Rao & Prasad (1993) provided a factor (SF) which could be used to calculate actual load capacities of helical piles and is given in equation (2), (3) and (4) for various ranges of S/D. Actual load capacities could be estimated by multiplying SF with computed load capacities (5). SF =1.0 for S / D ≤ 1.5 (2) SF =0.863+ 0.069(3.5 − S / D) for 1.5 ≤ S / D ≤ 3.5 (3) SF =0.700+ 0.148(4.6 − S / D) for 3.5 ≤ S / D ≤ 4.6 (4) 𝑆𝐹 =𝐴𝑐𝑡𝑢𝑎𝑙 𝑜𝑟 𝑒𝑥𝑝𝑒𝑟𝑖𝑚𝑒𝑛𝑡𝑎𝑙 𝑐𝑎𝑝𝑎𝑐𝑖𝑡𝑦𝐶𝑜𝑚𝑝𝑢𝑡𝑒𝑑 𝑐𝑎𝑝𝑎𝑐𝑖𝑡𝑦 (5) 10 2.2 Design of helical piles Helical piles and conventional piles are normally designed to take two types of load, vertical lift and compression. To be able to design for the anticipated loads, type of pile needs to be taken into consideration but the most important factor in the design is the properties of the soil in which it is to be placed. When a pile fails, it is not the pile material that is failing but is the soil that the pile is supported by. Experimentation of Sprince & Pakrastinsh (2010) on helical piles in different soils reinforced the significance of soil dependence. They observed differential increase in capacity of same sized helixes in different soils. So the most important part to the design of piles is the properties of the soil on site. Piles can be designed using two approaches: direct method and indirect method. Direct method calculates the soil unit resistance directly based on soil strength characteristics idealized from in situ testing. Indirect method uses some empirical approaches for this purpose. 2.2.1 Direct methods Direct methods include standard penetration test (SPT) and cone penetration test (CPT). SPT works by driving a thick walled tube into a borehole. This tube is driven by dropping a weight of 140 lb on it. The number of dropping blows is counted per six inch of penetration and used to estimate the unit end resistance of soil. The ultimate load carrying capacity can be determined by multiplying end resistance of soil by area of pile at toe. For helical piles, projected area of helixes can be used for this purpose. 11 Cone penetration test involve penetration of a cone into the soil. The resistance of cone penetration is recorded using a sensor at the tip of the cone, a pour water pressure sensor directly above it or located on the cone, and a friction sleeve following the pour water sensor. Using these three sensors the CPT test can give a good approximation of the soil properties. An extension of CPT is used for piles which is LCPC method. This stands for Laboratoire Central des Ponts et Chaussees in Paris and is based on 197 static loading tests that were conducted on different piles (Tappenden, 2007). LCPC method can directly measure the ultimate end bearing and skin resistance of piles. This method was documented by Bustamante and Gianeselli in 1982 and use factors related soil and pile type (Rodrigo & Junhwan, 1999). Tappenden & Sego (2007) argued that LCPC utilizes the coefficients used for cone penetration test to estimate resistances of pile. These coefficients used are derived from soil type and tip resistance of cone penetration test (Tappenden & Sego, 2007). However, it is not clear if this method is able to estimate resistance of cylindrical shear plane in case of helical piles. The basic relationship for LCPC is given in equation (6) and (7) after Rodrigo & Junhwan (1999). 𝑞𝑏 = 𝐾𝐶𝑞𝑐𝑎 (6) 𝑞𝑠 =1𝐾𝑠𝑞𝑐 (7) 12 Where qb = unit base or end resistance Kc= base resistance factor qca = equivalent cone resistance at pile toe Ks =shaft resistance factor qc = representative cone resistance factor for corresponding layer qs = unit shaft resistance 2.2.2 Indirect method Indirect methods estimate unit resistance/strength of soil using empirical relations. These relationships include t Torque method Bearing equations Dynamic load and settlement formulas Torque method is the most widely used method in industry because of its instant results. It is only used for helical piles. Bearing/load capacity is determined by multiplying installation torque with a suitable torque coefficient. On the other hand, bearing equations calculate this capacity by means of soil strength parameters such as cohesion and angle of internal friction. They use few factors which include bearing, shape and inclination factors. These equations can 13 be used for both helical and conventional piles and require laboratory analysis of soil. The third indirect method, Dynamic load settlement formulas, relate bearing capacity of foundations with instant settlement that are incurred in them as a result of dropping certain amount of load. These formulas do not need any lab or field analysis. Each of these methods are discussed here. 2.3 Torque method The torque method is widely known around the industry for its simplicity and zero setup effort. The torque method is able to determine the pile load capacity based on installation torque of the helical pile. This method has two basic advantages; it does not require any soil exploration data and ultimate load capacity of helical piles can be determined immediately after their installation. This method was established based on uplift load tests on helical piles (CFEM, 2006), but is equally used for axial compressive loads in industry and for research. Given the fact that deeper soil layers are denser and stronger than upper layers, the uplift capacity is always less than the compressive capacity. (Trofimenkov & Mariupolskii, 1965) tested two hundred helical piles and found that capacity in compression was 1.4 to 1.5 times the capacity in uplift. Similarly, Sakr (2009) conducted a helical pile test program and found that the compression capacities were 40% to 50% more than the uplift capacities . Researchers such as Livneh & El Naggar (2008) and Narasimha Rao et al. (1991) have studied the behaviour of helical piles in tension and compression and found that the load carrying capacity was higher in compression. To calculate ultimate resistance, the installation torque is multiplied with an appropriate torque coefficient. The basic relationship is shown in equation (8) below. 14 𝑄𝑈𝑙𝑡 = 𝐾𝑇 × 𝑇 (8) where QUlt = Ultimate load on pile KT = Torque coefficient T = Installation torque Canadian Foundation Engineering manual (CFEM, 2006) suggests that the torque coefficient is a function of the skin resistance along the length of the shaft, the skin resistance of the top and bottom of the helixes and the resistance of the leading edge of helixes. This coefficient depends on pier’s shaft size and shape. The recommended values of torque coefficient according to CFEM (2006) for use with equation (8) are as follows: 7/ft for pipe/round shafts of 90mm outer diameter, with this value decreasing to 3/ft for shaft diameters approaching 200mm. 10/ft (33/m) for square shafts of less than 90mm diameter, Installation torque can vary depending on the skin resistance of the shaft during installation. Different shaft types of the same size (i.e. round vs. square) may vary slightly in skin friction values. However, for small sized shafts which produce little or no skin resistance, the installation torque is completely dependent on the skin resistance of the helix, and the computed bearing capacity will be the same regardless of the type of shaft. The minimum shaft 15 diameter for which skin friction is applicable is 100mm (CFEM, 2006). The shaft diameter of installed helical piles (RS2875.203) is 2-7/8 inches (or 73 mm). Hence, a torque coefficient of 10/ft as suggested by CFEM (2006) for square shafts of less than 90mm can be used for RS2875.203 helical piles. Hoyt & Clemence (1989) analyzed 91 multihelix piles in uplift and obtained good approximation of loads using KT as 10/ft for square and round shafts having less than 3.5 inches (89mm) diameter. Furthermore, the same value of KT was used by Zhang (1999) for calculating the uplift capacity of helical piles having similar sized shafts. Although, this is a reasonable choice of torque coefficient for helical piles having 2-7/8 inch outside shaft diameter, 9/ft is a value that is also suggested for this shaft sizes. A.B CHANCE INC, a renowned company that specializes in helical piles, and they recommend to use 9/ft for RS2875.203 helical piles (A.B.CHANCE & Hubbell, 2014). Likewise, section 13.2.1 of AC 358 (ICC evaluation service, 2007) and the international building code 2009 (Willis & Ram Jack., n.d.) advocate same value of torque coefficient for helical piles having round shaft diameters equal to 2-7/8 inches. Changes in the KT can vary the resultant load capacities, but the values of installation torque used in torque method can alter the calculated ultimate load drastically. The Canadian Foundation Engineering manual (2006) advises to use a value of installation torque averaged over the entire depth of the pier, whereas, Hoyt & Clemence (1989) recommend using installation torque values averaged over the final length of penetration equal to three times the diameter of largest helix. Their research was based on analysis of 91 multihelix helical piles tested in uplift. 16 2.4 Bearing equations As discussed before, piles develop their full load carrying capacity from end bearing and skin resistance. In helical piles, a third component, known as the cylindrical shear plane, is added. Bearing equations are available for each of these components of bearing capacity. The basic form of the bearing equation for ultimate capacity of a pile is given in equation (9). 𝑄𝑈𝑙𝑡 = 𝑄𝑏 + 𝑄𝑠 = 𝑞𝑏𝐴𝑏 + 𝑓𝑠𝐴𝑠 (9) where QUlt = ultimate capacity Qb = tip resistance capacity Qs = skin resistance/friction qb = unit baring capacity of the soil Ab = area of pile at base/toe fs = unit skin friction As = shaft area This equation includes end/tip resistance and shaft resistance. The basic form of the bearing equation used to calculate tip resistance is given in equation (10) as proposed by Terzagi in 1943 (Das, 1999). This correlation gives the ultimate end bearing load of footing. The maximum allowable load is calculated by multiplying ultimate load with a suitable safety factor. 17 𝑄𝑈𝑙𝑡 = 𝐴(𝐶𝑁𝑐 + 𝑞′𝑁𝑞 + 0.5Ƴ𝐵𝑁𝑦) (10) where C= cohesion A= cross sectional area of footing Ƴ= unit weight of soil qʹ= overburden pressure NC, Nq, Ny= bearing capacity factors Overburden pressure is calculated by multiplying the unit weight of soil with the depth of footing. Unit weight of soil used depends on the water table depth below natural surface level. If the water table is located above the footing, the soil’s submerged unit weight must be used in the bearing equation. This can be calculated by subtracting unit weight of soil with unit weight of water. As depicted, this relationship estimates ultimate load capacity using three terms. Understanding these terms is vital when explaining the calculated results. Bowles (1988) provides an explanation of these terms. The first term, CNC, dominates in the case of cohesive soils, while the second term, qʹNq. is dominates when the soil is cohesionless. He describes the third term, 0.5yBNy, as the width of foundation’s contribution to bearing capacity. Additionally, cohesion can be replaced by undrained shear strength when clay is present. This is due to the low hydraulic conductivity of clays (Coduto, 2001). 18 The basic bearing equation for shallow and deep foundations is the same; however, they differ in regards to which bearing factors will be used. In the case of shallow foundations, separate additional factors such as shape, depth, and inclination are used. The bearing equations for shallow and deep foundations are explained below. 2.4.1 Load bearing equation for shallow foundations Shallow foundations differ from deep foundations primarily because of their shallower depth, but they can also feature different shapes such as strip, square, and circular. This basic form of the correlation given in equation (10) is used for strip footings. For circular and square foundations, additional factors known as shape factors are used. With the inclusion of shape factors, equation (10) can be written as shown in (11) and (12). 𝑄𝑢𝑙𝑡 = 𝐴(1.3𝐶𝑁𝑐 + 𝑞′𝑁𝑞 + 0.4Ƴ𝐵𝑁𝑦) (𝑠𝑞𝑢𝑎𝑟𝑒) (11) 𝑄𝑢𝑙𝑡 = 𝐴(1.3𝐶𝑁𝑐 + 𝑞′𝑁𝑞 + 0.3Ƴ𝐵𝑁𝑦) (𝑐𝑖𝑟𝑐𝑢𝑙𝑎𝑟) (12) The bearing capacity factors for shallow foundations given by Terzaghi are as follows in equation (13), (14) and (15) (Das, 1999). 𝑁𝑐 = 𝐶𝑜𝑡𝜙 (𝑒2(3𝜋4 −𝜙2)𝑡𝑎𝑛𝜙2𝐶𝑜𝑠2 (𝜋4 −𝜙2)− 1) = 𝐶𝑜𝑡𝜙(𝑁𝑞 − 1) (13) 19 𝑁𝑞 =𝑒2(3𝜋4 −ϕ2)𝑡𝑎𝑛ϕ2𝐶𝑜𝑠2(45 +ϕ2) (14) 𝑁𝑦 = 0.5 (𝐾𝑝𝐶𝑜𝑠2𝜙− 1) 𝑡𝑎𝑛𝜙 (15) where Kp is passive earth pressure coefficient These bearing capacity factors correspond to the soil failure plane angle, α, which is equal to friction angle (𝜙) . However, studies have shown that α is closer to 45+𝜙/2 rather than 𝜙 as shown in figure (2). Therefore, the values of bearing capacity factors were modified and applied in equation (16), (17) and (18) (Das, 1999). 𝑁𝑐 = (𝑁𝑞 − 1)𝐶𝑜𝑡𝜙 (16) 𝑁𝑞 = 𝑒𝜋𝑡𝑎𝑛𝜙𝑡𝑎𝑛2(45 +𝜙2) (17) 𝑁𝑦 = 2(𝑁𝑞 + 1)𝑡𝑎𝑛𝜙 (𝑉𝑒𝑠𝑖𝑐 1973) (18) 20 Figure 2. Failure plane of a typical footing. (Reprinted from Principles of foundation engineering (P. 157) by Das, 1999. Copyright 1999 by Cengage learning, Inc. Reprinted with permission) Bearing capacity factor, Ny, shown in equation (18) was initially presented by Vesic in 1973 (Das, 1999). Nowadays, there are several correlations used for Ny. Meyerhof in 1961 gave a new relationship (Meyerhof, 1963) shown in equation (19). Another correlation for Ny proposed by Davis and Booker (1971) is presented in equation (20) (CFEM, 2006). This relationship is suitable when angle of internal friction is more than 10 degrees (CFEM, 2006). 𝑁𝑦 = (𝑁𝑞 − 1)1.4𝜙 (19) 𝑁𝑦 ≅ 0.0663𝑒0.1623𝜙 (20) The bearing capacity equations mentioned above do not consider the shearing resistance above the bottom of footing along failure plane. To address this, Meyerhof (1963) introduced additional factors in order to account for the deficiencies of these equations (Das, 1999). These factors 21 included shape, depth and load eccentricity factors. The modified equation by Meyerhof can be written as given in equation (21). 𝑄𝑢𝑙𝑡 = 𝐴(𝐶𝑁𝑐 𝐹𝑐𝑠 𝐹𝑐𝑑 𝐹𝑐𝑖 + 𝑞′𝑁𝑞 𝐹𝑞𝑠 𝐹𝑞𝑑 𝐹𝑞𝑖 + 0.5Ƴ𝐵𝑁𝑦 𝐹𝑦𝑠 𝐹𝑦𝑑 𝐹𝑦𝑖) (21) where Fcs, Fqs, Fys= shape factors Fcd, Fqd, Fyd= depth factors Fci, Fqi, Fyi= load inclination factors If all these new factors are combined within thethe bearing factors, then equation (21) can be written as shown in equation (22) below. These bearing capacity factors: 𝑁𝑐∗, 𝑁𝑞∗ and 𝑁Ƴ∗ include the depth, shape and inclination factors. 𝑄𝑈𝑙𝑡 = 𝐴(𝐶𝑁𝑐∗ + 𝑞ʹ𝑁𝑞∗ + 0.5Ƴ𝐵𝑁Ƴ∗) (22) 2.4.2 Load bearing equations for deep foundations Deep foundation piles develop their load bearing capacity from two primary mechanics: toe resistance and skin resistance. The skin (or shaft) resistance of a pile is not considered if the shaft is less than 100mm in diameter (CFEM, 2006). The basic form of bearing equations for deep foundations is the same as for shallow foundations, except for the bearing factors. Unlike shallow foundations, bearing factors for deep foundations are complex and difficult to calculate. This is due to the variation in soil angle of internal friction, which is caused when the pile is 22 driven into the soil (Day, 2010), as well as the differences in soil-footing interaction. When piles are driven into sandy soils, they displace and densify the soil at the bottom of the pile. This action tends to increase the friction angle of the soil around the vicinity of the pile; therefore changing the bearing capacity factors, especially Nq (Day, 2010). The bearing factors for deep foundations/piles are explained below. 2.4.2.1 Bearing factors for pile foundations As discussed above, the bearing capacity factors for deep foundations are relatively complex. Various values for these factors can be found in literature. Abdel-rahim, Taha, & Mohamed (2013) use an NC value of 9 for deep foundations. This approximation of NC is also in use for helical piles (Thompson, n.d. ; A.B.CHANCE & Hubbell, 2014). However, a design engineer can override this value based on the above mentioned correlations for NC (Thompson, n.d.). Canadian foundation engineering manual (2006) recommends NC values for different pile toe diameters which are given in table (1). Table 1:Bearing capacity factors for piles Pile toe diameter (m) Nc less than 0.5 9 0.5-1 7 larger than 1 6 Winterkorn & Fang (1975) gave the following correlation for NC. 23 𝑁𝑐 = (𝑁𝑞 − 1)𝑐𝑜𝑡𝜙 (1) (23) The bearing capacity factor, Nq, depends on soil friction angle and the failure mechanism at the pile toe. The bearing factors are generally much higher for deep foundations as compared to shallow footings (Lambe & Whitman, 1969). The possible cause of this is the variations in friction angle of soil due to the soil consolidation around pile during pile installation as discussed above. Vesic (1967) gave the values of Nq for circular deep foundations/piles by several researchers which has been given in figure (3). Figure 3. Nq values for circular deep foundations. (Reprinted from “Ultimate loads and settlements of deep foundations in sand,” by Aleksandar. S. Vesic, 1967. Copyright 1967 School of engineering, Duke university) 24 For deep foundations, the width (toe diameter) is very small compared to the depth of foundation. For helical piles it is even smaller. Such small diameters result into very small values of 0.5ƳD𝑁Ƴ and therefore can be neglected. If B is less than 9.8 to 13.1ft, then this term can be ignored (Thompson, n.d. ; Hubbell & CHANCE, 2003). Various researchers such as Meyerhof, Vesic and Janbu have studied bearing equations for piles. Das (1999) explained Meyerhof and Vesic methods which are discussed below. 2.4.2.2 Meyerhof method In 1976, Meyerhof proposed recommendations for calculating the bearing capacity of pile foundations. He argued that the point bearing resistance and skin resistance of piles increases with the depth of pile inside the ground in homogeneous soils. The ultimate end bearing load for homogeneous cohesive soils from this method is given in equation (24) 𝑄𝑝 = 𝐴𝑃(𝐶𝑁𝐶∗ + 𝑞′𝑁𝑞∗) ≤ 𝐴𝑃𝑞1 (24) where Qp= ultimate end bearing load AP= cross-sectional area of pile q/= effective overburden pressure q1 = limiting value of unit point resistance 25 Nq*= bearing capacity factor with respect to overburden pressure NC*= bearing capacity factor for cohesion When using a helical pile, the area of pile in equation (26) is replaced by projected helix area (Ah). q1 is given as 0.5𝑁𝑞∗tan (𝜙), and has units of tons per square foot. Nq increases as embedment ratio increases, and reaches a maximum value when the embedment ratio becomes equal to half of the critical embedment/depth ratio. The critical embedment ratio corresponding to any specific friction angle of soil can be calculated using Meyerhof (1976) plot given in figure (4). Meyerhof states that in most cases, the embedment ratio of deep foundations is greater than half of critical embedment ratio. For this reason, maximum value of 𝑁𝑞∗ and 𝑁𝐶∗ should be used for determination of load carrying capacity. These bearing factors are given in figure (4) below. 26 Figure 4. Meyerhof Bearing capacity factors and critical depth (Reprinted from “Bearing capacity and settlement of pile foundations,” by G.G. Meyerhof, 1976, Journal of the geotechnical engineering division, 102(GT3), p. 197-228. Copyright 1976 American Society of Civil Engineering. Reprinted with permission. (This material may be downloaded for personal use only. Any other use requires prior permission of the American Society of Civil Engineers) 2.4.2.3 Vesic (1977) method Vesic proposed a different method for measuring the ultimate load on piles in 1977 (Das, 1999). The Vesic’s ultimate point/end bearing load on pile is given in equation (25) below. 𝑄𝑝 = 𝐴𝑝(𝐶𝑁𝐶∗ + 𝜎0ʹ𝑁𝜎∗) (25) where 27 σ0=mean normal ground effective stress at pile toe= (1+𝐾𝑜3)𝑞′ 𝑁𝜎∗= bearing capacity factor The bearing factor, 𝑁𝜎∗, is a function of reduced rigidity index. Reduced rigidity index can be computed using the rigidity index which in turn is dependant on friction angle, cohesion and modulus of elasticity of the soil. Das (1999) has provided, the bearing factors used in Vesic’s method for various values of reduced rigidity indexes and friction angles. 2.4.2.4 Load bearing equation for helical piles Helical piles, like conventional piles, derive their full load bearing capacity from two main mechanisms; skin friction and end bearing. Additionally, a third phenomenon known as cylindrical shear plane can develop depending upon inter helix spacing ratio and contribute to load bearing capacity. The summation of these three components yields the total bearing capacity of helical foundations. The end bearing of each helix is calculated unless the helixes are not close enough to form a cylindrical shear plane. In order to calculate total end bearing load capacity (QT), end bearing resistance of all individual helixes (Qh) are summed up as shown in equation (26). The ultimate individual helix bearing capacity can be derived using the same relations as for other deep foundations, because helical piles are a sub type of deep foundations (CFEM, 2006). However, the calculation for helical pile differs as area of pile is replaced by projected helix area, as seen in equation (27) 𝑄𝑇 = ∑ 𝑄ℎ (26) 28 𝑄ℎ = 𝐴ℎ(𝐶𝑁𝑐∗ + 𝑞ʹ𝑁𝑞∗ + 0.5Ƴ𝐷𝑁Ƴ∗) (27) where Ah= projected helix area Although, the bearing factors for deep foundations can be used for helical piles, some researchers have explained bearing factors specifically for helical piles. Tappenden & Sego (2007) argued that for helical piles, the Nq values recommended by Vesic (1963) may be used. Additionally, Nq values can also be calculated based on the work of Meyerhof (1976) (A.B.CHANCE & Hubbell, 2014;Thompson, n.d.). The relationship adopted from Meyerhof’s work is given in equation (28) which is Meyerhof Nq values multiplied by 0.5. This approach is used for long term applications (A.B.CHANCE & Hubbell, 2014;Thompson, n.d.). 𝑁𝑞 = 0.5(12 × 𝜙)𝜙/54 (28) Based on the above discussion, the friction angle of soil has come forward as a basic parameter in order to calculate bearing capacity factors. However, when angle of internal friction is not known, SPT blows can be utilized to approximate a value of internal friction angle. For this purpose the empirical data provided by Bowles in 1968 can be used which is represented by the equation (29) given below (Thompson, n.d.). ∅ = 0.28𝑁 + 27.4 (29) 29 Skin friction in helical piles is a resistance that develops along the shafts of the pile and can also be referred to as shaft friction/resistance. This type of friction follows a relationship with the shaft, where the resistance will increase as the shaft diameter increases. The relationship given in equation (30) can be used to calculate shaft friction (CFEM, 2006). The shaft friction is ignored if the shaft is less than 100mm in diameter (CFEM, 2006). 𝑄𝑠 = ∑[𝜋𝐷𝑓𝑠∆𝐿𝑓] (30) where, QS = shaft capacity D= Diameter of shaft ΔLf = Incremental pile length over which πD and fs are taken as constant fs= Sum of friction and adhesion between soil and pile The length of shaft for which skin friction is valid for is referred to as effective length of shaft (Heff). Effective length is not always equal to entire shaft length, rather, it varies depending upon the category of helical piles in terms of embedment ratio. Narasimha Rao, Prasad, & Veeresh (1993) conducted several tests on model helical piles in uplift and explained the implications of different effective lengths. They observed that the failure zone of shallow helical piles extended to the top of a pier at the surface, and that there is a gap between the shaft and soil at the top of pier. Based on these observations, they argued that the shaft resistance for shallow helical piles should be ignored. Similarly, skin resistance was present for transition piles; however, heaving of soil at the surface was also observed. This indicates that skin resistance is 30 not applicable for the entire shaft length and varies with different embedment ratios. For transition helical piles, effective shaft length was equal to 0.7 D1 - 0.9 D1 and 1.7 D1 - 2.5 D1 for embedment ratio of 3 and 4 respectively. No heaving of soil was observed for deep helical piles and effective length of shaft was in the range of (H - 1.4 D1) to (H - 2.3 D1). (Trofimenkov & Mariupolskii, 1965) analyzed helical piles in compression and tension and found that skin resistance was not applicable for the entire shaft length. They attributed it with the formation of compaction zone and hollow zone above top helix in tensile and compressive loads respectively. The length of shaft above the top helix for which skin resistance is not applicable is equal to the helix diameter (Trofimenkov & Mariupolskii, 1965). Zhang conducted helical pile load tests in 1999 and proposed an effective length of shaft for skin friction as given in equation (31). Where, H is installation depth of pile and D1 is diameter of top helix. 𝐻𝑒𝑓𝑓 = 𝐻 − 𝐷1 (31) 2.5 Pile load tests Pile load tests are the most accurate and precise way to determine the ultimate loads which a pile can carry. Direct and indirect methods of design do not give settlements/deflections in piles, and therefore lack accuracy. As discussed earlier, if a pile settle more than the specified limits, it is considered failed even if the designed allowable load is not reached. For this reason, pile load tests are conducted prior to construction to check for actual ultimate and allowable load. 31 Two types of load tests can be performed: axial tension and axial compression. In axial tensile load tests, tension loads are applied on a pile, whereas, for axial compressive load tests, compressive loads are applied in order to check for ultimate and allowable loads. The main product of a pile load test is to obtain load and settlement data. A graph is plotted between load and settlement which is regarded as load movement curve or loading curve. Using this plot, settlement/deflection of a pile at any load magnitude can be determined. This curve is useful because it describes the complete behaviour of a footing. Depending upon the allowable limits of settlements for any particular structure, a suitable ultimate and allowable load is selected from loading curve. There are several failure criterions which analyze a load movement curve and define ultimate loads. Since pile load tests are done in real field conditions, they represent actual loading capacity of footings for which settlements are maintained within limits. 32 Chapter 3 Installation of helical piles and site investigation North-West corner of Kelowna International airport was selected as the testing site to install helical piles. This site is away from the airport taxiway and is a good plain area with minimum disruptions. Prior to the testing, City of Kelowna was contacted to get the permission for testing on these grounds. Because of the presence of utilities and water pumps in vicinity, it was necessary to check for any underground utilities at the piling location. For this purpose, BC CALL ONE was contacted, and piling started after getting their confirmation. 3.1 Configuration and material properties of helical piles All piles installed and tested were AB CHANCE RS2875.203 helical piles. These single helix piles have shaft diameter of 2.875 inches and thickness of 0.203 inches. The diameter, pitch and thickness of helix is 10 inches, 3 inch and 0.375 inch respectively. A.B.CHANCE & Hubbell (2014) specifies the nominal capacity of RS2875.203 as 73.6 kips and maximum torque rating as 6000 ft-lb. The material used for the manufacturing of the shafts is ASTM A500 grade C material with a minimum yield strength of 50 ksi. The helix material is ASTM A572 with minimum yield strength of 50 ksi. The allowable load capacity of these piles is 66.3 and 36.8 kips based on LRFD and ASD design methods respectively (A.B.CHANCE & Hubbell, 2014). However, the load carrying capacity of deep foundations is governed by the geotechnical strength of soil, as the structural strength of deep foundations is much greater than the soil’s strength. 33 In total, seven helical piles were installed at the test site, and installation torques were measured digitally through a system supported by a wireless network. A Quick static axial compression test was performed on seventh pile which is identified as test pile here after. Test pile was installed in the middle of all the other piles. All piles were driven to a depth where they achieved an installation torques value close to 5500 ft-lb. The distance between piles was carefully selected in order to avoid any interaction between their developed pressure bulb within the ground. Helical pile extensions were used to drive them to the desired depth. The torque records of installed piles are given in figure (5). From these records, it is clear that the increase in torque was not gradual; but instead changed rapidly. (a) 015003000450060000 4 8 12 16 20 24Torque (ft-lb)Depth (ft)Pile No 134 (b) (c) 015003000450060000 4 8 12 16 20 24Torque (ft-lb)Depth(ft)Pile No 2015003000450060000 4 8 12 16 20 24Torque (ft-lb)Depth (ft)Pile No 335 (d) (e) 015003000450060000 4 8 12 16 20 24Torque (ft-lb)Depth(ft)Pile No 4015003000450060000 4 8 12 16 20 24Torque(ft-lb)Depth(ft)Pile No 536 (f) (g) Figure 5. Installation torques of installed helical piles 015003000450060000 4 8 12 16 20 24Torque (ft-lb)Depth(ft)Pile No 6015003000450060000 4 8 12 16 20 24Torque(ft-lb)Depth (ft)Test Pile37 3.2 Soil investigation In order to determine the geotechnical characteristics of the soil in which helical piles were installed, soil was excavated to a depth of 7ft and samples were collected. Various tests such as soil sieve analysis, hydrometer test and direct shear test were conducted to investigate the classification of soil and its strength characteristics. 3.2.1 Soil classification To classify the soil according to unified soil classification system, ASTM D2487 standards were followed. Soil sieve analysis revealed that the soil composition was sand with fines. In order to further classify the soil, liquid and plastic limit tests were performed according to ASTM D4318 standards. The liquid limit test apparatus used is shown in figure (6). Figure 6. Liquid limit test apparatus 38 Due to the significant fluctuations of liquid and plastic limit test results, a hydrometer test was performed to further classify the soil. This test is conducted to determine the particle size distribution of particles less than 0.075mm. As silt consists of particles sized from 0.05-0.002mm and clay particles are less than 0.002mm, the percentage of silt and clay in a soil sample can be identified by analyzing the particle size distribution curve given by a hydrometer test. Hydrometer test was conducted according to recommendations of ASTM D422. And the soil particle size distribution curve is given in figure (7). The results from the hydrometer test and soil sieve analysis are plotted together in figure (7) which explains the complete particle size distribution of the soil at the test site. As underscored by the distribution curve, the percent passing at diameter equal to 0.002mm and 0.038mm was about 3% and 8.17% respectively. Clearly, percentage of soil with particle size equal to 0.002mm was less than the percentage at size greater than 0.002mm; thus, the soil was classified as silty sand (SM). 39 Figure 7. Particle size distribution curve of airport soil 3.2.2 Direct shear test Direct shear test is performed to determine shear strength characteristics of soil (i.e. angle of internal friction and cohesion). For this purpose, the direct shear apparatus developed by GDS instruments was used. Parameters such as sample height, width, shearing rate and normal load on sample are input using the device’s built in computer software. It is able to record and save data every second. Two displacement transducers, vertical and horizontal, are attached to the apparatus which records respective displacements. Two types of soil samples can be sheared using the GDS apparatus, one is square and other is cylindrical. The shearing plane is predefined in case of square samples, whereas there is no control over shearing plane in the case of 0204060801001200.001 0.01 0.1 1 10Percent passingDiameter (mm)40 cylindrical samples. Several steel rings are used to confine the cylindrical soil samples, which ensure the anonymity of shearing plane. Figure 8. Shearing of cylindrical sample Direct shear test was performed according to the recommendations by ASTM D3080, and in total, four tests were conducted with square samples. Each sample was compacted to develop density similar to that in field. Each sample was sheared to a relative lateral displacement equal to 10% of the width of sample as per the recommendations of ASTM D3080, where the relative lateral displacement is defined as the distance between the top and bottom halves of shear box. Figure (9) represents the results of direct shear tests which were conducted. As observed in figure, the results show very good consistency with each other. The cohesion and friction angle was calculated to be 0.0036 ksi and 27 degrees. 41 Figure 9. Shear and normal stress of airport soil 3.2.3 Water contents of soil Water content is calculated by dividing the weight of water in a soil sample by the soil’s dry weight. For this purpose, four samples were taken and dried in an oven for 24 hours. Weights were measured before drying and after drying in order to calculate the water contents in the samples. Average water contents in soil was found to be 7.40%. y = 0.5098x + 0.003600.0040.0080.0120.0160 0.005 0.01 0.015 0.02 0.025Shear Stress (Ksi)Normal Stress (Ksi)42 3.2.4 Unit weight of soil The bearing capacity of a foundation is dependent on the soil unit weight or density in field. To measure the unit weight of soil, two steel moulds were inserted into the ground to collect undisturbed soil samples. The internal diameter and height of moulds were measured using Vernier calliper. Using these measurements, the volume of soil contained inside the moulds was calculated, and weights of the moulds with soil still inside was subtracted with the weights of empty moulds to determine the weight of soil. Unit weight was calculated by dividing weight of soil with the volume of soil. Average unit weight was found to be 14.90 KN/m3. 3.2.5 Maximum dry unit weight and optimum moisture content The maximum unit weight of the soil is dependant on its moisture content. As the moisture content in a soil increases, the unit weight will increase until it attains a maximum value at certain moisture content, which is known as the optimum moisture content. This test was performed based on the recommendations of ASTM D698 (standard proctor test), from which the compaction curve is given in figure (10). The optimum moisture content of soil was found to be 12.5% and maximum dry unit weight was 18.1 KN/m3. 43 Figure 10. Compaction curve for airport soil 16.51717.51818.50 4 8 12 16 20Dry density (KN/m3)Actual moisture content (%)44 Chapter 4 Static axial compression test 4.1 Introduction There are several methods to estimate the load capacity of foundations, such as empirical relations, bearing capacity equations and load tests. Empirical formulas for deep foundations include high strain dynamic loading tests and torque tests method. Dynamic testing is conducted by dropping a load from a variety of different heights to the foundation cap to induce some immediate settlements, which is linked to soil strength and eventually leads to load capacity. Similarly, torque method links installation torque with load carrying capacity and calculates load capacity by multiplying installation torque with a torque coefficient. The bearing capacity equations utilize soil strength characteristics to estimate the ultimate load carrying capacity. All these correlations indirectly calculate load capacity and possess default uncertainties, as there can be instrumental or human errors in data acquisition. The static axial compression test is conducted to estimate the actual load carrying capacity of foundations. In this loading method, a predefined load is applied on foundations in intervals. The resultant settlements are recorded at each load interval. As this test is conducted in field on installed piles, it results in the actual load capacity of foundations by taking real soil conditions into account which is not possible by other method. Other methods such as empirical and bearing equations estimate ultimate load carrying capacities regardless of settlements and sometimes may give results for which settlements are not within allowable range. 45 4.2 Types of static axial compression test The compression test on foundations can be performed in seven different ways. Load movement curve is the key product of each method. ASTM D1143 specifies the following seven types or methods of static axial compression test. i. Quick test ii. Maintained test iii. Loading in excess of maintained test iv. Constant time interval test v. Constant rate of penetration test vi. Constant movement increment test vii. Cyclic loading test i. Quick test: In this method, anticipated failure load is applied on foundations in increments of 5% and settlement for each load interval is measured. The load is applied until the foundation fails but is not exceeded by the safe structural load of foundations. During each increment, the load is kept constant for a time interval of 4 to 15 minutes. After completely loading the pile, the load is removed in 5 to 10 equal decrements keeping each decrement constant for 4 to 15 minutes. Movements (settlements) in pile head are noted down at 0.5, 1, 2, 4, 8 and 15 minutes during each load increment and decrement. ii. Maintained test: In this method, load equal to 200% of design load is applied in increments of 25% and is maintained unless rate of axial movement does not exceed 46 0.25mm per hour. Afterwards, load is removed when settlement do not exceed 0.25mm in one hour iii. Load in excess of maintained test: This test method is an extension of maintained test where the pile is loaded again to the maximum maintained load in increments equal to 50% of pile design load. For this, time interval between the load increments is kept 20 min. Additional load is applied in increments, keeping the same time intervals, equal to 10% of pile design load until the pile fails. If the pile fails, load is applied until the pile attains a settlement equal to 15% of pile diameter. Otherwise, the load is held for two hours and removed in four decrements keeping time interval of 20 minutes. iv. Constant time interval test: The same procedure is followed as for maintained load test. However, the load increments are 20% of design load with one-hour time interval between each increment. The pile is then unloaded with same time interval of one hour between loading decrements. v. Constant rate of penetration test: In this testing method, a magnitude of load is applied such that the pile penetrates with a constant rate equal to 0.25 to 1.25mm per minute for cohesive soils and 0.75 to 2.5mm for granular soils. Maximum applied load is held on pile unless it penetrates at least 15% of average pile diameter or stops penetrating and then load is removed gradually. vi. Constant movement increment test: Load is applied in increments such that the additions in settlement of pile head are equal to 1% of pile diameter. The load is continuously applied until the total settlement reaches 15% of pile diameter. Additional load is not applied until the load variation rate necessary to keep the settlement increments constant 47 is less than 1% of total applied load per hour. Load is then removed in four decrements until rate of load change is less than 1% per hour. vii. Cyclic loading test: Load is applied in 50, 100 and 150% increments of design load in the same manner as in case of maintained load. Each load increment is kept for one hour and load is removed in decrements which are equal to increments. Time between decrements is kept 20 minutes. After removal of each load, pile is then loaded again keeping load increments equal to 50% of design load and time interval of 20 minutes. 4.3 Components and requirements of loading assembly Following are the main components of loading assembly as recommended by ASTM D1143 in order to apply the load on pile or a pile in group. Hydraulic jack/ Weights Test beam Anchor piles/Reaction piles Steel test plate Load transfer beams Reference beams Settlement measuring device After installation of a pile, a steel test plate is placed above the pile head to uniformly distribute loads. The load is applied by means of hydraulic jack which is placed over the steel test plate. The load induced by hydraulic jack is transferred to the soil as tensile load by means of four reaction piles. For safety purposes, load capacity of each reaction pile is kept more than 48 the anticipated failure load of foundation. This loading assembly can be changed depending upon the conditions of sites and availability of different loading components. For example, the wooden planks and dead weights can replace reaction beams and hydraulic jack respectively in case they are not available. A standard loading assembly can be seen in figure (11) where load is applied using hydraulic jack. Reaction piles are fixed with load transfer beams which are placed horizontally normal to the test beam. When hydraulic jack is raised to apply load, the load is transferred from test beam to the load transfer beams, which further distribute load uniformly to four reaction piles. Figure 11. Standard loading assembly (Reprinted from Standard test method for deep foundations under static axial compressive load, ASTM D1143, in ASTM international. Retrieved from http://compass.astm.org/EDIT/html_annot.cgi?D1143+07(2013) . Reprinted with permission) 49 The requirement of each component of this assembly is important for reliable results. The test beam should be strong enough to sustain load and sufficient in size to prevent excessive deflections. The settlements can be measured using displacement transducers and dial gauges. Laser beams and surveyor’s level can also be used to observe settlement and should be referenced with a permanent bench mark. A minimum of two transducers or dial gauges capable of measuring at least 2 inches are required. These gauges are mounted on reference beams and their stems are kept parallel to longitudinal axis of pile or pile group. Reference beams should be levelled horizontally to avoid any errors in settlement measurements. 4.4 Interpretation of static axial compression test Static axial compression test provides load and settlement data. This data can be interpreted to determine the failure/ultimate load of foundations. There are several criterions that attempt to analyse this data and can be used to estimate failure load. These criterions/methods are discussed in the following section. Davisson offset limit load criterion 10%, 8% and 5% failure criterions L1-L2 failure criterion The Chin failure criterion The Brinch-Hansen 80% & 90% failure criterions De Beer failure criterion Decourt’s extrapolation 50 There is no specific method recommended by any standards. However, the International Building Code (2006) recognizes three methods by which load movement curve can be interpreted. These three methods are: 1- Davisson; 2-Brinch Henson, and; Butler-Hoy criterion (NeSmith & Siegel, 2009). The selection of these methods depend on individual cases but Davisson offset limit load and 10% failure criterion are the most widely used methods in the deep foundation analysis. 4.4.1 Davisson offset limit load criterion The Davisson offset limit load considers elastic compression of pile. It is mostly used for driven piles which are tested according to quick method (Fellenius, 1980;Sakr, 2011) It defines failure load as the load corresponding to a settlement which exceeds elastic compression by 0.15 inches plus a factor equal to diameter of pile divided by 120 (Fellenius, 1980). Equation (32) can be used to compute this settlement. S =PLAE+ 0.15 +D120 (32) where S= total settlement P= load A= cross sectional area of pile E= modulus of elasticity of pile material D= diameter of pile at toe in inches 51 The diameter of pile used in equation (32) should be considered as the diameter at the toe of the pile. For helical piles, this diameter should be the helix diameter. If the pile is expanded at the base, then the diameter of pile should be taken as the diameter of expanded base (CFEM, 2006). In order to determine the failure load, elastic compression line is first drawn on load movement curve. The scale of the load movement graph should be such that the elastic compression line makes an angle of 20 degrees with the load axis (Abdelrahman, Shaarawi, & Abouzaid, 2003). Canadian foundation engineering manual (2006) also recommends to draw this line at 20 degrees from the load axis. In the second step, offset line is drawn parallel to elastic compression line at a distance equal to 0.15 inches plus D/120 from elastic compression line. The load corresponding to the intersection point of offset line with the load movement curve is considered as the ultimate load on pile. Allowable or safe load is determined by dividing the ultimate load with a safety factor which is normally equal to 2. The primary advantage of this criterion is that the elastic compression and offset line can be drawn prior to the pile load test and test can be ended when load movement graph touches the offset limit line. Davisson method can also be analyzed to check the presence of skin friction along the pile shaft. The intersection of elastic compression line with the load movement graph indicates the presence of skin friction along the shaft of pile (A.B.CHANCE & Hubbell, 2003). This offset limit load method always yields conservative results (CFEM, 2006). 4.4.2 10%, 8% and 5% failure criterions According to 10% failure criterion, a pile can be considered failed when it settles to an amount equal to 10% of its diameter at toe. This criterion was first proposed by Terzaghi in 1942 for 52 compressive loads (Sakr, 2009). Subsequently, this criterion was proposed by Viggiani, Mandolini, & Russo (2014). Sakr (2011) proposed that full toe resistance of soil is normally mobilized at a settlement equal to 10% of helix diameter. Unlike the toe resistance, shaft resistance is mobilized at small settlements which are equal to 5mm to 10mm (CFEM, 2006). 10% failure criterion, when used for smaller helix plates, yields appropriate settlements. For piles with greater diameter of helix, 10% criterion can result into a failure load for which settlements will be too high. These excessive differential settlements can pose a major challenge for design engineers. For this reasons, further failure criterions have been introduced which could account for larger diameter helixes and are describer below. Livneh & El Naggar (2008) tested 19 helical piles in both tension and compression. They observed that most of the piles failed in compression when a settlement more than 8% of helix diameter had been achieved. They proposed a failure criterion which defines the failure load as a load corresponding to the settlement equal to 8% of largest helix plus elastic compression settlement of helical pile. This failure criterion is given in equation (33) below which is similar to Davisson offset limit load criterion. Failure load is also determined in the same way as for Davisson offset limit load criterion. 𝑆 =𝑃𝐿𝐴𝐸+ 0.08𝐷 (33) where S= Settlement at failure P= Applied load at failure 53 L= Length of pile A= Cross-sectional area of pile shaft E= Young’s modulus of steel D= Diameter of largest helix 5% failure criterion was proposed by Sakr in 2009 and defines the ultimate or failure load as a load corresponding to a settlement equal to 5% of average helix diameter. This criterion is intended for the uplift loads. In most cases, the uplift capacities of helical piles are less than the compressive capacities for the same soil characteristics. The testing of two hundred helical piles by Trofimenkov & Mariupolskii (1965) revealed that the compression load capacities were about 1.4 to 1.5 times the tension or uplift load capacities. Sakr (2009) tested few helical piles and suggested that compression capacities are about 40% to 50% more than the uplift capacity. For this reason, the 5% failure criterion can also be used for the compressive load capacities, as the results will always be safe. O’ Neill & Reese (1999) defined the axial compressive capacity of cast-in-place concrete piles as the load that produces a settlement in pile equal to 5% of the toe diameter. 4.4.3 L1-L2 failure criterion L1-L2 method was proposed by Hirany & Khulhawy in 1989 (Kulhawy & Hirany, 2009). The load movement curve can be divided in three regions as shown in figure (12): initial linear, transition and final linear. In initial linear portion, the graph remains more or less linear. Afterwards, it starts changing its shape which is transition region. Then comes a stage after transition region where a large settlements occur by applying small loads. This portion of graph 54 is called final linear portion. The point corresponding to initial linear region is L1 and point corresponding to final linear region is called L2. The load corresponding to point L1 and L2 are PL1 and PL2 respectively. Figure 12. Regions in load movement curve (Adapted from “Interpreted failure load for drilled shafts via Davisson and L1-L2,” by Fred. H. Kulhawy and A. Hirany, 2009, Proceedings of the 2009 International Foundation Congress and Equipment Expo, p.127-134. Copyright 2009 American Society of Civil Engineers. (This material may be downloaded for personal use only. Any other use requires prior permission of the American Society of Civil Engineers) 55 L1 is easy to point out as it is prominent in most of the cases. However, in absence of large displacement data, it is difficult sometimes to determine the exact location of point L2 on graph (Kulhawy & Hirany, 2009). Khulhawy and Hirany in 2009 analyzed this method with other existing methods and found out that most of the piles fail in transition region or in the final linear region of load movement curve. Their research revealed that the Davisson offset limit load was always higher than PL1 and lower than PL2 and failure occurred in transition zone. However, they argued that failure can also take place beyond final linear portion of graph. Sakr estimated ultimate loads for the helical piles in 2011 using Davisson, L1-L2 and 5% failure criterions. The helical piles that were tested by Sakr (2011) had shaft diameter range between 1234 to 16 inches and helix diameters were between 30 to 40 inches. He found out that capacity results from L1-L2 method were lower than the 5% criterion by 13% to 18%. However, Davisson criterion yielded load capacity values up to 47% lower than L1-L2 and 5%. Apart from being a simple method, this method has some major drawbacks. One drawback is that it can be effected significantly by the scale of graph as shape of load movement curve can change as a result of scale change. The shape of load movement graph can also vary considerably by the amount of skin friction in shafts of piles. This criterion can produce considerably high load capacity values when skin friction is present. The graph of load movement curve starts changing shape once skin friction reaches its threshold value. Thus, this method can produce considerably low capacity values for small diameter piles where skin friction is usually non-existent. Moreover, unlike other criterions such as Davisson, this criterion does not consider elastic compression. 56 4.4.4 Chin failure criterion This failure criterion was proposed by Chin in 1970. This method assumes a hyperbolic function between load and settlement (Abdelrahman et al., 2003). In order to calculate the failure load, each settlement value is divided by its corresponding load value and plotted along the ordinate. The settlements are plotted along the abscissa. This failure criterion works on the assumption that the plot shows a straight line at failure and the inverse slope of this straight line is failure load. The slope of plotted straight line is called C1 and y intercept is called C2. The equation of this straight line is given in equation (34). ∆𝑃= 𝐶1 ∆ + 𝐶2 (34) where, Δ= settlement P= load To identify a straight line, three points are required on a line. However, there are significant chances of arriving at wrong conclusions in locating this straight line. Canadian foundation engineering manual (2006) suggest that Chin failure criterion always results in ultimate load capacities more than the maximum applied load on pile and is therefore is less useful. However, it can be used to underscore the potential damages to the pile which appear as a sudden change in Chin graph (CFEM, 2006). 57 Chin method can be applied to both quick and slow tests but there has to be constant time increments between loads (Fellenius, 1980). This criterion is also dependent on extent of loading. Abdelrahman et al. (2003) monitored this dependency and suggested that if the pile is loaded close to failure loads then predicted ultimate loads are higher. As ASTM D1143 does not allow to load the footing beyond safe load capacity, the applied loads in static axial compression test are lower than the failure loads most of the times. For this reason, sometimes one need to extrapolate the load movement curve in order to estimate the failure loads. The extrapolation creates doubts and can result in wrong failure loads (CFEM, 2006;Abdelrahman et al., 2003). Fellenius (1980) suggested that this method is more effective for high load data points on the load movement curve. Usually a straight line does not form until it passes Davisson offset limit load and as a rule, Chin method results are 20% to 40% greater than Davisson offset limit load (Fellenius, 1980). 4.4.5 Brinch-Hansen 80% and 90% failure criterions Brinch-Hansen 80% criterion was proposed by Brinch & Hansen in 1961. It defines failure load as a load which induces four times the settlements as obtained at 80% of that load (Fellenius, 1980). Failure load is (PU) which induce settlement (ΔU) if 0.80PU gives a settlement value of 0.25ΔU. Similarly, Brinch-Henson 90% criterion defines failure load as a load that gives twice the settlements of pile head as obtained at 90% of that load (Fellenius, 1980). In most cases, 80% criterion is usually followed. To estimate failure load, a graph is plotted between ratio of square root of each deflection and the test load (√Δ/P) on ordinate and deflections on abscissa (Dotson, 2013). This method 58 assumes a straight line on its plot (CFEM, 2006). A best fit linear line is drawn and its slope (C1) and y intercept (C2) is calculated. PU and ΔU are given as follows in equation (35) and (36). 𝑃𝑢 = 12√𝐶1 𝐶2 (35) ∆𝑢 = 𝐶1𝐶2 (36) The load movement data with low load points can be ignored to plot Brinch Hansen curve because data with high load magnitude is more significant for this criterion (Dotson, 2013). Canadian foundation engineering manual (2006) suggest that this criterion is only valid if failure load (plunging failure) was reached during static axial compression testing of pier and the point (0.8PU, 0.25ΔU) lies on load movement curve of test pile. If this is not the case, load movement plot can be extrapolated. 4.4.6 De Beer failure criterion This failure criterion was proposed by De Beer in 1967 and again in 1972 by De Beer and Wallays. This method was initially developed for slow axial load tests (Fellenius, 1980). To calculate failure load, graph between load and settlement is plotted in a double logarithmic scale. Graph plot values falls in two straight lines. The intersection of these lines is defined as failure load. This method assumes that pile was loaded close to failure load, otherwise the graph can be a single straight line (Abdelrahman et al., 2003). 59 4.4.7 Decourt’s extrapolation criterion This criterion was proposed by Luciano Decourts in 1999 (Abdelrahman et al., 2003). Each load value is divided by its corresponding settlement and plotted against the load. Plot tend to be a straight line. The point where this straight line intersects load axis is regarded as ultimate/failure load. Ultimate load is calculated by extrapolating straight line to the load axis. The slope of the line is C1 and y intercept is C2. The ultimate load can directly be determined by dividing C2 with C1 (equation 37). 𝑄𝑢 = 𝐶2/𝐶1 (37) 4.5 Static axial compression test on test pile The helical pile on which this compression test was conducted is denoted as test pile. Test pile was the same as other installed helical piles i.e. RS2875.203. It was installed at the middle of the rest piles so as to keep it in same soil conditions as for other’s. 4.5.1 Test assembly The load set up is shown in figure (13). The components of the compression test assembly were the same as explained by ASTM D1143 and are given below. Hydraulic jack Test beam 60 Reaction piles Steel test plate Load transfer beams Reference beams Dial gauges The main part of this assembly is to have a strong and huge test beam. This test beam was sufficient in size to avoid excessive deflections and to sustain huge loads. The test beam was connected with load transfer beams to transfer the load to these beams. The load transfer beams were connected with reaction piles with the help of ringed steel rod bolted with the reaction piles. Four reaction piles were installed which were multi helix piles containing three helixes of 10, 12 and 14 inches. Multi helix pile were used as reaction piles for safety purposes as they have more capacity than a single helix pile in both tension and compression loadings. Reaction piles were driven to a depth and installation torque similar to the test pile. They all attained a similar torque value at an average depth of 21ft. 61 Figure 13. Loading assembly for compression test conducted There were two reference beams for mounting the dial gauges. They were accurately levelled to measure settlement precisely. In total, four dial gauges were used to measure the resultant settlement and were placed at four corners of steel test plate. The gauges were fixed on the reference beam with the help of magnetic bases. Net settlement was calculated by averaging the settlement of each dial gauge. The test set up with dial gauges is shown in figure (14). 62 Figure 14. Mounted dial gauges on reference beams 4.5.2 Hydraulic jack and calibration There were two hydraulic jacks which could be used to load the test pile. First was AME ram and other was Orbit ram. For this test, AME ram/hydraulic jack was used fig 15 (b). This ram is powered by a generator and is capable of applying a load up to 186,246 lbs. Whenever a load had to be applied on the test pile, pressure was induced in ram and its shaft moved up. The pressure inducing system of ram/hydraulic jack is shown in figure 15 (a). 63 Figure 15 (a): Pressure inducing system Figure 15 (b): Hydraulic jack Figure 15. Hydraulic jack system Like any other loading system, calibration of hydraulics with load is very important as it is responsible for accurate load application on the test pile. Calibration chart of load with hydraulic pressure is given in figure (16). Load increases linearly with the pressure and shows a perfect straight line. Using this calibration plot, the load magnitude can be determined at any pressure. The load in pounds can be calculated by multiplying pressure magnitude in Psi by 20.694 using equation (38). 𝐿𝑜𝑎𝑑 (𝑙𝑏𝑠) = 𝑃𝑟𝑒𝑠𝑠𝑢𝑟𝑒 (𝑃𝑠𝑖) × 20.694 (38) 64 Figure 16. Calibration chart of AME ram 4.5.3 Static axial compression test Static axial compression test is terminated when specific predetermined load magnitude is attained. This load, which can be regarded as the failure load, should not be more than safe or allowable structural load capacity of foundation as suggested by ASTM D1143. To determine the failure load for compression test conducted on a test pile, it is important to discuss the ultimate and allowable structural load capacities of RS2875.203 helical piles. 0.E+006.E+041.E+052.E+052.E+050 4000 8000 12000Load (lbs)Pressure (Psi)65 A.B CHANCE & Hubbell Inc specifies nominal/ultimate structural load capacity of RS2875.203 helical piles as 73.6 kips and allowable structural load capacity of RS2875.203 helical piles as 66.3 kips and 36.8 kips according to load and resistance factor design (LRFD) and allowable stress design (ASD) method. The difference between LRFD and ASD allowable loads is due to difference of safety factors in both design methods. Considering the recommendations of ASTM D1143 standards, the test pile could be loaded to a maximum load equal to 66.3 kips, as it is the maximum structural allowable load for RS2875.203 helical piles. However, 66.3 kips may not be the ultimate load of RS2875.203 helical piles as ultimate and allowable loads of piles are governed by the amount of settlement induced in piles. Deep foundations are considered failed when they settle more than the permissible settlement. The permissible amount of settlement depends upon the settlement tolerance of structures as structures are designed to sustain a certain amount of settlement. If settlement more than the permissible settlement is induced, this can damage the structural integrity. For most cases, when designing the structures, 1 inch (25mm) is considered as the maximum accepted value of settlement (Sakr, 2009). A settlement more than 1 inch can have serious damaging implications on the superstructure of a building. However, the permissible settlement value also depends on the soil on which foundation is lying. Helical piles like other deep foundations carry loads by mobilizing full end bearing strength of soil and cannot support load beyond this strength. If loaded further, they start to settle. End bearing strength is fully mobilized at a settlement equal to 10% of helix diameter (Sakr, 2011). And a pile is considered to have failed if the settlement in pile head reaches 10% of pile diameter (Viggiani et al., 2014). Thus the settlement that would mobilize full end bearing 66 strength of soil at the tip of RS2875.203 helical pile is 1 inch as helix diameter for this pile is 10 inch. Considering the above discussion, the permissible settlement for RS2875.203 helical piles can be fixed at 1inch. Thus, the maximum amount to which the test pile could be allowed to settle is 1 inch. Therefore, it was decided that the axial static compression test will be terminated at a load equal to 66.3 kips or load corresponding to a settlement equal to 1 inch, whichever comes first. This termination criterion served two purposes: safety of loading apparatus and sufficient data for analysis. The load was applied in increments of 5% of 66.3 kips and time interval for each load was kept at ten minutes. Settlement in pile was noted at half, one, two, four, eight and ten minutes of each loading interval. At the start of test, the load movement curve showed sharp increase in slope but this slope started decreasing gradually depicting the depleting load capacity of helical test pile. The test pile attained an average settlement of 0.957 inches at a load of 48.6 kips and the test was terminated at this stage. Load was removed afterwards in eight equal decrements keeping the time interval of decrements same as it was during loading phase. The load movement curve of test pile is shown in figure (17). Up to a load magnitude of 12.99 Kips, rate of increase in settlement was low but further loading of pile yielded a higher settlement rates. Upon removal of load completely, a regain of 0.297 inch was observed and net settlement was 0.66 inch at zero load. The data revealed that settlements occurred right after each load increment and remained almost the same till the end of that increment. 67 Figure 17. Load movement curve of the test pile 0.00E+002.00E+044.00E+046.00E+040 0.2 0.4 0.6 0.8 1 1.2Load (lbs)Settlement (Inches)68 4.6 Interpretation of test pile load movement curve The failure criterions described above have been used to interpret the load movement curve of the test pile in this section. 4.6.1 Davisson offset limit load The Davison offset limit load method has been used for load movement curve of a test pile and is given in figure (18). Elastic compression line is drawn at 20 degrees from the load axis. For convenience, the graph has been plotted with the pile settlements in millimetres. If the graph had been in inches, it would have been too small to draw these lines. 69 Figure 18. Davisson offset limit load The offset line is at distance equal to 5.91mm from the elastic compression line. The ultimate load for the test pile results in 33.55 kips. The allowable load results in 16.77 kips by applying a safety factor equal to 2. The settlements corresponding to ultimate and allowable load are about 0.42 inch and 0.093 inch respectively. The ultimate settlement is 42% of permissible settlement and leave a margin for additional settlement as 58%. The allowable settlement is 9.3% of permissible settlement and leaves an additional settlement margin equal to 90.7%. Although 0.00E+007.56E+031.51E+042.27E+043.03E+043.78E+044.54E+040 4 8 12 16 20 24Load (lbs)Settlement (mm)PUlt70 the ultimate settlement and corresponding additional settlement margin seem appropriate, but piles are subjected to the allowable load rather than the ultimate load. The allowable load and settlement is too small. The margin for additional settlement at allowable load is huge 90.7%, which indicate that Davisson offset limit load criterion results in conservative failure load for RS2875.203 helical piles. Canadian foundation engineering manual (2006) does not recommend considering skin friction for shafts having diameter less than 100 mm. Although, the shaft diameter of RS2875.203 helical pile is less than 100mm but Davisson plot indicates the presence of skin friction along the shaft of the test pile as elastic compression line intersects the load movement graph of the test pile. 4.6.2 10%, 8% and 5% failure loads The settlement corresponding to 10% failure criterion is 1 inch for the test pile. The average total settlement of the test pile was 0.957 inches, which is close to 1 inch. Hence, 0.957-inch settlement is considered the settlement corresponding to 10% failure criterion in this research. Thus, the maximum applied load on the test pile (i.e. 48.63 kips) is the ultimate load corresponding to 10% failure criterion. The allowable load using a safety factor of 2 results in 24.3 kips and the corresponding settlement is about 0.22 inch. This settlement is 22% of permissible settlement and leaves an additional settlement margin equal to 78%. The ultimate settlement corresponding to 10% failure criterion is equal to 1 inch, which is the permissible settlement for RS2875.203 helical piles. Thus, it can be argued that this failure criterion yields real value of failure or ultimate load for RS2875.203 helical piles. Also, there is a sufficient 71 margin of additional settlement at allowable loads which ensures the safety of structures. This indicate that 10% failure criterion yields appropriate values of ultimate and allowable loads without compromising on safety. The 8% failure criterion has been used for the test pile and can be seen in figure (19). The elastic compression line has been drawn in the same way as in Davisson offset limit method. The offset line did not intersect the load movement curve. The load movement curve was extrapolated to intersect the offset line drawn at a distance equal to 8% of 10 inches (i.e. 20.32 mm). The ultimate load for this failure criterion comes out to be about 50 kips. The allowable load results in 25 kips and corresponding settlement is about 0.23 inch. These allowable load and settlement are appropriate. However, the load movement curve should not be extrapolated to get the results because of the significant likelihood that load movement curve would not follow the same path in reality as extrapolated (CFEM, 2006). The extrapolation makes this criterion dubious and less useful for RS2875.203 helical piles. The load movement curve of test pile gives an ultimate load of approximately 36.5 kips corresponding to 5% failure criterion. The corresponding settlement is 0.5 inch which is 50% of the permissible settlement. The allowable load, with a safety factor equal to 2, results in 18.25 kips. The settlement corresponding to allowable load is approximately 0.12 inch and leaves additional settlement margin equal to 88%. 5% failure criterion yields conservative ultimate and allowable loads as compared with 10% failure criterion. This criterion was initially proposed for uplift loads. The compressive load capacities of helical piles are significantly higher than the uplift capacities as discussed in section 4.4.2. Thus, conservative results of this failure criterion 72 are justified. This failure criterion is more applicable for the helical piles having higher diameter helixes because it results in very small settlement for smaller diameter helixes. Figure 19. 8% failure load 4.6.3 L1-L2 failure load The L1-L2 plot for test pile is given in figure (20). The approximate values of PL1 and PL2 are 10140 lbs and 25991lbs respectively. The portion of graph between L1 and L2 is all transition 0.00E+009.06E+031.81E+042.72E+043.63E+044.53E+045.44E+046.34E+040 4 8 12 16 20 24 28 32Load (lbs)Settlement (mm)PUlt73 zone. As mentioned above that failure of pile occurs at transition region or at L2, load corresponding to point L2 (PL2) can be considered as ultimate load capacity of test pile. The ultimate load of this criterion can change if the scale of load movement curve is changed which is a major drawback of this criterion and make its results highly uncertain. Figure 20. L1-L2 failure plot 0.00E+001.50E+043.00E+044.50E+046.00E+040 0.4 0.8 1.2Load (Lbs)Settlement (Inches)L2L174 4.6.4 Chin failure load This criterion assumes that the piles was loaded till the plunging failure condition. The Chin plot for test pile is given in figure (21). It is evident that the straight line for this plot does not exist potentially due to the reason that test pile was not loaded till failure. Figure 21. Chin failure plot 0.00E+005.00E-061.00E-051.50E-052.00E-050 0.2 0.4 0.6 0.8 1Settlement/correspondig loadSettlement (inch)C1= 1.31×10-5C2=0.7×10-5 75 In order to achieve straight line for the figure (21), load movement curve of test pile needed to be extrapolated till the plunging failure load. Because of uncertainties associated with extrapolation, a best fit straight line was drawn and failure load was determined from this line. The ultimate/failure load was found to be 76301.61 pounds or 76.3 Kips from this Chin curve. By applying a factor of safety equal to 2, allowable load capacity becomes 38150.8 pounds or 38.15 Kips. The ultimate load is significantly higher than the maximum applied load on the test pile. The settlement corresponding to the ultimate load would certainly be significantly higher than permissible settlement which make this failure criterion invalid for RS2875.203 helical piles. 4.6.5 Brinch Hansen failure loads The Brinch-Hansen graph plot is shown in figure (22). The low load data point has been ignored and only high load points from load movement curve of test pile has been used to plot Brinch Hansen curve. The calculations show the values of PU and ΔU as 48492.86 pounds and 1.62 inches respectively. It was found that the points 0.8Pu and 0.25Δu did not plot on the load movement curve of test pile. The potential reason is that the failure was not achieved during static axial compression testing of test pile. Failure load for the test pile therefore cannot be determined using Brinch Hansen 80% and 90% criterions. 76 Figure 22. Brinch-Hansen failure plot 4.6.6 De Beer failure load De Beer plot for the test pile has been given in figure (23). It can be seen from the plot, just after the third point, line has changed its course and intersection point can be distinguished. This intersection point between two lines shows a failure load of approximately 10,000 pounds or 10 Kips. 0.00E+005.00E-061.00E-051.50E-052.00E-052.50E-050 0.2 0.4 0.6 0.8 1 1.2(√Settlement)/Test loadSettlement (inch)C277 With a safety factor of 2, allowable load becomes 5000 pounds or 5 Kips which is significantly less as compared with other methods discussed in this chapter. The settlements at ultimate and allowable load are about 0.038 inch and 0.026 inch respectively. These settlements are 3.8% and 2.6% of the permissible settlement which are very small and leave huge margins for additional settlements as 96.2% and 97.4% respectively. Thus, it can be argued that this failure criterion underestimates the true bearing potential of RS2875.203 helical piles which may lead to highly uneconomical pile designs and therefore is not recommended for these piles. Figure 23. De Beer failure plot 1.00E+001.00E+011.00E+021.00E+031.00E+041.00E+050.01 0.1 1Load (Lbs)Settlement (Inches)78 4.6.7 Decourt’s failure load Decourt’s extrapolation for the test pile is given in figure (24). A best fit straight line has been drawn. A few initial data points of load movement curve were discarded because they were not lying in straight line. A best fit line is drawn to calculate the ultimate load. From extrapolation, ultimate load was found to be 62,748.64 pounds or 62.74 Kips. By applying a safety factor equal to 2, allowable load becomes 31374.32 pounds or 31.37 Kips. Figure 24. Decourt’s extrapolation y = -2.9389x + 184412R² = 0.92630.00E+004.00E+048.00E+041.20E+051.60E+051.00E+04 2.00E+04 3.00E+04 4.00E+04 5.00E+04Load/SettlementLoad(lbs)79 The settlement corresponding to allowable load (31.37 kips) is approximately 0.4 inch. The ultimate load is more than the maximum applied load on the test pile and corresponding settlement would certainly be higher than permissible settlement which make this failure criterion invalid for RS2875.203 helical piles. 80 Chapter 5 Estimation of ultimate loads from torque method and bearing equations The ultimate loads of installed helical piles were determined from the torque method and are compared with the ultimate loads estimated using bearing equations and different interpretation criterions of load movement curve of the test pile. The comparison between these techniques will provide useful information regarding suitability and soundness of the torque method for RS2875.203 helical piles. The installation torque is representative of the soil strength. The strength of soil is an important parameter because it defines the load bearing capacity of foundations. Records of installed helical piles indicate that each separate pile achieved an identical magnitude of torque at different depths. Thus, it is justified to say that each pile attained the same load carrying capacity at different depths. Although, these depths were not same but were close to each other. For comparison purposes, it is imperative for each pile to have the same torque magnitude at their toes, in order to ensure that the soil strength is consistent. The test pile has an installation torque equal to 5257 lb-ft at its toe. Hence, for each pile, final toe depth is considered at a depth where it achieves a torque equal to 5257 lb-ft. Average torque values to be used in the torque method are determined considering this final toe depth. Furthermore, by fixing the depth of each helical pile according to the test pile’s toe torque, it can be said that the load capacity and strength characteristics of the soil for each helical pile is similar to the test pile. Thus, the settlement response of each installed helical pile can be determined from the load movement curve of the test pile (figure 17). 81 The ultimate load values calculated using the bearing equations are also compared with those calculated using torque method. Bearing equations use the following soil strength parameters to estimate loading capacities: cohesion and angle of internal friction. These parameters are determined from soil samples taken from a depth 7ft below the natural surface level (NSL). Thus, the bearing equations were used to determine the load capacities of installed helical piles at 7ft from NSL. To be able to compare with the bearing equation results, the torque method is used to estimate ultimate loads using a pile depth of 7ft. The projected helix area used in the bearing equations is 0.531ft2 or 0.0493m2 for RS2875.203 (A.B.CHANCE & Hubbell, 2014). This area also includes the shaft area, which reflects the formation of a soil plug inside the shaft. The projected helix area less the shaft area for RS2875.203 is 0.485ft2 or 0.0451m2. It is important to mention that the ultimate load of each pile is computed from the torque method at two different depths: depth corresponding to a torque of 5257 ft-lb and at 7ft. Ultimate loads at 7ft depth are calculated for comparison with bearing equation’s results. Whereas, loads at a depth corresponding to the torque of 5257 ft-lb are compared with those of static compression test and its interpretation criterions. 5.1 Ultimate loads estimated from the torque method and the corresponding settlements In order to calculate ultimate load from the torque method, an average torque value is multiplied by KT. This average torque can be determined by taking average over the entire depth of pile or taking average over the final length of penetration equal to three times the helix dimeter, as discussed in section 2.4. The length equal to three times the helix diameter is equivalent to 2.5ft 82 for RS2875.203 helical piles, as they have a single helix with 10inch diameter. For these helical piles, two torque coefficients can be used as discussed in section 2.4: 9/ft and 10/ft. Both averages of torque and KT are used to calculate ultimate loads of installed helical piles. Figure (25) below represents the ultimate loads calculated using torque coefficients equal to 9/ft and 10/ft respectively. Blue colored bars in these figures depict load values calculated using torque averaged over entire pile depth, whereas orange bars show ultimate loads using torque averaged over final 2.5ft of depth. Pile number seven in figure is the pile on which static axial compression test was conducted, i.e. the test pile. It can be seen that the ultimate load of the fifth pile shows a considerable dissimilarity from the rest, possibly due to the extraordinary soft soil layers encountered during its installation. For this reason, fifth pile has been ignored in all successive sections. Before discussing the settlement response of installed helical piles, it is imperative to discuss the permissible amounts of settlement. The structural members of any construction project are designed based on a certain threshold amount of settlement. This threshold value is normally taken as 1inch (Sakr, 2009). Aside from the structural design limitations, there is certain amount of settlement required to mobilize the end bearing capacity of helical piles. Typically, end bearing capacity is fully mobilized at settlements equal to 10% of the helix diameter (Sakr, 2011). For RS2875.203 helical piles, this required settlement is equal to 1inch, as these piles have single helix of 10 inches. Moreover, regardless of this permissible settlement, a settlement greater than 1inch could damage the superstructure. Thus, 1inch can be regarded as the maximum permissible amount of settlement for installed helical piles. 83 (a) (b) Figure 25. Ultimate loads calculated from torque method 16.3418.5213.7418.6612.6517.96 17.9640.2941.5636.9833.5526.9938.0839.25010203040501 2 3 4 5 6 7Ultimate load (kips)Pile noKT=9/ft18.1520.5815.2720.7314.0519.95 19.9544.7646.1841.0837.2829.9842.3143.61051015202530354045501 2 3 4 5 6 7Ultimate caoacity (Kips)Pile noKT=10/ft84 The settlements at loads corresponding to the torque averaged over last 2.5ft are given in table (2). The maximum ultimate load for KT =10/ft is in pile number 2 which is 46.18 kips, can be seen in figure (25). Settlement at this load is 0.83 inches as given in table (2). For the other helical piles, settlements are less than 0.83 inch, which is less than the permissible amount. Although settlements at calculated ultimate loads are within limits, foundations are only subjected to the allowable loads which include a safety factor. By applying a safety factor of 2 to the ultimate load, maximum allowable load is calculated to be 23.09 kips. The settlement at this allowable load is about 0.19 inch as given in table (2). This settlement is 19% of the permissible amount, leaving a margin of 81%. Therefore, there is a considerable settlement margin left behind which ensures the safety of structure. Table 2: Settlements at ultimate and allowable loads corresponding to the torque averaged over last 2.5ft Settlement at ultimate loads (inch) settlement at allowable loads (inch) Pile no KT =10/ft KT =9/ft KT =10/ft KT =9/ft 1 0.78 0.64 0.18 0.14 2 0.83 0.68 0.19 0.15 3 0.66 0.52 0.15 0.11 4 0.52 0.41 0.12 0.09 5 0.33 0.26 0.07 0.06 6 0.7 0.56 0.16 0.12 Test pile 0.75 0.6 0.18 0.13 When a KT value of 9/ft is used instead of 10/ft, ultimate loads are further reduced. The corresponding settlements at these reduced loads are less as compared with the settlements at loads corresponding to KT equal to 10/ft. Thus, results become more safe. Ultimate and 85 allowable settlements of helical piles for KT =9/ft can be seen in table (2). The maximum ultimate load is 41.56 kips, and the corresponding settlement at this load is approximately 0.68 inch. When safety factor of 2 is applied on this load, the maximum allowable settlement is estimated as 0.15 inch. This is equal to 15% of 1inch leaving an 85% margin. When torque averaged over the entire depth of helical pile is used, the maximum ultimate load using KT as 10/ft and 9/ft is 20.73 kips and 18.65 kips respectively. Settlements at these loads are determined from the test pile’s load movement curve (figure 17), and are found to be 0.158 and 0.12 inch respectively. When a safety factor of 2 is applied on these ultimate loads, settlements at resultant allowable loads corresponding to KT =10/ft and 9/ft are 0.037 inch and 0.034 inch. These settlements are 3.7% and 3.4% of the permissible amount, leaving a margin of 96.3% and 96.6% behind. This huge settlement margin indicates that such small ultimate and allowable settlements are highly uneconomical for designs and will only mobilize a small fraction of the end bearing potential of soil. To summarize, when a torque value averaged over the final length equal to three times the helix diameter is used, the torque method yields ultimate and allowable loads for RS2875.203 helical piles which are realistic, economical and safe for both values of KT (10/ft and 9/ft). The corresponding settlements for these loads are appropriate and within permissible limits. However, when torque averaged over the entire depth of pile is used in torque method, resultant settlements are very small and are not sufficient to utilize true bearing potential of soil. Such small settlements are highly uneconomical for designs. 86 5.2 Torque method and compression test interpretation criterions The static axial compression load test gives a complete depiction of a foundation’s behavior in different soils. Various failure criterions that tend to interpret load movement curve are discussed earlier in chapter 4. The Davisson offset limit load, 10% and 5% criterions are most noteworthy methods to interpret a load movement curve. The other criterions discussed are unable to interpret the load movement curve of a test pile because of various reasons. Brinch Hansen, Chin and De Beer failure criterions require the pile to be loaded close to the plunging failure, whereas the compression test was aborted before this stage was reached. Moreover, Chin and Decourt’s criterions result in loads that are significantly greater than the maximum applied load on test pile and therefore are not deemed valid for RS2875.203. The comparison of ultimate loads with Davisson, 10%, 5% and 8% failure criterions are given below. Pile no five is ignored here as mentioned in section 5.1. As torque averaged over entire depth of pile results in uneconomical and unrealistic results, only ultimate loads of helical piles calculated using torque averaged over final 2.5ft are discussed here. Figure 26(a) compares ultimate loads estimated using the torque method with KT=10/ft to failure loads corresponding to 10% and Davisson offset limit criterion. Figure 26(b) compares ultimate loads using KT=10/ft to failure loads corresponding to the 5% and 8% criterions. “Ultimate loads” refer to the loads calculated using torque method and “failure loads” refer to those determined by various interpretation criterions. 10% failure load (i.e. 48.63 kips) occurs at a settlement equal to 1inch for installed helical piles as they have single helix of 10inch dimeter. Since 1inch is maximum permissible settlement for RS2875.203 helical piles, loads corresponding to 1inch of settlement can be considered as 87 actual ultimate load for RS2875.203 helical piles. This makes the 10% criterion the most important out of other mentioned criterions for RS2875.203 helical piles. The ratios between ultimate loads calculated from the torque method versus the 10% failure load range between 0.76 to 0.949, which indicated an agreement between the two methods. However, the 10% failure load is higher than the ultimate loads, which suggests that the torque method yields safer results. On the contrary, the Davisson offset failure load is always less than ultimate loads. When compared with the Davisson offset failure load, calculated ultimate loads are up to 37% and 24% greater for KT equal to 10/ft and 9/ft respectively. As discussed in section 4.6.1, Davisson failure criterion results in relatively small and uneconomical loads for RS2875.203 helical pile, while the torque method seems to provide a good approximation of ultimate loads. The 5% failure load is relatively close to the ultimate loads, as ratios of ultimate and 5% failure load range from 1.02 to 1.26. Although 5% failure load is less than the ultimate loads, it was proposed for helical piles loaded in uplift. Uplift load capacities of helical piles are normally less than the compression load capacities (Livneh & El Naggar, 2008;Narasimha Rao et al., 1991). The lower failure load of 5% criterion is thus justifiable. 8% failure load is 50 kips, which is very close to the maximum applied load on test pile (48.63 kips). Ultimate loads calculated using torque method range between 74% to 92% of this failure load. This failure load represents a good agreement with torque method; however, it is important to mention that load movement curve was extrapolated to obtain the failure load. Extrapolation can produce unrealistic results, as the load movement curve might not follow the extrapolated path in reality and failure could occur before reaching the anticipated load. The 88 (a) (b) Figure 26. Comparison of calculated loads for torque averaged over last 2.5ft and failure loads for KT =10/ft 010203040501 2 3 4 5 6 7Load (kips)Pile noUltimate load 10% Davisson01020304050601 2 3 4 5 6 7Load (kips)Pile noUltimate load 5% 8%89 Canadian foundation engineering manual also discourages the extrapolation of a load movement curve. Validity of 8% criterion in this case is therefore dubious. A similar to the results corresponding to KT=10/ft is evident for KT=9/ft. Figure (27) compares the results of static axial compression test criterions to the torque method using KT=9/ft. Ultimate loads are 69-85.5% of 10% failure load, which are 76-94.9% for KT =10/ft. Similarly, ultimate loads are up to 23.8% more than the Davisson offset limit load. The ratios of ultimate loads to 8% failure load range between 0.67 to 0.83. However, loads calculated with KT=9/ft depict a better understanding with the 5% criterion as their ratios are from 0.91 to 1.13. (a) 01020304050601 2 3 4 5 6 7Load (kips)Pile noUltimate load Davisson 10%90 (b) Figure 27. Comparison of calculated loads for torque averaged over final 2.5ft and failure loads for KT=9/ft 5.3 Estimation of ultimate loads using bearing equations The basic form of bearing equations is the same for all foundations; however, the bearing factors used in the equations differ for shallow and deep foundations. Bearing factors suggested by Terzaghi (1943) and Vesic (1963) have been used to calculate ultimate loads on helical piles. Meyerhof (1976) and Vesic (1977) bearing equations are also used for this purpose. Table (3) represents the calculated ultimate loads of installed helical piles at 7ft from NSL. For Terzaghi (1943) and Vesic (1963), values of Nq * are determined from figure (3) and NC * is taken as 9 as suggested by CFEM (2006). The ultimate load corresponding to Meyerhof (1976) is also 01020304050601 2 3 4 5 6 7Load (kips)Pile noUltimate load 8% 5%91 determined and found to be controlled by the limiting value of point/end resistance as discussed in section 2.5.2.2. Vesic (1977) method is used to compute ultimate load at minimum value of reduced rigidity index i.e. 10 which is the minimum mentioned value in Das (1999). This minimum value is used because reduced rigidity index for the airport site could not be determined due to the research limitations. Results show a slight deviation from each other. Where the ultimate load upper bound is is 10.74 kips and lower bound is 8.57 kips, governed by Vesic (1977) and Meyerhof (1976) respectively. Table 3: Calculated ultimate loads of helical piles at 7ft Bearing factors Ultimate load (Kips) Meyerhof (1976) 8.57 Vesic (1977) 10.74 Terzaghi (1943) 9.46 Vesic (1963) 9.06 Ultimate loads corresponding to various combinations of bearing factors as recommended by A.B.CHANCE & Hubbell (2014) are determined and given in table (4). To calculate Nq, equation (28) is used. This equation represents values of Nq reduced by 50% (9 in this case) for a long term application. The non reduced value of Nq i.e. 18 is also used. Similarly, a NC value of 9 represents the recommendation of Canadian foundation engineering manual and 18 is determined from the equation (23). The calculated value of ultimate load is appropriate and within the upper and lower bounds set by Vesic and Meyerhof equations when a value of 9 and 18 is used for NC and Nq respectively. This combination of bearing factors, using the A.B CHANCE & Hubbell method, seems suitable to calculate the ultimate load of RS2875.203 helical piles. 92 Table 4: Bearing factor from A.B CHANCE & Hubbell Inc at 7ft NC Nq QUlt (kips) 9 9 5.65 9 18 8.828 33.36 18 15.53 15.7 9 7.5 5.4 Torque method and bearing equations In order to compare with the bearing equations, the torque method is used to calculate the ultimate load of piles at 7ft. This means that the average torque values are determined considering final depth of 7ft for each pile. The comparison with the bearing equations will check the soundness of torque method. Although bearing equations used for conventional pile foundations are also used for helical piles, the comparison of the ultimate loads of these equations with torque method’s result will provide further understanding to identify the most appropriate equation for helical piles. The calculated ultimate loads are given in table (5). Table 5: Ultimate loads of installed helical piles calculated from torque method at 7ft depth Ultimate loads (kips) Pile no KT=10/ft KT= 9/ft 1 10.27 9.25 2 9.21 8.29 3 10.41 9.36 4 11.6 10.44 5 9 8.1 6 10.2 9.18 TP 9.89 8.9 93 A comparison of ultimate loads estimated by the torque method and bearing equations at 7ft is given in figure (28). Torque method results using both torque coefficients show a good correlation with predicted loads calculated from the bearing equations. Ratios between loads corresponding to different bearing methods and the torque method using both torque coefficients ignoring pile no 4 are given in table (6). Pile no 4 is neglected because its results are showing some deviation from the rest and may affect the comparison. Ratios for pile no 4 are given in table (7). An ultimate load value of 8.828 kips is used for the A.B CHANCE and Hubbell method in tables (5) and (6). Vesic (1977) equation always yield higher ultimate loads than those computed with torque method. However, ultimate loads from other bearing methods are in sound agreement with those from the torque method. The ratios given in table (6) do not exceed 1 significantly, which reflects the suitability of torque method with respect to the bearing equations. Moreover, ultimate loads calculated using a torque coefficient of 9/ft are relatively closer to loads calculated from bearing equations as indicated by their ratios. Ultimate load of the fourth pile deviates from the rest piles, and for this reason, its ratios are different. 94 Figure 28. Comparison of loads computed from torque method and bearing equations Table 6 : Ratios between ultimate loads calculated from bearing equations and torque method KT=10/ft KT=9/ft Terzaghi (1943) 0.90-1.05 1.01-1.16 Meyerhof 0.82-0.95 0.91-1.05 Vesic (1963) 0.87-1 0.96-1.11 A.B CHANCE & Hubbell 0.84-0.98 0.94-1.08 Vesic (1977) 1.03-1.19 1.14-1.32 024681012141 2 3 4 5 6 7Load (kips)Pile noKt=9/ft Kt=10/ft Meyerhof (1976) Vesic (1977)Terzaghi (1943) Vesic (1963) CHANCE95 Table 7: Ratios for pile no 4 KT=10/ft KT=9/ft Terzaghi (1943) 0.81 0.9 Meyerhof 0.73 0.82 Vesic (1963) 0.78 0.86 A.B CHANCE & Hubbell 0.76 0.84 Vesic (1977) 0.92 1.02 When an average of the ultimate loads from the torque method at 7ft for all installed helical piles except pile no. 4 is taken, their correlation with bearing equation’s results are reinforced. Table (7) below shows the ratios between average ultimate loads from the torque method using both KT and loads corresponding to different bearing equations. A KT of 10/ft depicts a reasonable agreement, but 9/ft proves to be the best suited torque coefficient in this case since its ratios are approximately equal to 1. Surprisingly when bearing factors recommended by A.B CHANCE and Hubbell are used to calculate ultimate loads, they show an excellent correlation with the torque method for KT=9/ft as this ratio is equal to 0.99. These results indicate that KT=9/ft generates relatively safe results which are equivalent to ultimate loads predicted by different bearing equations. Table 8: Ratios between ultimate load calculated from bearing equations and average ultimate load of helical piles from the torque method KT=10/ft KT=9/ft Terzaghi (1943) 0.96 1.06 Meyerhof 0.87 0.96 Vesic (1963) 0.92 1.02 A.B CHANCE & Hubbell 0.89 0.99 Vesic (1977) 1.09 1.21 96 5.5 Numerical modelling of RS2875.203 The process of design for a geotechnical structure are relatively complex. For example, the design of foundations is comprised of two parts; firstly, a structural design and then a subsequent geotechnical design. The load carrying capacity of deep foundations/piles is governed by the geotechnical properties of the soil, because piles are designed structurally to carry enormous loads so that the soil strength will fail before the structural strength of pile. Deep foundations sustain their loads through two mechanisms: skin friction and end bearing. Helical piles bear their loads in same way as a conventional deep foundation; however, the helix the tip enhances end bearing resistance substantially. Additionally, a third component called cylindrical shear plane can exist, as discussed in section 2.2.2. In order to account for all these complexities, a software especially designed for helical piles called HelixPile was used in this research. HelixPile is developed by Deep Excavations. Deep Excavations is a private organization which deals with the modern and robust design of deep excavations as well as helical piles. HelixPile software incorporates all aspects of helical piles that are vital for their analysis and design. Where any specific helical pile can be chosen from a database; however, users are able to define their own specifications. This computer program is capable of designing and analyzing following kinds of helical piles. Square shaft helical pile Round shaft helical pile Square micro pile 97 Round micro pile Pipe grouted shaft with smaller core Once pile geometry inputs are given, it calculates all parameters such as projected area of pile and section modulus of helical pile automatically. The soil profiles can be changed to match the actual soil characteristics in field. Helical piles can be analyzed in different stages while incorporating different analysis options such as multiple failure modes. HelixPile is also able to investigate cylindrical failure plane, making this program a unique modern way to analyze helical piles. Moreover, SPT test records can be inputted in to directly estimate helical pile capacity. Helical piles can be analyzed in tension as well as compression. Incorporation of other effects, such as buckling and installation disturbances enable users to analyze the helical pile in detail. HelixPile simulates input soil and pile parameters to predict the load settlement behaviour, and can estimate failure load corresponding to Davisson Offset limit load, New York City criterion (2011-011) and ICC 358 criterion. This software can also calculate ultimate load using bearing capacity equations. Various options of bearing capacity factors are available, such as Meyerhof/Hansen and Vesic. Shaft resistance can also be incorporated. 98 Figure 29. Model of single helix helical pile A simple model of an installed helical pile created using HelixPile is presented in figure (29). HelixPile is used to simulate field conditions for installed piles, assuming the same soil conditions as determined at 7ft depth. Ultimate loads, settlements, stresses and Davisson offset limit loads are determined at various depths of helical piles and are summarized in table (9). The depths at which these values have been calculated using HelixPile conform to the depths of the installed helical piles. 24ft of depth represents the depth of test pile. Similarly, a depth of 17ft conforms to the pile number 1 and 5, and 18ft to pile number 2. A depth equal to 21ft conforms to the pile number 3 and 4. And depth equal to 23ft depicts pile number 6. The ultimate compression loads predicted by HelixPile are significantly less than those calculated from the torque method. HelixPile gives ultimate compression load and settlement as 24.06 kips and 99 0.67inch respectively at 24ft, which is the same depth to which the test pile was driven. Test pile has a settlement of about 0.22 inch corresponding to the load equal to 24.06 kips. It should be noted that the soil strength parameters input to HelixPile at various depths (given in table (9)) are the same as determined at 7ft. The soil’s strength is significantly enhanced after 7ft, as indicated by the sudden exceptional increase in torque magnitude of installed helical piles. Thus, the considerably lower ultimate loads of helical piles and higher settlement of the test pile predicted by HelixPile are justified. However, when ultimate load predicted by HelixPile at 7ft is compared with the ultimate load predicted by the torque method, a sound correlation is seen. Average ultimate loads of installed helical piles calculated from the torque method at 7ft were 9.83 and 8.84 kips, using KT values of 10/ft and 9/ft respectively. On the other hand, HelixPile gives ultimate load as 8.27 kips at 7ft as shown in table (9). There is only a minor difference between these loads, which underscores the fact that HelixPile can predict good approximations of ultimate loads. Table 9: Simulated results of RS2875.203 helical piles using HelixPile Depth (ft) Settlement (inch) Ultimate compression load (kips) Davisson offset load (kips) Settlement at Davisson offset load (inch) Helix stress (ksi) Shaft tip stress (ksi) 7 0.24 8.27 8.24 0.25 5.01 5.11 17 0.47 17.44 13.2 0.29 8.97 9.07 18 0.5 18.43 13.77 0.29 9.36 9.46 21 0.58 21.46 15.57 0.31 10.55 10.65 23 0.63 23.54 16.85 0.33 11.34 11.44 24 0.67 24.06 16.69 0.33 11.78 11.88 100 Chapter 6 Conclusions Helical piles are a type of deep foundations that has potential to provide significant advantages over other conventional deep foundations. The torque method is the most widely used technique to estimate the load bearing capacities of helical piles. Generally, the failure of deep foundations is governed by the amount of settlement induced; however, the torque method does not account for the induced settlements, rendering its results inaccurate due the possibility of excessive settlements at calculated loads. In order to minimize the knowledge gap in helical piles, this research aimed on investigating the torque method for the settlements via installation of helical piles. Seven (7) helical piles were installed at a site in the Kelowna International Airport. In these field tests, RS2875.203 helical piles were utilized. In addition, static axial compression test was conducted on the seventh pile (i.e. test pile). Based on the results and analysis, the following conclusions are summarized for RS2875.203 helical piles: 1. The torque method can be used to estimate the ultimate compressive loads. 2. Both values of torque coefficients (i.e. 10/ft & 9/ft) can be utilized to calculate ultimate compressive loads. 3. When installation torque is averaged over the entire depth of pile, the torque method yields ultimate compressive loads that are drastically low, and the end bearing potential of the soil remains immobilized. These load values are highly uneconomical for pile design. 101 4. When installation torque is averaged over a final length of pile equal to three times the diameter of helix, the torque method yields a good approximation of ultimate compressive loads. The settlements at calculated loads are appropriate and are within permissible limits. 5. The bearing factors used for conventional piles give a good approximation of ultimate loads for helical piles. 6. Torque method yields ultimate compressive loads which are in a good conjunction with the ultimate compressive loads anticipated using bearing equations. 7. 10% failure criterion is the best suited criterion to interpret the load movement curve of the static axial compression test, whereas Davisson offset limit load criterion yields conservative results. When the torque is averaged over the entire depth of a helical pile, it depicts the overall soil strength above the helix. The tensile loads on helical piles are supported by the soil above the helix. The soil strength above the helix is thus more related to the uplift load capacities. On the contrary, compressive loads are supported by the soil beneath the helix. Given the small thickness of helix (i.e. 0.375 inch), the soil just below the helix has virtually the same characteristics as possessed by the soil which is just above the helix. Thus, the soil which is just above the helix gains more significance when pile is loaded in compression. Therefore, the installation torque averaged over the final length of pile equal to three times the helix diameter can be said to represent the soil strength below the helix. Consequently, this torque average yields a good approximation of ultimate compressive load capacities. 102 In this research RS2875.203 single helix helical piles were used due to their widespread availability. The possible directions of future research could be the testing of multihelix helical piles. The investigation of torque method for the formation of cylindrical shear plane in multihelix helical piles will provide valuable information and will help to optimize this technique. Furthermore, helical piles featuring different types and sizes of shaft could be tested to recognise the soundness of torque method in regards to the skin friction of the shaft. Moreover, these piles could be tested in different types of soils to reinforce the findings of this research. Testing a variety of helical piles in various soils will investigate the ultimate limits of the torque method. 103 References A.B.CHANCE, & Hubbell. (2003). Helical Screw Foundation System Design Manual for New Construction. Retrieved from http://www.vickars.com/screwpile_manual. A.B.CHANCE, & Hubbell. (2014). Technical design manual. Abdel-rahim, H. H., Taha, Y. K., & Mohamed, W. E. din E. S. (2013). The compression and uplift bearing capacities of helical piles in cohesionless soil. 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Hoyt, R. ., & Clemence, S. . (1989). Uplift Capacity Of Helical Anchors in Soil. In 12th International Conference on Soil Mechanics and Foundation Engineering (pp. 1019–1022, vol. 2.). 13-18 August 1989. Rio de Janeiro, Brazil: A.A. Balkema. ICC evaluation service, I. (2007). Acceptance criteria for helical foundation systems and devices. 105 Kulhawy, F. H., & Hirany, A. (2009). Interpreted Failure Load for Drilled Shafts via Davisson and L1-L2. In Proceedings of the 2009 International Foundation Congress and Equipment Expo (pp. 127–134). http://doi.org/10.1061/41021(335)16 Lambe, T. W., & Whitman, R. V. (1969). Soil mechanics. New York: John-Wiley & sons Inc. Livneh, B., & El Naggar, M. H. (2008). Axial testing and numerical modeling of square shaft helical piles under compressive and tensile loading. Canadian Geotechnical Journal, 45(8), 1142–1155. http://doi.org/10.1139/T08-044 Meyerhof, G. G. (1963). Some recent research on the bearing capacity of foundations. 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Soils and Foundations (Japanese Society of Soil Mechanics and Foundation Engineering), 31(2), 35–50. Narasimha Rao, S., Prasad, Y. V. S. N., & Veeresh, C. (1993). Behaviour of embedded model screw anchors in soft clays. Geotechnique, 43(4), 605–614. NeSmith, W. M., & Siegel, T. C. (2009). Shortcomings of the Davisson offset limit applied to axial compressive load tests on cast-in-place piles. In Contemporary Topics in Deep Foundations (pp. 568–574). http://doi.org/10.1061/41021(335)71 O’ Neill, M. W., & Reese, L. C. (1999). Drilled Shafts: Construction procedures and design methods. Office of infrastructure, Federal highway administration. Washington, D.C. Publication No. FHWA-IF-99-025. Retrieved from http://isddc.dot.gov/OLPFiles/FHWA/011594.pdf Rodrigo, S., & Junhwan, L. (1999). Pile design based on cone penetration test results. Report no FHWA/IN/JTRP-99/8. Joint transportation research program Purdue university. Sakr, M. (2009). Performance of helical piles in oil sand. 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Estimation of compressive load bearing capacity of helical piles using torque method and induced settlements Khan, Muhammad Umair Shabbir 2016
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Title | Estimation of compressive load bearing capacity of helical piles using torque method and induced settlements |
Creator |
Khan, Muhammad Umair Shabbir |
Publisher | University of British Columbia |
Date Issued | 2016 |
Description | Helical piles are deep foundations that have a helix at the end. The traditional approaches to determine the load capacity such as loading tests and in situ tests (i.e. SPT, CPT and LCPC) are not economically feasible for the small scaled constructions, for which helical piles are generally recommended. In order to estimate the ultimate load that helical piles can carry, torque method is thus mostly used. Torque method does not account for the possible settlements induced at calculated loads. Settlement induced is the main load capacity governing factor for deep foundations, as they are considered failed when a settlement more than the permissible amount is attained. The possibility that the piles might fail well before the calculated load is achieved because of excessive settlements make the results of torque method dubious. This research attempts to investigate the torque method for the settlements and for its precision. For this purpose, seven RS2875.203 helical piles were installed and their ultimate compressive loads are calculated using the torque method. On seventh pile, static axial compression test was conducted. The settlements at torque method’s ultimate loads are determined from the load movement curve of compression test. Results underscore that the settlements at torque method’s ultimate and allowable loads are within the permissible amount. The load movement curve of compression test is interpreted using different failure criteria to calculate the failure load. Results show that 10% failure criterion is the most suitable criterion to interpret the load movement curve of RS2875.203 helical piles. Additionally, different bearing equations are used to compute the ultimate compressive loads of helical piles. Result suggest that the loads calculated using torque method and bearing equations correlate well with each other. |
Genre |
Thesis/Dissertation |
Type |
Text |
Language | eng |
Date Available | 2016-09-08 |
Provider | Vancouver : University of British Columbia Library |
Rights | Attribution-NonCommercial-NoDerivatives 4.0 International |
DOI | 10.14288/1.0314176 |
URI | http://hdl.handle.net/2429/59121 |
Degree |
Master of Applied Science - MASc |
Program |
Civil Engineering |
Affiliation |
Applied Science, Faculty of Engineering, School of (Okanagan) |
Degree Grantor | University of British Columbia |
Graduation Date | 2016-11 |
Campus |
UBCO |
Scholarly Level | Graduate |
Rights URI | http://creativecommons.org/licenses/by-nc-nd/4.0/ |
Aggregated Source Repository | DSpace |
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