The effects surcharging has on the rate of secondary settlement on clays along the Wasatch front

Secondary compression of foundation soils can cause long-term settlement damage to bridges, their foundations and approach embankments, overlying pavements and other nearby constructed works. Because this type of settlement is long-term and manifests itself many months to years following embankment construction, it sometime goes unnoticed until it damages overlying or nearby infrastructure. Surcharging or preloading of the earthen embankments and underlying compressible soils is the most commonly deployed strategy to reduce the magnitude of secondary compression. Surcharging or overconsolidating of the foundation soils can be used to reduce the post-construction secondary settlement. In the course of this research twenty-two consolidation tests and eighty-eight time rate tests were performed on Pleistocene and recent fine-grained, cohesive, lacustrine deposits comprised of Lake Bonneville and more recent clays, most likely of Utah Lake origin located along the Wasatch Front. Prior to analyzing the data, the test results were screened using the sample quality designation (SQD). Plots of the adjusted amount of surcharge (AAOS) were plotted versus the normalized rate of secondary settlement (C?-/C?) and compared with the research performed by Ng (1998). The data from this thesis plots higher than that reported by Ng (1998). This higher trend agrees better with the long-term settlement performance monitoring data obtained from the I-15 Reconstruction Project. Data from the time rate tests were used to determine the C?/CR ratio giving a mean value of C?/CR = 0.0456. This value was also compared with the research performed by Ng (1998) which has a value of C?/CR = 0.0433, this correlates well. A plot of moisture content vs. CR was developed and compared with research done by Bartlett and Lee (2004). The data from this thesis tendline is slightly lower than that reported by Bartlett and Lee (2004), but still correlates well. The correlation of moisture content vs. the C?/CR ratio was explored which shows promise, but more observations are needed to improve the statistical support for this relation.

THE EFFECTS SURCHARGING HAS ON THE RATE OF SECONDARY SETTLEMENT ON CLAYS ALONG THE WASATCH FRONT by Zach Montgomery Gibbs A thesis submitted to the faculty of The University of Utah in partial fulfillment of the requirements for the degree of Master of Science Department of Civil and Environmental Engineering The University of Utah December 2015 Copyright © Zach Montgomery Gibbs 2015 All Rights ReservedThe University of Utah Graduate School STATEMENT OF THESIS APPROVAL The thesis of Zach Montgomery Gibbs has been approved by the following supervisory committee members: Steven Bartlett , Chair 12/10/2014 Date Approved Evert Lawton , Member 12/10/2014 Date Approved Richard Porter , Member 12/10/2014 Date Approved and by Michael Barber , Chair/Dean of the Department/College/School of and by David B. Kieda, Dean of The Graduate School. Civil and Environmental Engineering ABSTRACT Secondary compression of foundation soils can cause long-term settlement damage to bridges, their foundations and approach embankments, overlying pavements, and other nearby constructed works. Because this type of settlement is long-term and manifests itself many months to years following embankment construction, it sometime goes unnoticed until it damages overlying or nearby infrastructure. Surcharging or preloading of the earthen embankments and underlying compressible soils is the most commonly deployed strategy to reduce the magnitude of secondary compression. Surcharging or overconsolidating of the foundation soils can be used to reduce the postconstruction secondary settlement. In the course of this research, twenty-two consolidation tests and eighty-eight time rate tests were performed on Pleistocene and recent fine-grained, cohesive, lacustrine deposits comprised of Lake Bonneville and more recent clays, most likely of Utah Lake origin located along the Wasatch Front. Prior to analyzing the data, the test results were screened using the sample quality designation (SQD). Plots of the adjusted amount of surcharge (AAOS) were plotted versus the normalized rate of secondary settlement (Cα'/Cα) and compared with the research performed by Ng. The data from this thesis plot higher than those reported byNg. This higher trend agrees better with the long-term settlement performance monitoring data obtained from the I-15 Reconstruction Project. Data from the time rate tests were used to determine the Cα/CR ratio, giving a mean value of Cα/CR = 0.0442. This value was also compared with the research performed by Ng, which had a value of Cα/CR = 0.0433. This correlates well. A plot of moisture content vs. CR was developed and compared with research done by Bartlett and Lee. The data from this thesis trendline are slightly lower than that reported by Bartlett and Lee, but still correlate well. The correlation of moisture content versus the Cα/CR ratio was explored which shows promise, but more observations are needed to improve the statistical support for this relation. iv TABLE OF CONTENTS ABSTRACT ........................................................................................................ iii LIST OF TABLES .............................................................................................. vii ACKNOWLEDGEMENTS ................................................................................. viii 1 INTRODUCTION ............................................................................................. 1 1.1 Overview ............................................................................................ 1 1.2 Scope and Purpose of Research ....................................................... 3 2 BACKGROUND ............................................................................................... 5 2.1 General Discussion ............................................................................ 5 2.2 Mesri et al. Concept of Secondary Compression ............................... 8 2.3 Surcharging to Reduce Secondary Compression .............................. 9 2.4 Surcharge Design Using Methodology Developed by Mesri ............ 10 2.5 Surcharge Design Using Methodology Developed by Ladd ............. 12 2.6 Application of Ladd's Method to the I-15 Reconstruction Project ..... 16 3 RESEARCH OBJECTIVES AND TASKS ...................................................... 24 3.1 Research Objectives ........................................................................ 24 3.2 Research Plan .................................................................................. 24 3.3 Tasks ................................................................................................ 25 4 FIELD INVESTIGATION ................................................................................ 27 4.1 Introduction ...................................................................................... 27 4.2 Field Investigations .......................................................................... 27 5 LABORATORY TESTING .............................................................................. 52 5.1 Test Procedures ............................................................................... 52 5.1.1 Introduction ............................................................................. 52 5.1.2 Testing Equipment .................................................................. 53 5.1.3 Sample Setup .......................................................................... 555.2 Laboratory Test Program ................................................................. 58 5.2.1 Determining of the Preconsolidation Stress ............................ 61 5.2.2 Determination of the Rate of Secondary Compression Cα ...... 63 5.2.3 Determining C'α ....................................................................... 64 6 RESULTS AND INTERPRETATIONS ........................................................... 65 6.1 Lab Tests and Data Screening ......................................................... 65 6.2 Relationships for Cα, Cα′, and Cα′/Cα ................................................ 72 6.3 Cα/CR Ratio ...................................................................................... 79 6.4 Moisture Content Correlations .......................................................... 82 7 CONCLUSIONS ............................................................................................. 88 7.1 Summary of Thesis Objectives ......................................................... 88 7.2 Mesri's Concept of Secondary Compression ................................... 88 7.2.1 Cα/CR ....................................................................................... 89 7.2.2 Creep Behavior as a Function of AAOS ................................... 89 7.2.3 The Time Before Creep Resumes After the Removal of a Surcharge ......................................................................................... 90 7.3 Recommendation for a Laboratory Testing Program ....................... 90 7.4 Additional Design Guidance ............................................................. 91 7.5 Recommendation for Further Testing ............................................... 92 Appendices A PLOTS FOR PRECONSOLIDATION STRESS ............................................. 93 B PLOTS OF RATE OF SECONDARY SETTLEMENT .................................. 357 C LABORATORY TESTING PROCEDURE ................................................... 446 D CPT PLOTS OF SOIL BEHAVIOR TYPES ................................................. 457 E COMPARISON OF INCREMENTAL LOADING AND INSTANT LOADING 462 REFERENCES ............................................................................................... 466 vi LIST OF TABLES 4.1 Boring locations, depths, and drilling dates ................................................. 33 4.2 CPT locations .............................................................................................. 33 6.1 Listing the sites, depths, moisture content, preconsolidation stress, compression and recompression ratio, and the rate of secondary settlement at different OCRs .................................................................................................. 66 6.2 Values of strain at σ'vo and the corresponding rating of SQD...................... 67 6.3 Site, depth, effective vertical stress, and SQD ............................................ 69 ACKNOWLEDGEMENTS I would like to thank Dr. Steven Bartlett for approaching me and giving me this great opportunity. I have gained a lot of knowledge and experience while performing the research for this thesis. I would also like to thank him for the help and guidance he has provided. I would like to thank Ramesh Neupane and Shun Li for the help they both provided with gathering lab readings and the support they offered. I would like to thank my wife, Stacie, for her unwavering support and patience through the course of this thesis and for always reminding me to be diligent. 1 1 INTRODUCTION 1.1 Overview Secondary compression or secondary settlement or creep settlement is a continuation of the volume change of a compressible soil under a constant (i.e., nonchanging) loading without the associated changes in the effective stress of the soil fabric. This behavior begins to be manifested near the end of primary consolidation and continues indefinitely, but at a nonlinear diminishing rate. In contrast to primary consolidation which is associated with compression due to pore water pressure dissipation, secondary compression begins when the specimen achieves a constant effective stress after essentially all excess pore water pressure has dissipated that was induced by the initial loading event (Holtz et al., 2011). Secondary compression of foundation soils at deep, compressible, soil sites can cause long-term settlement damage to bridges, their foundations and approach embankments, overlying pavements, and other nearby constructed works. Because this type of settlement is long-term and manifests itself many months to years following embankment construction, it sometime goes unnoticed until it damages overlying or nearby infrastructure. For example, the collective secondary compression is often significantly large enough to produce a severe "bump" at pile-supported bridges where the approach embankment has settled differentially relative to the bridge and bridge abutments. 2 The magnitude and potential deleterious effects of secondary compression on the future performance of the interstate system were important geotechnical design and performance considerations during the reconstruction of I-15 in the northern part of Salt Lake Valley, Utah during 1998 to 2001. Surcharging or preloading of the earthen embankments and underlying compressible soils was the most commonly deployed strategy to reduce the magnitude of secondary compression; however, soil improvement and light-weight embankment materials were also used as settlement mitigative measures. An important part of embankment design for the I-15 project was a systematic evaluation of the required amount (i.e., height) of surcharge to reduce the secondary compression to acceptable, postconstruction, performance goals. Associated with this issue is also the required time that such surcharge is to remain in place to achieve the desired long-term settlement performance goal. For the I-15 project, the performance goal was to surcharge the foundation soils enough that the embankment in the bridge approach area did not settle more than 3 inches in a 10-year, postconstruction period. Whatever the desired outcome, the settlement performance goals should be clearly defined by the project team in consultation with the owner. For fast-paced construction, the corresponding settlement calculations and design, construction settlement monitoring, and project communication are vital if these goals are to be realized. In addition to the amount of surcharging employed, the time or duration that the surcharge is to remain in place (i.e., surcharge duration) strongly impacts the postconstruction settlement performance and the construction schedule. 3 Because the surcharge duration can be long for deep soil sites, this can significantly impact the construction schedule; hence, there is an inherent tendency by the contractor and the project team to try to shorten the surcharge duration in order to expedite the construction. Therefore, construction settlement monitoring to assess the progression of primary consolidation settlement and a decision framework for selecting when to remove the surcharge are essential in achieving the settlement performance goals and delivering a timely project. 1.2 Scope and Purpose of Research The primary purpose of this research is: (1) to quantify the effects that surcharging (i.e., preloading) has on secondary compression settlement for fine-grained soils located along the Wasatch Front, Utah area using one-dimensional (1D) consolidation tests performed on conventional table top oedometers and (2) to develop the information, relations, equations, charts, etc. required to develop and implement a surcharge design for these sediments using the framework developed by Ladd (1989) and Ng (1998), (3) to confirm or recommend changes, if any, to the relations required for surcharge design as presented by Ladd (1989) and Ng (1998), and (4) to make other recommendations about the implementation of the results of this research in regard to surcharge calculations and design. These purposes and objectives will be explored via: (1) reviewing the geotechnical literature that supports the approaches of Mesri et al. (Mesri and Castro, 1987; Mesri et al., 1994) and of Ladd (Ladd, 1989 and, Ng 1998), (2) undisturbed sampling of cohesive soils from four soft soil sites located along the 4 urban Wasatch Front, Utah, (3) laboratory testing of these specimens in conventional oedometers to determine the rate of secondary compression as a function of preloading (i.e., overconsolidation ratio), (4) evaluating and presenting the results of the laboratory test program in the analysis framework developed by (Ladd, 1989 and Ng 1998), and (5) making recommendations, if any, about modifications to the relationships or methods developed by Ladd (1989) and Ng (1998). 5 2 BACKGROUND 2.1 General Discussion Consolidation settlement of soil occurs from three general mechanisms: (1) quasi elastic compression of the soil fabric upon reloading that occurs below the preconsolidation stress, (2) primary consolidation settlement resulting from significant compression of the soil fabric from an applied stress that exceeds the preconsolidation stress, and (3) secondary consolidation or secondary compression of the soil fabric which is a complex combination of processes that initiates near the end of primary consolidation and continues as a long-term process under a constant load or unchanging effective stress (Holtz et al., 2011). Secondary compression is generally thought of as void ratio change in the soil fabric occurring at a relatively slow rate after primary consolidation is essentially completed. However, some researchers have noted that secondary compression occurs in conjunction with primary consolidation settlement, but at a slower rate; hence, its effects are in a large part masked by the significantly greater magnitude and faster rate of primary consolidation settlement realized during the initial part of the consolidation process. Therefore, it is difficult to distinguish secondary compression from experimental data when the sample is undergoing large void ratio changes associated with primary consolidation (Takeda et al., 2013). Although such distinctions are important for the advancement of 6 consolidation theory, this research will adopt the classical construction shown in Figure 2.1 to define the time corresponding to the end of primary consolidation, tp, which also marks the beginning of secondary compression (Raymond and Wahls, 1976). In this definition, tp is calculated as the intersection of the straight lines that define primary consolidation and secondary compression on a void ratio, e, versus log of time plot. As the rate of primary consolidation diminishes, secondary compression becomes the dominate process. At this point, almost all of the excess porewater pressure (i.e., porewater pressure above hydrostatic) has dissipated from the soil fabric that was caused by the initial loading event. Hence, secondary compression is also defined as void ratio change or settlement occurring when the effective stress in the soil fabric is no longer significantly changing (Holtz et al., 2011). The continued settlement at a diminishing rate is a result of creep, viscous behavior of the soil fabric, compression of organic matter, and other processes. Figure 2.1 Definition of Cα from 1D time rate of consolidation test (after Raymond and Wahls, 1976; Ng 1998). 7 Holtz et al. (2011) suggest that the following assumptions must be adopted to provide a working hypothesis about the behavior of fine-grained sediments undergoing secondary compression based on work by Ladd (1971) and Raymond and Wahls (1976). They discuss the relative merits and practical consequences associated with these assumptions which have been briefly summarized below: 1. The rate of secondary compression is independent of time, at least during the time span of engineering interest. (This assumption is discussed later in this report.) 2. The rate of secondary compression is independent of the soil layer. 3. The rate of secondary compression is independent of the load increment ratio (LIR), as long as some primary consolidation occurs. 4. The ratio of the rate of secondary compression to the compression index is approximately constant for many geo-materials over the range of engineering stresses (also discussed later in this report). The amount of volume change during secondary compression is calculated from the secondary compression index, Cα, which represents the rate of secondary compression defined by: Cα = Δ e / Δ log t (2-1) where: Δ e is the change in void ratio, Δ log t is log t - log tp. The value of Cα represents the change in void ratio, e, divided by the change in log of time for the portion of the time rate of consolidation curve extending beyond the end-of-8 primary (EOP) consolidation (Figure 2.1) (Holtz et al., 2011). When plotted on a semi-log plot, Cα represents the slope of this semi-log linear portion of secondary compression that occurs beyond EOP consolidation (Figure 2.1). The magnitude of secondary compression or settlement for a specimen or layer is typically calculated by the following formula for 1D consolidation: S = [Cα / (1+ eo)] Ho log t / tp (2-2) where: Ho is the height of the specimen or layer, eo is the initial void ratio, Cα is the rate of secondary compression, t is the elapsed time after the end of primary consolidation and tp is the time required to reach the end of primary consolidation (Figure 2.1) (Holtz et al., 2011; Terzaghi et al., 1996). 2.2 Mesri et al. Concept of Secondary Compression Mesri et al. have shown that the secondary compression index for a normally consolidated soil, Cα NC, is correlated with and can be estimated from the virgin compression index, Cc, or the compression ratio CR of that soil (Mesri and Castro, 1987; Ladd 1989; Mesri and Feng, 1991; Mesri et al., 1994; Terzaghi et al. 1996; Ng, 1998; Saye and Ladd, 2000) where the compression ratio is defined as: CR = Cc / (1 + eo). (2-3) Because of this correlation, the method proposed by Mesri et al. (Mesri and Castro, 1987; Mesri et al., 1994) is often used to estimate the rate of secondary compression for a given geologic unit. In the approach proposed by Mesri et al., 9 the ratio of Cα/Cc or Cα/CR has been found to be considered relatively constant for sediments of the same geologic origin. Therefore, this ratio can be used to estimate Cα NC if Cc or CR has been determined for the soil of interest. It should be noted that the Cα/CR ratio used in this research is the same as Cα/Cc ratio of Mesri and Castro (1987) and Mesri et al. (1994) because for Cα/CR, the unit of strain in Cα = dεv / dlog t and that found in CR = dεv / dlog σ'vc cancel each other; and for Cα/Cc, the unit of change in void ratio Cα = de / dlogt and that found in Cc = de / dlog σ'vc cancel each other (Ng, 1998), thus Cα/CR or Cα/Cc can be used interchangeably. To implement the Mesri et al. method (Mesri and Castro, 1987; Mesri et al., 1994), values of Cα/Cc or Cα/CR are typically determined from a laboratory consolidation testing program from each geologic unit of interest. Once this ratio is established, additional estimates of Cα can be made for a given deposit using laboratory or field estimates of Cc or CR and the corresponding values of Cα/Cc or Cα/CR ratio for that deposit. 2.3 Surcharging to Reduce Secondary Compression Terzaghi et al. (1996) recognized the fact that if the final in situ state of stress resulting from a loading event imparted to a foundation soil is higher than the original preconsolidation stress of the soil and if the time for primary consolidation, tp, is small perhaps due to the installation of prefabricated vertical drains (PVD), then the amount of secondary compression settlement can be relatively large. However, this can be reduced to acceptable levels by using surcharging of the foundation soil during the last stage of embankment 10 construction. Surcharging has the effect of preloading the soil (i.e., overconsolidating) and reducing the rate of secondary compression when compared to the rate of secondary compression for a normally consolidated soil (i.e., a soil that has not been surcharged). Surcharge methodologies developed by Ladd (1989) and by Mesri (1986) have been used in engineering practice to develop a surcharge approach to reduce the effects of secondary compression associated with embankment construction atop on relatively soft, compressible, foundation soils. The next two sections of this report describe Mesri's and Ladd's methodologies. The final section of this report discusses how Ladd (1989) methodology was applied to the I-15 Reconstruction Project to reduce the effects of secondary compression. The data developed from this roadway project in conjunction with additional field and laboratory testing and evaluations performed as part of this research become, in part, the basis for the surcharge design guidance developed herein. 2.4 Surcharge Design Using Methodology Developed by Mesri Mesri (1991) has shown the behavior of a soil subjected to surcharging (Figure 2.2). The removal of the surcharge leads to rebound of the specimen, including primary rebound up to the time tpr and secondary rebound that levels off at time tl and is followed by secondary compression occurring at a nonlinear rate on a log of time plot. In this figure, tpr, tl, and t are measured from the time when the surcharge load was removed (i.e., t's). The postsurcharge secondary compression behavior, C'α, shown in Figure 2.2 is initially small and subsequently gradually increases with time. Mesri has shown that at large values of t, the 11 Figure 2.2 Mesri's basic concepts of the effects of surcharge on secondary compression (Mesri 1991 from GEO-COAST 91, unpublished proceedings). behavior of the secondary compression depends on the initial shape of the EOP e vs log σ'v curve at the state of stress with the surcharged load applied. Hence, because C'α is not constant with time, a secant value C''α is used in evaluations where the slope of C''α is defined by the line connecting tl with t. In Mesri's approach, the surcharging effort is expressed as the total surcharge ratio: Rs = (σvs / σ'vf) - 1 (2-4) where: σvs is equal to σ'vf + Δσvs and σ'vf is the final effective vertical stress after removal of surcharge and Δσvs is the total stress applied by the surcharge load. The surcharging time ratio, t's / t'ps, affects the behavior of the curve where t's is the duration of the surcharge, and t'ps is the time to EOP compression under the surcharge load. For cases where the surcharge load is removed before EOP compression, the above equation is rewritten as: 12 R's = (σ'vs / σ'vf) - 1 (2-5) where: σ'vs is the maximum effective vertical stress reached before the removal or surcharge. Hence, when t's / t'ps = 1, then Rs = R's. If t's exceeds the time to the EOP compression, then the value of R's is adjusted to reflect the aging and the effective surcharge ratio, R's, is equal to: R's = (σp′ - σvf′) / σvf′ (2-6) where: σp′ is the apparent preconsolidation stress due to aging of the soil under the surcharge load. 2.5 Surcharge Design Using Methodology Developed by Ladd The methodology of Ladd (1989) has many aspects that are similar to that of Mesri, but the part of the curve that defines secondary compression has been simplified (Figure 2.3). The most important difference is that Ladd's method assumes that the rate of secondary compression is linear when plotted on a semi-log plot. The linear portion begins after the start of secondary compression, ts, and continues thereafter (Figure 2.3). Hence, Cα' is calculated from the slope of a line fitted through the linear most part of the vertical strain measurements that follow ts. This construction makes Ladd's method easier to apply than that of Mesri. In Ladd's method, if the soil is surcharged (i.e., overconsolidated) and subsequently aged under this surcharge load, it undergoes secondary compression at a reduced rate, Cα′, when compared with the, unsurcharged, unaged rate of secondary compression Cα (Figure 2.3). The aging time of the soil 13 Figure 2.3 The effects of surcharging (i.e., preloading) on the rate of secondary compression (after Ladd, unpublished notes). under the surcharge stress (i.e., elapsed time between tp and tr) reduces the rate of secondary compression from Cα, the normally consolidated value, to a lesser, overconsolidated value, Cα′, which has a reduced slope (Figure 2.3). Hence, if a soil can be surcharged and aged, the amount of postconstruction creep settlement is reduced when compared with the unsurcharged, normally consolidated value. Therefore, in applying this concept to developing the surcharge design for an embankment and its foundation soil, an evaluation is made to provide sufficient surcharging of the foundation soil so as to reduce Cα′ to a value that will reduce the amount of secondary compression. The value of Cα′ required is a function of the thickness of the foundation soil layer undergoing secondary compression and the postconstruction settlement goal selected by the 14 project team. The amount of secondary compression settlement for normally consolidated sediments (i.e., unsurcharged soils) is calculated from: Ss = H1 Cα log (t / tp) (2-7) where: Ss is the amount of secondary compression settlement, H1 is the thickness of the layer undergoing secondary compression, Cα is the normally consolidated rate of secondary compression, tp is time to end of primary consolidation and t is the time beyond tp. If the soil has been surcharged (i.e., overconsolidated) and then unloaded, then a reduced rate of secondary compression, Cα′, is used in lieu of Cα: Ss = H1 Cα′ log (t / tp) (2-8) When using Ladd's methodology for a surcharge design, the soil is loaded from (σ'v) to the surcharge stress (σ'vs) and then is unloaded to the final stress (σ'vf). The difference between these values is defined as the amount of surcharge (AOS) which is determined from the following: AOS = (σ'vs - σ'vf)/ σ'vf (2-9) For staged embankment construction, σ'vs should be calculated using the full embankment height plus the height of surcharge and σ'vf should be calculated using the final embankment height after surcharge removal, but including the weight of the overlying pavement system, if present. If the surcharge stress σ'vs exceeds the preconsolidation stress of the soil (i.e., primary consolidation is initiated) and primary consolidation is allowed to go to completion, but significant secondary compression is not allowed under the 15 surcharge load (i.e., soil is not allowed to age by removing the surcharge, tr, at the same time as tp is achieve), then Equation 2-9 is appropriate and Equation 2-8 should be used to calculate the secondary compression of the soil using Cα′ appropriate for the AOS achieved by the surcharge load. The AOS and the overconsolidation ratio, OCR, are related by: AOS = OCR - 1 (2-10) If the soil is aged by allowing the surcharge to remain in place for some time after the end of primary consolidation (i.e., tr > tp), then the AOS is adjusted to account for the apparent increase of σ'p above σ'vs, due to the aging. The adjusted amount of surcharge (AAOS) is determined using the following: AAOS = (σ'p- σ'vf)/ σ'vf (2-11) where: σ'p = σ'vs(tr/tp)Cα/CR. Figure 2.3 also shows a time delay between the removal of the surcharge, tr, and the initiation of the reduced rate of secondary compression, ts. Initially, there is a brief heave event, followed by the initiation of secondary compression at a reduce rate represented by Cα'. The length of this time delay is a function of AAOS (Ladd, 1989 and Ng, 1998). The time delay is longer for higher AOS values, as discussed in the next section using data from the I-15 Reconstruction Project (Ng, 1998). This delay is also beneficial in reducing the amount of secondary compression occurring in the postconstruction period. For evaluation purposes, the value of ts represents the point in time when the soil has reached its maximum heave value (Figure 2.3). 16 2.6 Application of Ladd's Method to the I-15 Reconstruction Project The I-15 Project was a fast-paced reconstruction project that began during the spring of 1998 and ended in the fall of 2001, just prior to the 2002 Winter Olympic Games in Salt Lake City, Utah. At that time, it was the largest public highway construction project to be accomplished using a design-build project delivery system. During this 3.5-year period, the design-build consortium demolished and rebuilt 26 km (16.2 miles) of urban interstate, widening the roadway from 6 to up to 12 lanes at a total cost of about $1.4 billion. A large part of this cost was spent erecting 144 overpass bridge structures, constructing 160 mechanically stabilized earth (MSE) retaining walls and placing 3.8 million m3 (134 million ft3) of new embankment. The design-build contract featured a 50-year design life and an optional 10-year corrective maintenance agreement (Farnsworth et al., 2008). The strict project completion date presented unique challenges to the design-build team. Perhaps the most demanding was developing strategies to address the impacts of consolidation settlement in the northern segment of the project near the downtown area. Here, compressible, fine-grained lacustrine sediments deposited by Pleistocene-age Lake Bonneville underlie about 5 m (16.4 ft) of Holocene alluvium (Figure 2.4). The lacustrine sediments are approximately 15 m (49.2 ft) thick, consisting of inter-bedded silty clay and clayey silt (CL, ML), plastic clays and silts (CH, MH), and fine clayey and silty sands (SC, SM) and are lightly overconsolidated (OCR ≈ 1.5). Interbedded, subaqueous silts, fine sands and low plasticity clays are found in the middle of 17 Figure 2.4 Typical cone penetrometer (CPT) log and soil descriptions for downtown segment of I-15 Reconstruction Project, Salt Lake City, Utah (Unpublished I-15 data). the Lake Bonneville sediments and separate the upper and lower Lake Bonneville clays. These upper and lower clay units are compressible (CR values ranging from 0.1 to 0.35), have relatively low undrained shear strength (25 to 50 kPa) and require substantial time to complete primary consolidation. In this regard, settlement records from the mid-1960s construction of I-15 show that a consolidation settlement over a period of 2 to 3 years. For example, Figure 2-5 shows a settlement record from the mid-1960s construction, for an embankment constructed over the typical soil conditions represented in Figure 2-4. The record 0 5 10 15 20 25 30 35 40 0 8 USCS CH, Fat Clay MH, Elastic Silt CL, Sandy Lean Clay ML, Silt SC, Clayey Sand SM, Silty Sand 0 5 10 15 20 25 30 35 40 0 250 500 fs (kPa) 0 5 10 15 20 25 30 35 40 0 500 1000 u (kPa) 0 5 10 15 20 25 30 35 40 0 10 20 30 40 Depth (m) qc (MPa) Recent Alluvium Upper Bonneville Clay Interbeds Lower Bonneville Clay Deeper Alluvium Deeper Clay 18 Figure 2-5 Typical settlement record for the mid-1960s construction of I1-5 in downtown Salt Lake City, Utah. typical 8 to 10 m high embankment underwent 1 to 1.5 m of primary shown in Figure 2-5 is typical of those recorded during the mid-1960s construction for this type of soil condition. This figure shows that fill placement was performed in multiple stages to reach the peak loading condition and then the primary settlement was allowed to take place prior to removal of the surcharge. These large magnitudes of settlement (1.4 m or 4.5 ft) and long consolidation settlement durations (approximately 2 years) can be attributed directly to the soft, thick, compressible Lake Bonneville clay layers. In the mid-1960s, the bridge foundations, bridge, approaches and pavement were not placed until such settlement was essentially finished (Farnsworth et al., 2008). 19 The I-15 Reconstruction Project team established a long-term performance goal to limit the amount of postconstruction settlement (i.e., secondary compression settlement) of the foundation soils to 3 inches, or less, in a 10-year postconstruction period. The design surcharge height and surcharge duration were calculated to meet this performance goal (Saye and Ladd, 2000). The I-15 Reconstruction Project utilized three geotechnologies to address the large and potentially damaging effects of primary and secondary consolidation originating from compression of the soft foundation soils prevalent beneath much of the northern part of the project. The first and most widely utilized approach was to apply surcharging in conjunction with the construction of drain (PVDS) were installed in the foundation soils prior to wall or embankment construction. Once the surcharged embankments had reached their design height, primary consolidation settlement of the foundation soil was allowed to take place, followed by surcharge removal. The second approach was to essentially eliminate most of the potential foundation settlement by using light-weight fill (e.g., scoria and EPS geofoam), thus greatly minimizing the loading condition imposed on the foundation soils. The third approach involved strengthening the foundation soils by installing lime cement columns prior to placing an MSE wall, thus reducing the magnitude of settlement within the stiffened foundation soils (Farnsworth et al., 2008). Ladd's (1989) method was used to develop the surcharge design for the MSE walls and earthen embankment construction. For this, it was important to 20 determine the thickness of the compressible layer(s) and to estimate the reduced rate of secondary compression for the underlying sediments as a function of the amount of surcharge and the surcharge duration. The former was determined from field investigations (i.e., soil boring and CPT soundings) at various locations along the project, and the latter soil properties were evaluated from a laboratory test program using "undisturbed" samples obtained from the field investigation program (Ng, 1998). The laboratory testing was done in a relatively rigorous manner for the project because C. C. Ladd was retained by Woodward-Clyde Consultants as a senior consultant and reviewer for the project. The consolidation and shear strength testing to support the design were done at Massachusetts Institute of Technology (MIT) soil mechanics laboratory under the supervision of C. C. Ladd and the results were reported in Ng (1998). While the evaluations of Ng (1998) were being performed and finalized, the project team used interim values of Cα /CR equal to 0.0425 from preliminary laboratory testing for the Lake Bonneville deposits performed by Woodward-Clyde Consultants (Saye and Ladd, 2000). Estimates of CR values were back-calculated using soil models developed in M.S. Excel spreadsheets for the Lake Bonneville deposits. To calibrate the CR values used in these models, foundation settlement versus time data were used as obtained by the Utah Department of Transportation (UDOT) from the original I-15 embankment construction records from the northern part of Salt Lake Valley. To constrain the layer thickness used in these models, soil layering was developed from borehole logs at the corresponding locales as obtained from baseline geotechnical investigations 21 performed just prior to the I-15 Reconstruction Project by various geotechnical consulting firms. As the MIT report (Ng, 1998) became available, the average Cα/CR of 0.0433 was adopted based on laboratory testing by Woodward-Clyde Consultants (WCC) and MIT for the Lake Bonneville deposits, see Figure 2.6. In addition to this design chart, the effects of AAOS on Cα′ and the time of initiation of secondary compression, ts, were needed to complete the design. For the AAOS versus Cα′/Cα relation (Figure 2.7 top), the maximum reduction line (i.e., bottom solid line surrounded by black dots) was used. This was selected because it was based on site-specific samples obtained from the I-15 Project as tested by Ng (1998). Figure 2.7 (bottom) quantifies this the time delay between surcharge removal, tr, and start of secondary compression, ts, as a function of Figure 2.6 Relationship between rate of secondary compression and compression ratio for Lake Bonneville clays. MIT data are labeled by MIT and Woodward-Clyde Consultants are labeled WCC (Saye and Ladd, 2000, unpublished). 22 Figure 2.7 Plots showing the relationship of AAOS on Cα′ and the time of initiation of secondary compression, ts. (Top) Cα'/ Cα as a function of AAOS (Bottom) Log(ts/tr) ratios as a function of AAOS (Saye and Ladd, 2000, unpublished). 23 AAOS based on testing done by Ladd (1989) and Ng (1998). This figure shows that the start of secondary compression has greater delay for higher amounts of surcharge. Such delay is beneficial in reducing the amount of secondary compression over a given postconstruction period. In regards to ts/tr versus AAOS, the average line was selected for design purposes for the I-15 Reconstruction Project (Saye and Ladd, 2000). 24 3 RESEARCH OBJECTIVES AND TASKS 3.1 Research Objectives The primary objectives addressed in this research are as follows: 1) corroborate Mesri's concept of secondary compression (i.e., Cα/CR is relatively constant) for the Lake Bonneville deposits along the Wasatch Front in Utah, 2) supplement and/or revise, as necessary, the design relationships developed by Ng (1998) for the I-15 surcharge design using a larger set of field and laboratory test data, (3) recommend an appropriate laboratory testing and evaluation program to support project-specific surcharge design for future highway embankment projects sponsored by the Utah Department of Transportation (UDOT) in the Wasatch Front Area. 3.2 Research Plan To accomplish these research objectives, a field investigation and collection of undisturbed samples of soils in the Wasatch Front area will be performed. The specimens acquired during the field investigation will be tested to develop design charts consistent with the design parameters required to implement Ladd's (1989) method. In addition, the data acquired from this research will be evaluated and compared with existing data and relations developed by Ng (1998). 25 3.3 Tasks The major tasks needed to achieve the above research objective are: 1) Review of the existing literature regarding secondary compression and how a surcharging program can be implemented to reduce secondary settlement. This will include a description of Ladd's and Mesri's methodologies. This has been completed and is summarized in Chapter 2. 2) Obtain undisturbed samples from 4 locations located along the Wasatch Front, Utah including: (1) 400 South Street in Salt Lake City, (2) Provo South Interchange, (3) Springville 400 South Overpass Structure, and (4) Layton Interchange. The undisturbed samples will be obtained using mud rotary drilling and piston sampling at sites where long-term monitoring of settlement has been ongoing as part of instrument arrays sponsored by the UDOT. Soil samples from the area near these arrays will be used in the laboratory test program. 3) Develop and implement a laboratory test program to acquire secondary compression consolidation data and the design parameters associated with Ladd's (1989) method. The undisturbed samples and the associated laboratory tests will be performed on fine-grained, cohesive soils to determine design parameters such as σp, CR, RR, Cα, and Cα'. 4) Evaluate the design parameters obtained from the laboratory testing program and compare them with those published by Ng (1998). Specifically, the parameters compared will be the Cα/CR ratio of Ng (1998) and the normalized Cα'/Cα and log(ts/tr) versus AAOS. 26 5) Make recommendations regarding the implementation of laboratory testing program and the steps and procedures required to implement a site-specific surcharge design for future UDOT embankment projects founded on soft soil sites. 27 4 FIELD INVESTIGATIONS 4.1 Introduction A field investigation was performed for the purpose of collecting samples of fine-grained soils along the Wasatch Front at sites where long-term settlement data were available. These sites were selected because long-term instrumentation and settlement monitoring had been performed at these sites over the past 10 years. These undisturbed soil samples were tested in the U of U Civil and Environmental Engineering Soil Mechanics laboratory to determine soil properties and design parameters required for engineering evaluations such as σp, CR, RR, Cα, and Cα'. 4.2 Field Investigations The four drilling sites were: (1) 400 South at 400 South and 800 West just east of I-15 in Salt Lake City Utah (see Figure 4.1) (2) South Layton at Layton Parkway and Main Street in South Layton Utah. This site is in an empty lot northeast of the Layton Parkway and Main Street intersection west of I-15 (see Figure 4.2). (3) Springville at 400 South and just south of 750 East in Springville Utah. This site is west of the railroad tracks and on the south side of the railroad tracks overpasses (see Figure 4.3). (4) Provo at the University Avenue I-15 southbound on ramp in Provo Utah. This site is at the point where the 28 Figure 4.1 400 South site (400 South 800 West). (Top) Vicinity map for drilling site (Bottom) Close up of drilling site. 29 Figure 4.2 South Layton site (Layton Parkway and Main Street). (Top) Vicinity map for drilling site (Bottom) Close up of drilling site. 30 Figure 4.3 Springville site (400 South and 1500 West). (Top) Vicinity map for drilling site (Bottom) Close up of drilling site. 31 southbound on ramp for University Avenue and the on and off ramp for southbound traffic for 1860 South meet see Figure 4.4. For longitude and latitude, depths, and drilling dates, see Table 4.1 The drilling was performed with the use of a truck-mounted CME 75 drill rig using mud rotary drilling in a 4-inch casing. The primary purpose of the drilling was to obtain piston samples from the cohesive, fine-grained soils of Lake Bonneville and recent lacustrine deposits. The piston sampling used standard galvanized-steel Shelby tubes with a 2.8 inch inner diameter and a 3.0 inch outer diameter, with an overall length of 30 inches. The depths selected for soil sampling were determined using CPT logs that were performed by others at or near the locations of the boreholes. The locations of the CPT soundings are shown in Table 4.2. For the CPT logs, see Figures 4.5 to 4.9. The samples were immediately logged and labeled by location and depth. The logs for the boreholes are given in Figures 4.10 to 4.14. Samples were sealed with plastic Shelby caps and wrapped with duct tape to maintain their in situ moisture content. They were carefully transported and stored in a humidified room in the University of Utah Concrete Laboratory until the consolidation testing could be performed. 32 Figure 4.4 Provo site (Southbound I-15 on ramp for University Avenue). (Top) Vicinity map for drilling site (Bottom) Close up of drilling site and CPT location. 33 Table 4.1 Boring locations, depths, and drilling dates # Site Latitude Longitude BH Depth (ft.) Drilling Date 1 400 South 40°45'40.02"N 111°54'45.61"W 92 12/5/2012 2 S. Layton 41° 3'23.25"N 111°57'47.63"W 142 2/5/2013 3 Springville 40° 9'39.78"N 111°38'18.43"W 129 4/1/2013 4 Provo 40°12'26.41"N 111°39'39.43"W 127 4/5/2013 Table 4.2 CPT Locations Site CPT Latitude Longitude 400 South 06-SC-159 40° 45' 39.06"N 111° 54' 49.02"W S. Layton CPT-01 41° 3' 16.42"N 111° 57' 47.63"W Springville CPT-01 40° 9' 39.89"N 111° 38' 20.03"W Springville CPT-07 40° 9' 39.90"N 111° 38' 17.75"W Provo CPT-07 40°12' 25.08"N 111° 39' 39.30"W 34 Figure 4.5 CPT for 400 South 06-SC-159. 35 Figure 4.6 CPT for South Layton CPT-01. 36 Figure 4.6 Continued. 37 Figure 4.6 Continued. 38 Figure 4.7 CPT for Springville CPT-01. 39 Figure 4.7 Continued. 40 Figure 4.7 Continued. 41 Figure 4.8 CPT for Springville CPT-07. 42 Figure 4.8 Continued. 43 Figure 4.8 Continued. 44 Figure 4.9 CPT for Provo CPT-07. 45 Figure 4.9 Continued. 46 Figure 4.9 Continued. 47 Figure 4.10 Test Hole Log for 400 South. By Zach GibbsElevation (ft)Sample Depth (ft)Visual Soil DescriptionSample Recovery (in)Soil SymbolsPenetration N (blows/ft)Remarks and raw SPT data4196- No samples taken418510-12SAND, medium grained, gray to green gray (SP)2420- 5 9 9 1112.5-14.5CLAY, silty, gray (CL)240- 0 weight of hammer15-17SILT, soft, gray (MH)24ST17.5-19.5SILT, soft, gray (MH)24ST417520-22SILT, soft, gray (MH)24ST- sluff sand had to clean out the hole22.5-24.5CLAY, silty, gray (CL)24ST25-27CLAY, silty, gray (CL)24ST- Inter beds27.5-29.5CLAY, silty, gray (CL)24ST416530-32CLAY, silty, gray (CL)24ST32.5-34.5SILT, soft, gray (ML)24ST35-37No recovery0ST- probably sands37.5-39.5CLAY, silty, gray (CH)24ST415540-42CLAY, silty, gray (CH)24ST42.5-44.5CLAY, silty, gray (CL)24ST45-47CLAY, silty, gray (CL)24ST47.5-49.5CLAY, silty, gray (CL)24ST414550-52CLAY, silty, with sand, gray (CL)24ST52.5-54.5CLAY, silty, with sand, gray (CL)24ST55-57No recovery014- put a spt after no recovery 5 4 5 9413560-62CLAY, silty, with sand, gray (CL)24ST65-67CLAY, silty, gray (CL)24ST412570-72SAND, silty (SM)24ST411580-82SAND, clayey, fine grained (SC)24ST410590-92CLAY, gray (CL)24STSoil SymbolsOther SymbolsDriller :BedkeSand:Boring Number :400 SouthSilt:Date Drilled :12/5/2012Clay:Job Number :1Salt Lake City, Ut.400 SouthNotes :Test Method :ASTM D 1586Automatic Trip Hammer140 lbShelby tube24 in. SamplerMud Rodery4" casingCME 75(Truck Mounted)Drilling Method :Make of Drilling Rig :Test Hole I400 SouthSite Location :Hammer Type :Sampler :48 Figure 4.11 Test Hole Log for South Layton Test Hole I. By Zach GibbsElevation (ft)Sample Depth (ft)Visual Soil DescriptionSample Recovery (in)Soil SymbolsPenetration N (blows/ft)Remarks and raw SPT data43414337.52.5-4.5SAND, silty (SM)24ST43355-7SAND, silty (SM)24ST4332.57.5-9.5SAND, silty (SM)24ST433010-12SAND, silty (SM)24ST4327.512.5-14.5CLAY (CL)24ST432515-17CLAY (CL)24ST4322.517.5-19.5CLAY (CL)24ST432020-22CLAY, sandy (CL)24STSoil SymbolsOther SymbolsDriller :BedkeSand:Boring Number :South LaytonSilt:Date Drilled :2/5/2013Clay:Job Number :2South Layton Ut.Layton Parkway and Main streetNotes :Test Method :ASTM D 1586Automatic Trip Hammer140 lbShelby tube24 in. SamplerMud Rodery4" casingCME 75(Truck Mounted)Make of Drilling Rig :South LaytonTest Hole ISite Location :Hammer Type :Sampler :Drilling Method :49 Figure 4.12 Test Hole Log for South Layton Test Hole II. By Zach GibbsElevation (ft)Sample Depth (ft)Visual Soil DescriptionSample Recovery (in)Soil SymbolsPenetration N (blows/ft)Remarks and raw SPT data43414337.52.5-4.5SAND, silty (SM)24ST43355-7SAND, silty (SM)24ST4332.57.5-9.5SAND, silty (SM)24ST433010-12SAND, silty (SM)24ST4327.512.5-14.5CLAY (CL)24ST432515-17CLAY (CL)24ST4322.517.5-19.5CLAY (CL)24ST432020-22CLAY, sandy (CL)24ST429050-52SAND, clayey (SC)24ST428555-57SAND, clayey (SC)24ST426080-82CLAY (CL)24ST425585-87CLAY (CL)24ST425090-92CLAY (CL)24ST424595-97CLAY (CL)24ST4240100-102CLAY (CL)24ST4235105-107CLAY (CL)24ST4230110-112CLAY (CL)24ST4225115-117CLAY (CL)24ST4220120-122CLAY (CL)24ST4215125-127CLAY (CL)24ST4210130-132CLAY (CL)24ST4205135-137CLAY (CL)24ST4200140-142CLAY (CL)24STSoil SymbolsOther SymbolsDriller :BedkeSand:Boring Number :South LaytonSilt:Date Drilled :2/5/2013 - 2/6/2013Clay:Job Number :2South Layton Ut.Layton Parkway and Main streetNotes :BH II is located 5 ft. north of BH ITest Method :ASTM D 1586Automatic Trip Hammer140 lbShelby tube24 in. SamplerMud Rodery4" casingCME 75(Truck Mounted)South LaytonTest Hole IISite Location :Hammer Type :Sampler :Drilling Method :Make of Drilling Rig :50 Figure 4.13 Test Hole Log for Springville. By Zach GibbsElevation (ft)Sample Depth (ft)Visual Soil DescriptionSample Recovery (in)Soil SymbolsPenetration N (blows/ft)Remarks and raw SPT data4520- No samples taken450415-17SILT, sandy (ML)24ST- little sandy449920-22No recovery0ST448930-32CLAY, silty (CL)24ST448435-37CLAY, silty (CL)24ST447940-42CLAY, silty (CL)24ST446950-52CLAY, silty (CL)24ST446455-57CLAY, silty (CL)24ST445960-62CLAY, silty (CL)24ST445465-67CLAY, silty (CL)24ST444970-72CLAY (CL)24ST443980-82CLAY (CL)24ST443584-86CLAY, silty (CH)24442990-92SILT (ML)244418101-103SILT, sandy (ML)12- little recovery, less than half4402117-119SILT (ML)244397122-124SILT (ML)244392127-129SILT (ML)24Soil SymbolsOther SymbolsDriller :BedkeSand:Boring Number :SpringvilleSilt:Date Drilled :4/1/2013Clay:Job Number :3Springville Ut.400 South 1500 WestNotes :Test Method :ASTM D 1586Automatic Trip Hammer140 lbShelby tube24 in. SamplerMud Rodery4" casingCME 75(Truck Mounted)Make of Drilling Rig :SpringvilleTest Hole ISite Location :Hammer Type :Sampler :Drilling Method :51 Figure 4.14 Test Hole Log for Provo. By Zach GibbsElevation (ft)Sample Depth (ft)Visual Soil DescriptionSample Recovery (in)Soil SymbolsPenetration N (blows/ft)Remarks and raw SPT data4498- No samples taken- drove casing to 10 ft (very dense)448512-14SILT, clayey (ML)24ST448017-19SILT, clayey (ML)24ST- gravels- drove casing to 40 ft- casing broke/came apart445740-42No Recovery0ST- fine sands and silts444750-52CLAY, gray (CH)24ST443760-62CLAY, silty, gray (CL)24ST442770-72CLAY, gray (CL)24ST441780-82CLAY, gray (CL)24ST440790-92CLAY, gray (CL)24ST- had artesian conditions at 90 ft.4387110-112CLAY, gray (CH)24ST4382115-117CLAY, gray (CH)24ST4372125-127CLAY, silty, gray (CL)24STSoil SymbolsOther SymbolsDriller :BedkeSand:Boring Number :ProvoSilt:Date Drilled :4/5/2013Clay:Job Number :4Provo Ut. University Ave.University Ave./ I-15 onrampNotes :Test Method :ASTM D 1586Automatic Trip Hammer140 lbShelby tube24 in. SamplerMud Rodery4" casingCME 75(Truck Mounted)Make of Drilling Rig :ProvoTest Hole ISite Location :Hammer Type :Sampler :Drilling Method :52 5 LABORATORY TESTING 5.1 Test Procedures 5.1.1 Introduction One of the outcomes of this research is to produce a recommended laboratory testing program that can be routinely executed by geotechnical consulting firms to develop surcharge evaluations and design for support of highway transportation projects. Recent research by the University of Utah (Ozer et al., 2012; Bartlett and Ozer 2005) and that of Ng (1998) has shown that the controlled rate of strain consolidation (CRS) test (D4186M-12e1) is generally preferable to standard incremental loading oedometer tests (ASTM D2435M-11) for producing high-quality laboratory data for consolidation evaluations. CRS testing better defines the shape of the consolidation curve due to the higher density of data points produced by this test, especially as the specimen transitions from recompression to virgin compression behavior. Notwithstanding, conventional incremental load oedometer tests were selected instead of CRS consolidation tests for this research. This was done in discussions with the UDOT technical advisory committee, and the reasons for this selection were: (1) incremental load tests are the standard of practice, (2) multiple secondary compression tests needed to be performed simultaneously for this research and there was not sufficient CRS devices at the University for such 53 testing, (3) it is unclear if the CRS test offers any advantage over incremental load oedometer tests when secondary compression is the primary topic of the research. Therefore, it is hoped that careful sample preparation and incremental load testing in conventional tabletop oedometers would produce data that are of sufficient quality to be incorporated in the research plan. In addition, it is hoped that if consulting engineers and technicians review and follow, as applicable, the procedures and recommendations developed herein, sufficient data quality and quantity can be obtained to support future surcharge design and evaluation strategies for UDOT projects. The following sections describe the test procedures that were used to perform the laboratory testing that supports this research. 5.1.2 Testing equipment The equipment used for the laboratory test program were tabletop oedometers (Figure 5.1) located in the University of Utah Soil Mechanics Laboratory. These devices consisted of a 1.0-inch high and 2.5-inch diameter stainless steel consolidation ring, a standard consolidation cell with base plate having an O-ring, two 2.5-inch porous stones, a top reservoir, a top cap to provide evenly dispersed pressure on the sample, and three bolts to hold the cell together (Figure 5.2). Filter paper was used between the porous stones and the soil sample to prevent plugging of the stones. Suspended weights, which produced a reaction on the sample equivalent to 0.25, 0.5, 1.0, and 2.0 tsf, were 54 Figure 5.1 Tabletop oedometer. Figure 5.2 Consolidation ring and cell. 55 used to load the specimens. These weights were used in various combinations to produce the loading increments discussed in Section 5.2. The dial gages used for the vertical settlement readings had a precision of 0.0001 inch. 5.1.3 Sample Setup All Shelby tubes collected from the sites remained sealed and were stored in a humidified room in the University of Utah Concrete Laboratory to preserve their initial water content. The Shelby tubes were cut with a commercial ban saw (Figure 5.3). The sample was cut approximately 1 inch above and below the test specimen to provide sufficient soil for trimming and moisture content testing without producing a lot of waste. Ng (1998), Ladd (1999), and Bartlett and Ozer (2005) have found that radiography can serve as an aid to detecting variations in the soil fabric (e.g., heterogeneity, layer, anomalies, disturbance), if desired, but radiography was not Figure 5.3 Horizontal ban saw used to cut samples for testing. 56 performed on specimens used for this research. In addition, prior to extrusion of soft samples, it is recommended that a piano wire be inserted between the outer edge of the specimen and the cut Shelby tube and carefully used to cut or break the perimeter adhesion bond between the tube and the specimen to reduce sample disturbance (Ladd et al., 1998). The samples were carefully extruded from the tubes using the sample extruder shown in Figure 5.4. Immediately following extruding, they were trimmed to fit the consolidation ring using a turntable (Figure 5.5), a fine-gauge wire saw, and reference straight edge. Soil trimmings were immediately weighed to prevent change in mass due to drying and were subsequently placed in the drying oven for moisture content determinations. It is important to obtain accurate moisture content measurements of the specimens for determination of the initial void ratio, which is required in the consolidation calculation, and also as an index of compressibility. Regarding this latter point, Bartlett and Lee (2004) have shown that the compression index of the soil, Cc, can be reasonable estimated from the in situ moisture content of the soil, because moisture content is highly correlated with void ratio, which is in turn correlated with soil compressibility. Because of this, it is recommended that moisture content and other index properties be obtained for the specimens for further correlation with laboratory-determined consolidation properties. The height, weight, and diameter of the consolidation ring were recorded using precision scales and a micrometer. The consolidation rings used had 57 Figure 5.4 Extruding sample from Shelby tube. . Figure 5.5 Sample in turntable. 58 heights of 1 inch and diameters of 2.5 inches. To reduce the friction between the soil and the consolidometer rings, the inner circumference of the rings were lubricated with a low-friction, silicone-based lubricant. The samples were then carefully placed into the consolidation ring and trimmed flush with the top and bottom of the ring (Figure 5.6). If there were any small voids present on the top or bottom of the specimens from the trimming process, these imperfections were carefully filled with trimmed soil. After the trimming and preparation were completed, the weights of the rings with the soil present were then recorded. Porous stones were boiled in demonized water and soaked in de-aired water prior to assembly. Standard 0.15-mm thick filter paper was used between the sample and the porous stones on both top and bottom of the specimens to prevent clogging of the stones and the loss of the solids during the tests. The samples in the consolidation ring were then assembled in the cell with porous stones and filter paper, and the top caps were placed on the top reservoir and secured with bolts to prevent leakage (Figure 5.7). The cell were then placed in tabletop oedometers and deionized, de-aired water was used to fill the reservoir to saturate the specimen (Figure 5.8). Testing was then performed using the interpretive methodologies described in the following section. 5.2 Laboratory Test Program The specimens selected for testing were determined by evaluating nearby CPT soundings and selecting the sampling depth having the most clay-like behavior. This was usually manifest on the adjacent CPT sounding as low tip 59 Figure 5.6 Sample in consolidation ring. Figure 5.7 Sample fully assembled in cell. 60 Figure 5.8 A fully assembled cell in table top oedometer. resistance and relatively high excess pore water pressure relative to hydrostatic conditions. The CPT sounding used in the evaluations are found in Chapter 4. At the 400 South site, a CPT sounding was not available; hence, the sampling depths started at 15 feet (i.e., below the alluvium) and sampling was done every 5 feet until a depth of 50 feet was reached. All of these samples from 400 South appeared to be very similar to each other and were cohesive sediments. For each sampling interval, the specimens were used to determine the following consolidation properties: (1) preconsolidation stress (σp), (2) rate of secondary compression for normally consolidated specimen, Cα, (3) rates of secondary compression for overconsolidated specimens, C'α, for OCR values of 61 1.25, 1.5, and 2.0. The preconsolidation stress was determined using 1-D incremental loading tests with time rate of consolidation measurements taken for each loading increment. The following describes the procedures used to determine these properties. 5.2.1 Determination of the Preconsolidation Stress The preconsolidation stress of the specimens was evaluated using a 1-D incremental loading consolidation test with the sample preparation described in section 5.1. The following loading schedule was implemented: 0.25, 0.5, 1.0, 2.0, 4.0, 8.0, and 16.0 tsf (500, 1000, 2000, 4000, 8000, 16000, and 32000 psf, respectively), and the unloading schedule was 16.0, 4.0, 1.0, and 0.25 tsf (8000, 2000, and 500 psf, respectively). This is equivalent to applying a load increment ratio, ΔP/P, of unity (1.0), which is most commonly used in practice. However, Jamiolkowski et al. (1985) recommend that the ΔP/P ratio be reduced to about 0.5 to obtain better defined consolidation curves in the vicinity of the preconsolidation stress as used by Canadian practice for sensitive clays in eastern Canada. In standard geotechnical practice, each load increment is usually maintained for one day (24 hours) to define the consolidation curve and estimate the preconsolidation stress. However, because the end of primary consolidation usually occurs in less than 1 hour after the newly applied load, the virgin portion of consolidation curves based on 24-hour waiting periods is displaced downward by one or more cycle of secondary compression (Jamiolkowski et al., 1985). This is not desirable for two obvious reasons: (1) it makes interpretation of the 62 preconsolidation stress more variable, especially for soft soils, (2) data needed for interpretation of the rate of secondary compression at the applied stress are not obtained. Thus, contrary to standard practice and in order to produce as little as possible secondary compression between each incremental loading, time rate of consolidation tests were done for every loading increment. These data were used to decide when end of primary consolidation was essentially completed so that the next load increment could be applied. To this end, Taylor's square root of time method (as found in Holtz and Kovacs 2011) was performed on the time rate of consolidation data to determine when 90% of primary consolidation had occurred. To accelerate the consolidation process 90% of primary consolidation was selected; when consolidating soft clays, it can take substantially more time to reach 95 or 100% of consolidation. When this value was reached, the next load increment was then applied. This accomplished three things: (1) it allowed for the incremental loading test to be completed in a more expeditious manner similar to that of CRS testing, (2) it allowed for a more consistent interpretation of the preconsolidation stress, and (3) it produced more repeatable results for the rate of secondary compression. The incremental loading results were then plotted and the pre-consolidation stress (σ'p) was determined using the work/strain method (Becker et al. 1987) and Casagrande's method (Casagrande 1936). The compression ratio (CR) and recompression ratio (RR) were also determined from plots of log of applied stress versus vertical strain. 63 5.2.2 Determination of the Rate of Secondary Compression Cα The rate of secondary compression varies with the preconsolidation stress and amount of aging (Jamiolkowski et al., 1985; Ng 1998). For this research, it is important to determine rate of secondary compression for the normally consolidated condition, i.e., Cα. This was done by first determining the preconsolidation stress from the incremental load tests described in the previous section. After this, specimens from the same depth interval and borehole were loaded to a new stress state that was 1.5 to 2.0 times the in situ σ'p value. This ensured that specimens had reached a new normally consolidated state and any effects of aging or past preconsolidation had been removed. The method of Ladd (1989) was used to interpret the time rate of consolidation data for secondary compression (Figure 2.3). When using Ladd's methodology, the value of secondary settlement Cα is the slope of the line through the linear most portion of the data, after primary consolidation has occurred, on the strain vs. log of time plot. The reading schedule used for this part of the test was, 4 min, 8 min, 15 min, 30 min, 1 hour, 2 hour, and 4 hours. Then readings were taken about once a day for the remainder of the test, usually about once every 24 hours. The first few readings were removed because they have very little effect on the value of Cα, it was decided to take the first reading at 4 minutes to increase the number of tests that could be ran at one time. This test ran for 1 to 2 weeks to be sure that a good value of Cα is achieved. 64 5.2.3 Determining C'α The process for determining C'α was done using the same procedure to determine Cα but with a few variations. The specimen is loaded to a state of stress that is 1.5 to 2.0 times the in situ σ'p value as was done in the section stated above. After the 1 or 2 hour reading was taken, the load was then reduced to a known OCR of ether 1.25, 1.5, or 2.0.These values were selected because this range is likely to bracket the values used for surcharge design. The reading schedule for this part of the test is the same as stated above and will run for 1 to 2 weeks. The method of Ladd (1989) was also used to interpret the time rate of consolidation data for secondary compression. Additional detail was provided in Chapter 2. 65 6 RESULTS AND INTERPRETATIONS 6.1 Lab Tests and Data Screening In the course of this research, a total of five 1-D consolidation tests were performed at each sampling depth from the individual boreholes completed during the field investigations. For each depth, one test was performed to determine the preconsolidation stress, and the remaining four tests consisted of time rate of consolidation tests to determine the rate of secondary compression at OCR values of 1.0, 1.25, 1.5, and 2.0. These OCR values were selected to represent a reasonable range of overconsolidation states that could be effectively achieved in the foundation soils during embankment construction. At the 400 South Street site in Salt Lake City, Utah, specimens from eight sample depths were tested. At the South Layton, Utah, site located just off of Layton Parkway and Main Street a total of four sample depths were tested. At the Springville, Utah, 400 South Street site, specimens from seven sample depths were tested. Similarly at the Provo, Utah, interchange, specimens from seven sample depths were tested. Thus, in total twenty-six consolidation tests and one hundred and four time rate tests were performed on Pleistocene and recent fine-grained, cohesive, lacustrine deposits comprised of Lake Bonneville and more recent clays, most likely of Utah Lake origin. A list of the locations, sample depths and results of these tests are presented in Table 6.1. 66 Table 6.1 Listing the sites, depths, moisture content, pre consolidation stress, compression and recompression ratio, and the rate of secondary settlement at different OCR's Site Avg. depth (ft) ω (%) Casagrande σ'p (psf) Work σ'p (psf) CR RR Cα Cα' OCR 1.25 Cα' OCR 1.50 Cα' OCR 2.00 400 South 16b 53.0 2671 2780 0.21 0.021 0.0164 0.0230 0.0017 0.0009 21 52.0 6338 6420 0.20 0.026 0.0127 0.0089 0.0044 0.0020 26 49.2 2989 3040 0.18 0.021 0.0095 0.0072 0.0028 0.0007 31 47.2 5473 5280 0.15 0.018 0.0066 0.0022 0.0012 0.0003 39 40.3 5423 5520 0.15 0.024 0.0105 0.0058 0.0015 0.0010 41a 46.7 1610 1660 0.13 0.016 0.0103 0.0103 0.0029 0.0010 46a 40.3 4262 4240 0.12 0.018 0.0036 0.0047 0.0026 0.0010 51a 48.5 2350 2120 0.17 0.016 0.0049 0.0047 0.0041 0.0018 S. Layton 16 32.8 5067 5120 0.13 0.012 0.0030 0.0017 0.0007 0.0004 91 34.7 6510 6580 0.07 0.013 0.0026 0.0015 0.0006 0.0005 106 28.4 6408 7140 0.04 0.008 0.0023 0.0018 0.0011 0.0005 131 25.5 5974 6320 0.11 0.017 0.0048 0.0033 0.0012 0.0006 Springville 31 26.7 7101 7060 0.09 0.023 0.0018 0.0012 0.0006 0.0003 41 26.8 6682 6780 0.09 0.014 0.0021 0.0011 0.0003 0.0002 66 23.6 5810 6040 0.05 0.008 0.0024 0.0022 0.0014 0.0010 71 27.7 12125 12420 0.12 0.018 0.0040 0.0026 0.0012 0.0012 76 32.8 6274 6280 0.15 0.022 0.0051 0.0050 0.0019 0.0007 81a 41.9 6643 6880 0.19 0.023 0.0074 0.0060 0.0016 0.0007 85a 40.6 5503 5503 0.17 0.023 0.0047 0.0025 0.0006 0.0001 Provo 13 45.5 7358 7820 0.11 0.017 0.0046 0.0023 0.0007 0.0004 18 35.2 2969 3120 0.14 0.013 0.0036 0.0021 0.0008 0.0016 51a 51.1 3803 4000 0.12 0.020 0.0037 0.0014 0.0010 0.0004 61a 32.6 4740 4800 0.10 0.017 0.0032 0.0026 0.0015 0.0009 81 47.8 6331 6340 0.18 0.023 0.0075 0.0027 0.0018 0.0018 91 34.1 7014 7300 0.13 0.015 0.0045 0.0040 0.0007 0.0005 111a 41.3 5512 5520 0.14 0.021 0.0037 0.0030 0.0008 0.0005 a. Data that has been removed from analysis due to a high SQD value. b. Data was removed from analysis due to poor result for Cα by not running long enough. 67 Prior to developing the consolidation parameters from the specimens, the consolidation test results were screened using the sample quality designation (SQD) developed by Andresen and Kolstad (1979) of the Norwegian Geotechnical Institute (NGI). This method uses the recompression vertical strain during the initial reloading loading of the specimen back to the in situ effective vertical stress. For example, in this method, a SQD value of 4 indicates that the specimen underwent 4 % vertical strain during the reloading. Andresen and Kolstad (1979) developed a SQD nomenclature that corresponds with the vertical strain values given in Table 6.2. For example, a SQD value between 2 to 4 receives a "fair" designation. Saye and Ladd (2000) used SQD values of 4, or greater, to screen out (i.e., omit) consolidation data from their evaluations. The 1D consolidation data used by these authors were obtained by the various geotechnical firms during the baseline investigations for the I-15 Reconstruction project and had varying levels of quality. Therefore, in general, based on Andersen and Kolstad (1979) and Saye and Ladd (2000), it is recommended that a SQD criterion of 4, or higher, be used as a screening criterion for future project evaluations. Table 6.2 Values of strain at σ'vo and the corresponding rating of SQD Strain on Reloading to σ'vo (%) Sample Quality Designation (SQD) <1 A, Very good to excellent 1-2 B, Good 2-4 C, Fair 4-8 D, Poor >8 E, Very poor 68 However, some of the specimens obtained for this research exceeded this recommended screening threshold (Table 6.3). In order to improve the sample size for the statistical analysis herein, it was decided to slightly relax this criterion. Thus, consolidation tests results having SQD greater than 6 were excluded from the subsequent analysis herein. The footnotes in Table 6.1 give information on which specimens were screened from the statistical analyses due to poor SQD, or other testing or data reduction issues. Some of the specimens sampled from the 400 South Street site in Salt Lake City, Utah and from the Provo, Utah South Interchange site had relatively high SQD values indicating higher amounts of sample disturbance. The reasons for this are unclear, but may be partly attributable to the softer soil deposits found at these locations. Such soils may be more susceptible to disturbance effects associated with sampling, handling, and preparation processes. The consequences of sample disturbance appear to have had a larger impact on specimens tested at or slightly above the normally-consolidated state of stress (i.e., OCR ≈ 1). For example, when the screened tests results for Cα and Cα′ from this research are plotted versus OCR values (Figure 6.1), the variance (i.e., scatter) of the data is higher for results obtained at lower OCR values when compared with that obtained at higher OCR values. This holds true for both the total variance (scatter of data from all sites) and for the sample variance (scatter of data obtained from a particular site). Some of this variation is due to natural variability of the layered sediments at a given site, because not all specimens at each site were obtained from the same layer. However, it is also likely that some 69 Table 6.3 Site, Depth, Effective Vertical Stress, and SQD Site Avg. Depth (ft) σ'v (psf) SQD 400 S 16 1198 2.7 21 1436 2.7 26 1674 5.5 31 1912 4.7 38.5 2269 5.6 41 2388 11.2 46 2626 8.2 51 2864 11.8 S. Layton 16 1483 3.0 91 5803 5.0 106 6667 3.3 131 8107 7.0 Springville 31 2067 2.8 41 2593 4.3 66 3908 3.0 71 4171 2.4 76 4434 5.2 81 4697 8.9 85 4908 6.5 Provo 13 1245 1.4 18 1508 3.2 51 3244 17.5 61 3770 8.1 81 4822 7 91 5348 6.4 111 6400 13.5 70 Figure 6.1 Plot of Cα and Cα′ vs. OCR for 400 South, South Layton, Springville, and Provo Sites of the variation can be attributed to the effects of sample disturbance. Therefore, it is reasonable to conclude that the apparent decrease in variance with increasing OCR values is, in part, due to the ameliorating effect of overconsolidation of the soil specimens prior to performing the time rate of consolidation tests to determine Cα′. A similar beneficial effect of overconsolidating soils prior to performing undrained shear strength tests have been discussed by Ladd and Foott (1974) in developing the SHANSEP (Stress Cα = 0.1189e-2.55OCR SD=0.0023, N=16, R² = 0.7569 Cα = 0.0145e-1.673OCR SD=0.0023, N=20, R² = 0.4991 Cα = 0.0226e-1.759OCR SD=0.0013, N=16, R² = 0.5951 Cα = 0.0179e-1.844OCR SD=0.0008, N=16, R² = 0.8287 Cα = 0.0276e-1.937OCR SD=0.0021, N=68, R² = 0.5557 0.000 0.002 0.004 0.006 0.008 0.010 0.012 0.014 1.0 1.2 1.4 1.6 1.8 2.0 Cα and C'α OCR OCR vs. Cα and C'α 400 South Springville Provo S. Layton Expon. (All)71 History and Normalized Soil Engineering Properties) method. A plot of preconsolidation stress versus depth for the higher quality samples shows that the soils at the various test sites are overconsolidated at all depths (Figure 6.2). This is a typical finding which has been documented by many geotechnical investigations for the surficial alluvium and underlying lacustrine sediments in the Wasatch Front Area. The apparent overconsolidation originates from aging, void ratio change due to repeated drying and wetting Figure 6.2 Plot of preconsolidation stress vs. depth for 400 South, South Layton, Springville, and Provo. 0 10 20 30 40 50 60 70 80 90 100 110 120 130 0 5000 10000 15000 Depth (ft) σ'p (psf) Preconsolidation Stress vs. Depth 400 South Springville Provo S. Layton 400 S. σ'v (estimated) Springville σ'v (estimated) Provo σ'v (estimated) S. Layton σ'v (estimated) 72 cycles resulting from ground fluctuations and in some cases, minor cementation from calcium carbonate. The effective vertical stress profile is also shown on Figure 6.2 using an estimation of the water table level determined from piezometer readings for CPT soundings performed nearby. The CPT soundings were used to infer the equilibrium water table condition instead of the boreholes because the drilling operations required that the boreholes be "abandoned" (filled with grout) soon after the drilling had ceased. The in situ OCR values range from 1.3 to 5.8, where the higher values of OCR tend to be at shallow depths and then decreases with depth. 6.2 Relationships for Cα, Cα′, and Cα′/Cα The plot Cα and Cα′ versus OCR shows that there is a non linear trend that exists between the dependent and independent variables. This trend is best fit using the exponential trendline feature in MS Excel. This plot and its corresponding nonlinear relationships indicate a relatively small change in Cα′ for OCR values greater than about 1.5. At higher OCR values, the Cα′ values appear to converge to a value slightly less than 0.001 (Figure 6.1). From an application standpoint, this behavior suggests that there is a point of diminishing return when surcharging soils beyond an OCR value of about 1.5 for the sediments tested in this study. While the Cα′ values for the various sites tend to converge at an OCR value of 2.0, it is also apparent that the fitted relationship for the 400 South Site in Salt Lake City is higher than the average fitted trendline; whereas the fitted 73 trendlines for Springville and South Layton tend to be lower than the average trendline. A possible reason for this difference in behavior may be attributed to nature of the lacustrine sediments. In the Salt Lake Valley, the lacustrine sediments sampled are solely from Lake Bonneville deposits, whereas at the Provo and Springville sites, in Utah Valley, the sampled sediments consisted of recent Lake Utah and earlier Lake Bonneville sediments. In general, it appears that the Utah Lake sediments are siltier than those obtained from Lake Bonneville, and this may be causing the difference in the Cα′ values. Because of these relatively large differences in the trendlines for the various sites, a better interpretation and graphical representation of the data are required. This can be done by normalizing Cα′ using Cα (i.e., forming Cα′/Cα ratios) and plotting the normalized values versus OCR (Figure 6.3). This method produces a normalized average trendline for Cα′/Cα that fits all data reasonably well which can be used as a good representation of the average of all data. Ladd (1989) introduced the concepts of amount of surcharge (AOS) and adjusted amount of surcharge (AAOS) instead of OCR to represent the data trends. These factors are more useful for applied surcharge purposes (see Chapter 2). A plot of Cα′/Cα versus AAOS is shown in Figure 6.4 Ng (1998) plotted C'α/Cα versus AAOS values from testing developed for the I-15 Reconstruction Project on a semi-log plot. The data from this study have been superimposed on the Ng (1998) relationship for comparative purposes in Figure 6.5. The average trendline for this research plots significantly higher than that developed by Ng (1998). There are several possible explanations for this, as 74 Figure 6.3 Plot of Cα′/Cα vs. OCR for 400 South, South Layton, Springville, and Provo Sites discussed in the next few paragraphs. First, the soil specimens from this research may be significantly different from those tested by of Ng (1998). The Springville and Provo sites are located several tens of miles to the south of Salt Lake City in an adjacent valley and appear to be siltier in nature due to the presence of near shore sediments deposited in the Utah Lake. Only one test site (400 S. Street in Salt Lake City) C'α/Cα = 12.758e-2.566OCR SD=0.15, N=12, R² = 0.9031 C'α/Cα = 5.106e-1.673OCR SD=0.19, N=15, R² = 0.6959 C'α/Cα = 4.6454e-1.759OCR SD=0.20, N=12, R² = 0.6293 C'α/Cα = 5.8783e-1.844OCR SD=0.14, N=12, R² = 0.8649 C'α/Cα = 6.5278e-1.955OCR SD=0.17, N=51, R² = 0.7473 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.0 1.2 1.4 1.6 1.8 2.0 C'α/Cα OCR OCR vs. C'α /Cα 400 South Springville Provo S. Layton Expon. (400 South) Expon. (Springville) Expon. (Provo) Expon. (S. Layton) Expon. (All)75 Figure 6.4 Plot of AAOS vs. Cα'/Cα using an average exponential trend line was geographically near the drill hole locations where Ng (1998) and Woodward-Clyde Consultants obtained soil samples for the I-15 project. When the results for the 400 South Street site (blue line in Figure 6.5) are compared with those of Ng (1998), significant differences still remain. However, the lower bound data points from the 400 South Site plot within the upper range of the Ng (1998) data. C'α/Cα = 0.9245e-0.02AAOS SD=0.18 R² = 0.7473 N=68 0.0 0.2 0.4 0.6 0.8 1.0 0 20 40 60 80 100 C'α/Cα AAOS (σp-σvf)/σvf (%) AAOS vs. C'α/Cα for all sites 400 South Springville Provo South Layton76 77 Second, the trend of the Ng (1998) relationship appears to overstate the reduction in Cα′/Cα as a function of AAOS. For example, if extrapolated to an AAOS value of 50 % (i.e., OCR = 1.5), the average trendline of Ng (1998) would predict a Cα′/Cα ratio of near zero, which appears to be unlikely, especially when considering that this research shows a minimum value of Cα′/Cα of about 0.1 at an OCR value of 2.0. This latter result appears to be more realistic and intuitive based on the data presented in this report. Third, long-term settlement performance monitoring data obtained from the I-15 Reconstruction Project for the surcharged earthen embankments and MSE walls show that the measured creep settlement is somewhat larger than desired performance goal (Figure 6.6) (Farnsworth et al., 2008). The settlement performance goal adopted by the project was to limit the creep settlement to 75 mm, or less, in a 10-year postconstruction period. Large earthen embankments located at 400 S. and 2400 S. Streets were constructed with surcharged embankments designed to meet this performance goal. However, the 10-year post construction settlement at these sites is projected to exceed this settlement goal by a factor of about 1.5 to 2 (Figure 6.6). This suggests that the rate of secondary compression in the subsurface soils at these locales is greater than that anticipated in the surcharge design. One possible reason for the underestimation of the actual settlement could rest in the value of Cα′ selected for the design calculations. Because the amount of secondary settlement is directly proportional to Cα′ based on Equation 2-8, the additional settlement incurred at these sites may have resulted from an underestimation of the actual Cα′ for the 78 Figure 6.6 Rate of foundation creep extrapolated to 10 years of postconstruction (from Farnsworth et al., 2008). foundation settlement for these locales. For example, if the design Cα′ values were approximately 1.5 to 2.0 times higher than those reported by Ng (1998), then such a change would produce a more reasonable result that is in better agreement with the average Cα′/Cα trendline developed for the 400 South street site in Salt Lake City from this research. Nonetheless, despite the various interpretations of the existing laboratory and field data that could be offered, it is clear that site-specific field and laboratory evaluations are needed for future sites to avoid pitfalls associated with applying data and relationships developed from other sites that may have significantly differing soil conditions than the site of interest. 79 6.3 Cα/CR Ratio Mesri has shown that the Cα/CR ratio is relatively constant for a given soil type. Knowledge of this ratio has proven to be very helpful in performing secondary settlement calculations because Cα can be estimated if values of CR and the Cα/CR ratio are known for a particular soil or can be reasonably estimated. Values of CR easily attainable from standard consolidation testing and Cα/CR ratios can be estimated from this research and that of Ng (1998), as shown in Figures 6.7 and 6.8, respectively. The ratios developed by this research (Figure 6.7) were calculated using a linear trendline function and by forcing the trendline through the origin; hence, the slopes of these lines also represent the Cα/CR ratio. These results show that each individual research site has a slightly different Cα/CR relationship when compared with the average trendline. In short, the 400 South street site in Salt Lake City has a somewhat steeper slope (higher ratio) than the South Layton, Springville, and Provo sites. However, when the average slope of all the data from the four research sites is calculated, the corresponding value is Cα/CR = 0.0442. This average ratio correlates reasonably well with that of Ng (1998) of Cα/CR = 0.0433. The Ng (1998) average relation included consolidation tests performed at MIT and by Woodward-Clyde Consultants for the Lake Bonneville deposits (Figure 6.8). Although this research supports a similar average Cα/CR ratio when compared with Ng (1998), the results for the 400 South street site in Salt Lake City plot somewhat above the average trendline of Ng (1998). In addition to this, the time delay for when secondary compression80 81 82 resumes is shown in Figure 6.9 as a plot of AAOS vs. LOG (ts/tr). When comparing the data from this research and the data from MIT and WWC (Ng, 1998), the trendline is slightly lower but compares well with previous work. 6.4 Moisture Content Correlations Many researchers have shown that the moisture content, ω, of soil is highly correlated with soil compressibility (i.e., Cc and CR) for saturated soils. This is because when the soil fabric is saturated, the in situ moisture content is directly correlated with the in situ void ratio for soils with a given specific gravity. For many cohesive soils, there is a relatively minor variation in the specific gravity of the soil solids; hence, moisture content is an excellent predictor of void ratio. In addition, void ratio in turn is highly correlated Cc and CR because soils with high voids have more opportunity for compression (i.e., void ratio reduction) upon loading. Figure 6.10 shows a ω versus CR relation using the test results from this research. The data indicate a relatively good correlation between these properties. In addition to this, Bartlett and Lee (2004) developed moisture content and compressibility correlations for the Lake Bonneville deposits. These correlations were made from laboratory data obtained from various geotechnical reports associated with the I-15 Reconstruction Project. Test results with a SQD value greater than 4 were screened (excluded) from their evaluations (Figure 6.11). The data from this research have been superimposed on the Bartlett and Lee (2004) plot for comparative purposes. The trendline developed from this research plots somewhat lower than that of Bartlett and Lee (2004). However, 83 84 Figure 6.10 Plot of moisture content vs. virgin compression ratio this does not imply that the two equations are inconsistent for the following reasons: (1) The soils from this research appear to be somewhat siltier, on average, than those used by Bartlett and Lee (2004), (2) The Bartlett and Lee (2004) relation has more statistical support because of the larger sample size, (3) The data from this research do not plot outside the data range of the Bartlett and Lee (2004) relation, suggesting the two data sets are not entirely inconsistent. For application purposes, it is recommended the Bartlett and Lee relation be used because of its greater statistical support. In addition to the correlations just discussed, correlations that included rate of secondary compression properties and moisture content were explored. CR = 0.0036ω - 0.0072 SD= 0.03 R² = 0.5743 N=17 0.00 0.05 0.10 0.15 0.20 0.25 0 10 20 30 40 50 60 Virgin Compression Ratio, CR Moisture Content, ω (%) Moisture Content vs. Virgin Compression Ratio 400 South Springville Provo S. Layton85 Figure 6.11 Plot of moisture content vs. virgin compression ratio (Bartlett and Lee, (2004) The correlations as attempted included: Cα, Cα'/CR, Cα'/Cα as shown in Figures 6.12, 6.13, and 6.14. Based on these plots, there is poor to very poor correlation between Cα and Cα'/Cα and ω (Figures 6.12 and 6.14, respectively). However, the correlation between Cα/CR and ω has some promise for future development and application (Figure 6.13). However, more observations are needed to improve the statistical support for this relation. CR = 0.0036ω - 0.0072 SD= 0.03 R² = 0.5743 N=17 400 South Springville Provo S. Layton86 Figure 6.12 Plot of moisture content vs. rate of secondary settlement Figure 6.13 Plot of moisture content vs. Cα/CR ratio Cα = 0.0001ω - 0.0019 SD=0.0024 R² = 0.1882 N=68 0.000 0.002 0.004 0.006 0.008 0.010 0.012 0.014 0 10 20 30 40 50 60 Rate of Secondary settlement Cα Moisture Content, ω (%) Moisture Content vs. Cα 400 South Springville Provo South Layton Cα/CR = 0.0008ω + 0.0122 SD=0.012 R² = 0.2761 N=17 0.00 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0 10 20 30 40 50 60 Ratio of Cα/CR Moisture Content, ω (%) Moisture Content vs. Cα/CR ratio 400 South Springville Provo South Layton87 Figure 6.14 Plot of moisture content vs. normalized rate of secondary settlement Cα'/Cα = -0.0058ω + 0.5764 SD=0.24 R² = 0.047 N=51 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 0 10 20 30 40 50 60 Normalized rate of secondary settlement Cα'/Cα Moisture Content, ω (%) Moisture Content vs. C'α/Cα 400 South Springville Provo South Layton88 7 CONCLUSION 7.1 Summary of Thesis Objectives The main objectives in this research are 1) corroborate Mesri's concept of secondary compression (i.e., Cα/CR is relatively constant) for the Lake Bonneville deposits along the Wasatch Front in Utah, 2) supplement and/or revise, as necessary, the design relationships developed by Ng (1998) for the I-15 surcharge design using a larger set of field and laboratory test data, 3) recommend an appropriate laboratory testing and evaluation program to support project-specific surcharge design for future highway embankment projects sponsored by the Utah Department of Transportation (UDOT) in the Wasatch Front Area, and 4) develop additional design guidance and/or recommendations for developing and evaluating the surcharge design. 7.2 Mesri's Concept of Secondary Compression Based on the results of the oedometer tests run at MIT and WCC they have recommended values of the following creep properties: 1) Cα/CR 2) creep behavior as a function of AAOS 3) the delay to the amount of time before creep resumes after the removal of a surcharge as a function of AAOS. The data acquired during testing from this research are being compared with that produced by MIT and WCC. 89 7.2.1 Cα/CR An analysis from this research, the Cα/CR ratio with the regression line passing through the origin (Figure 6.8) gives an average value of Cα/CR = 0.0442. The site of 400 South Cα/CR = 0.0598, South Layton Cα/CR = 0.0346, Springville Cα/CR = 0.0308, and Provo Cα/CR = 0.0359 (Figure 6.7). The trend at the 400 S. Street site in Salt Lake City has a somewhat steeper slope (higher ratio) than the South Layton, Springville and Provo sites. The average ratio obtained though this research of Cα/CR = 0.0442 correlates reasonably well with that obtained by the results from Ng's (1998) of Cα/CR = 0.0433, thus confirming that Mesri's concept of secondary compression is constant along the Wasatch Front (i.e., the Cα/CR ratio). 7.2.2 Creep Behavior as a Function of AAOS Using the methodology developed by Ladd (1989), when plotting AAOS vs. C'α/Cα on a semi-log plot and comparing this with research done by Ng (1998), these data show that the trend seems to be higher than that estimated by Ng (1998) (Figure 6.5). In fact the trend seems to be closer to the upper bound of the data from Ng (1998). During the I-15 reconstruction project, the lower bound was used for the calculations of long-term settlement. The 10-year post-construction settlement at these sites is projected to be almost 1.5 to 2 times that of what was calculated (Figure 6.6). This suggests that the rate of secondary compression in the subsurface soils at these locales is greater than that anticipated in the surcharge design. 90 In Figure 6.5, it can be seen that with AAOS above 50% (OCR = 1.5) the average trendline of Ng (1998) would predict a Cα′/Cα ratio of near zero. The data from this research show a minimum value of Cα′/Cα of about 0.1 at an AAOS of 100 percent (OCR = 2.0) (Figure 6.5). However, the data seem to fit better with the use of an exponential trend line (Figure 6.4). Therefore, it is recommended that when estimating the value of C'α/Cα from AAOS to use the exponential equation C'α/Cα = 0.9245e-0.02(AAOS), or if using the plot produced by MIT and an AAOS below 50% to use the upper bound trend line. 7.2.3 The Time Before Creep Resumes After the Removal of a Surcharge When comparing the data from this research with that done by MIT and WWC (Ng 1998) for the plot of AAOS vs. Log(ts/tr) (Figure 6.9), it can be seen that the average trend line through the origin is lower than that produced by MIT. Where Ng (1998) has the equation of Log(ts/tr) = 0.0206 (AAOS) and the equation for the average trend from this research is Log(ts/tr) = 0.0174 (AAOS). This trendline is slightly lower but compares well with previous work. 7.3 Recommendation for a Laboratory Testing Program When performing laboratory testing to determine the Mesri's Cα/CR ratio, the procedure is as follows: (1) perform a 1-D consolidation test on the sample; this is done to determine the consolidation properties such as σp, CR, and RR, (2) determine the rate of secondary compression for normally consolidated specimen, Cα by loading the soil sample to 1.5 to 2.0 times that of σp (to remove 91 any disturbance from the soil sample and to be sure that the soil is normally consolidated, (3) determine the rates of secondary compression for overconsolidated specimens, C'α. Refer to Appendix C for the detailed procedure used in this research. It is recommended that when determining the preconsolidation stress, each loading step should be moved to the next step with as little as possible secondary settlement occurring. If a large amount of secondary settlement occurs, the soil then becomes aged and it can have an impact on the results of the consolidation data. It is also advantageous to move to the next loading step with as little secondary compression occurring because the tests can be run in a shorter amount of time as opposed to the traditional 24-hour loading steps. 7.4 Additional Design Guidance The moisture content can be used for the estimating of Cc and CR of a saturated soil (Figures 6.10 and 6.11). For application purposes, it is recommended the Bartlett and Lee (2004) equation CR = 0.0053(ω) - 0.0283 be used because of its greater statistical support. Correlations for Cα, Cα'/CR, and Cα'/Cα with ω were explored in this research (Figures 6.12 to 6.14). Based on these plots, there is poor to very poor correlation between Cα and Cα'/Cα with ω (Figures 6.12 and 6.14, respectively). The correlation between Cα/CR and ω has some promise for future development and application (Figure 6.13). However, more observations are needed to improve the statistical support for this relation. 92 7.5 Recommendation for Further Testing It is recommended that when future testing is being performed, Atterburg limits and fines wash to be performed on every test specimen for a better classification of the soil being tested. This will also lead to a better understanding of how each type of soil behaves during long-term settlement. It is also recommended that in the future, a more advanced testing with pore water pressures mesurements be performed, so as to know when primary settlement is complete and to have very little secondary settlement occur. 93 APPENDIX A PLOTS FOR PRECONSOLIDATION STRESS 94 A1 400 South at 15-17 feet 1.34 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.10 1.00 10.00 Strain (ΔH/H) Stress (tsf) Log of Stress vs. Strain 1.39 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0 1 2 3 4 5 6 7 8 9 Work (ton/ft3) Stress (tsf) 95 A2 400 South at 15-17 feet 0.5 0.7 0.9 1.1 1.3 1.5 1.7 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf 0.5 0.7 0.9 1.1 1.3 1.5 1.7 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 96 A3 400 South at 15-17 feet 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 97 A4 400 South at 15-17 feet 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 98 A 5 400 South at 15-17 feet 0 1 2 3 4 5 6 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf 0 1 2 3 4 5 6 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 99 A6 400 South at 15-17 feet 0 1 2 3 4 5 6 7 8 9 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf 0 1 2 3 4 5 6 7 8 9 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 100 A7 400 South at 15-17 feet 0 1 2 3 4 5 6 7 8 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf 0 1 2 3 4 5 6 7 8 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 101 A8 400 South at 15-17 feet -0.38 -0.36 -0.34 -0.32 -0.30 -0.28 -0.26 -0.24 -0.22 -0.20 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf -0.40 -0.38 -0.36 -0.34 -0.32 -0.30 -0.28 -0.26 -0.24 -0.22 -0.20 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 102 A9 400 South at 15-17 feet -0.65 -0.60 -0.55 -0.50 -0.45 -0.40 -0.35 -0.30 -0.25 -0.20 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf -0.65 -0.60 -0.55 -0.50 -0.45 -0.40 -0.35 -0.30 -0.25 -0.20 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 103 A10 400 South at 15-17 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 104 A11 400 South at 15-17 feet -1.0 -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf -1.0 -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 105 A12 400 South at 15-17 feet -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 1000 10000 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 5 10 15 20 25 30 35 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 106 A13 400 South at 20-22 feet 3.17 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.10 1.00 10.00 100.00 Strain (ΔH/H) Stress (tsf) Log of Stress vs. Strain 3.21 0 0.2 0.4 0.6 0.8 1 1.2 1.4 0 2 4 6 8 10 12 14 16 18 Work (ton/ft3) Stress (tsf) 107 A14 400 South at 20-22 feet 0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf 0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 0 0.5 1 1.5 2 2.5 3 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 108 A15 400 South at 20-25 feet 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.1 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.1 0 0.5 1 1.5 2 2.5 3 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 109 A16 400 South at 20-25 feet 0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3 1.4 1.5 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf 0.5 0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3 1.4 1.5 0 0.5 1 1.5 2 2.5 3 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 110 A17 400 South at 20-22 feet 0.5 0.7 0.9 1.1 1.3 1.5 1.7 1.9 2.1 2.3 2.5 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf 0.5 0.7 0.9 1.1 1.3 1.5 1.7 1.9 2.1 2.3 2.5 0 0.5 1 1.5 2 2.5 3 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 111 A18 400 South at 20-22 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 0 0.5 1 1.5 2 2.5 3 3.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 112 A19 400 South at 20-22 feet 0 1 2 3 4 5 6 7 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf 0 1 2 3 4 5 6 7 0 0.5 1 1.5 2 2.5 3 3.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 113 A20 400 South at 20-22 feet 0 1 2 3 4 5 6 7 8 9 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 16.0 tsf 0 1 2 3 4 5 6 7 8 9 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 16.0 tsf 114 A21 400 South at 20-22 feet -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 115 A22 400 South at 20-22 feet -0.65 -0.6 -0.55 -0.5 -0.45 -0.4 -0.35 -0.3 -0.25 -0.2 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf -0.65 -0.6 -0.55 -0.5 -0.45 -0.4 -0.35 -0.3 -0.25 -0.2 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 116 A23 400 South at 20-22 feet -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 117 A24 400 South at 20-22 feet -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 118 A25 400 South at 20-22 feet -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 119 A 26 400 South at 20-22 feet -1.6 -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf -1.6 -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 120 A27 400 South at 25-27 feet 1.49 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.10 1.00 10.00 100.00 Strain (ΔH/H) Stress (tsf) Log of Stress vs. Strain 1.52 0 0.2 0.4 0.6 0.8 1 1.2 1.4 0 2 4 6 8 10 12 14 16 18 Work (ton/ft3) Stress (tsf) 121 A28 400 South at 25-27 feet 0.5 0.7 0.9 1.1 1.3 1.5 1.7 1.9 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf 0.5 0.7 0.9 1.1 1.3 1.5 1.7 1.9 0 0.5 1 1.5 2 2.5 3 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 122 A29 400 South at 25-27 feet 0.0 0.5 1.0 1.5 2.0 2.5 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf 0.0 0.5 1.0 1.5 2.0 2.5 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 123 A30 400 South at 25-27 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 124 A31 400 South at 25-27 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 125 A32 400 South at 25-27 feet 0 1 2 3 4 5 6 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf 0 1 2 3 4 5 6 7 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 126 A33 400 South at 25-27 feet 0 1 2 3 4 5 6 7 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf 0 1 2 3 4 5 6 7 8 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 127 A34 400 South at 25-27 feet 0 1 2 3 4 5 6 7 8 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 16.0 tsf 0 1 2 3 4 5 6 7 8 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 16.0 tsf 128 A35 400 South at 25-27 feet -0.36 -0.34 -0.32 -0.30 -0.28 -0.26 -0.24 -0.22 -0.20 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf -0.36 -0.34 -0.32 -0.30 -0.28 -0.26 -0.24 -0.22 -0.20 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 129 A36 400 South at 25-27 feet -0.65 -0.6 -0.55 -0.5 -0.45 -0.4 -0.35 -0.3 -0.25 -0.2 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf -0.65 -0.6 -0.55 -0.5 -0.45 -0.4 -0.35 -0.3 -0.25 -0.2 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 130 A37 400 South at 25-27 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 131 A38 400 South at 25-27 feet -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 132 A39 400 South at 25-27 feet -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 133 A40 400 South at 25-27 feet -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf -1.4 -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 134 A41 400 South at 30-32 feet 2.74 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.10 1.00 10.00 100.00 Strain (ΔH/H) Stress (tsf) Log of Stress vs. Strain 2.64 0 0.2 0.4 0.6 0.8 1 1.2 0 2 4 6 8 10 12 14 16 18 Work (ton/ft3) Stress (tsf) 135 A42 400 South at 30-32 feet 0.5 0.7 0.9 1.1 1.3 1.5 1.7 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf 0.5 0.7 0.9 1.1 1.3 1.5 1.7 0 0.5 1 1.5 2 2.5 3 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 136 A43 400 South at 30-32 feet 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 0 0.5 1 1.5 2 2.5 3 3.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 137 A44 400 South at 30-32 feet 0.0 0.5 1.0 1.5 2.0 2.5 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 0 0.5 1 1.5 2 2.5 3 3.5 4 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 138 A45 400 South at 30-32 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0 0.5 1 1.5 2 2.5 3 3.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 139 A46 400 South at 30-32 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 0 0.5 1 1.5 2 2.5 3 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 140 A47 400 South at 30-32 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 0 0.5 1 1.5 2 2.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 141 A48 400 South at 30-32 feet 2.0 2.5 3.0 3.5 4.0 4.5 5.0 5.5 6.0 6.5 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 16.0 tsf 2.0 2.5 3.0 3.5 4.0 4.5 5.0 5.5 6.0 6.5 0 0.5 1 1.5 2 2.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 16.0 tsf 142 A49 400 South at 30-32 feet -0.32 -0.31 -0.30 -0.29 -0.28 -0.27 -0.26 -0.25 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf -0.32 -0.31 -0.30 -0.29 -0.28 -0.27 -0.26 -0.25 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 143 A50 400 South at 30-32 feet -0.60 -0.55 -0.50 -0.45 -0.40 -0.35 -0.30 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf -0.60 -0.55 -0.50 -0.45 -0.40 -0.35 -0.30 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 144 A51 400 South at 30-32 feet -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 145 A52 400 South at 30-32 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 146 A53 400 South at 30-32 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 147 A54 400 South at 30-32 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 148 A55 400 South at 37.5-39.5 feet 2.71 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.10 1.00 10.00 100.00 Strain (ΔH/H) Stress (tsf) Log of Stress vs. Strain 2.76 0 0.2 0.4 0.6 0.8 1 1.2 0 2 4 6 8 10 12 14 16 18 Work (ton/ft3) Stress (tsf) 149 A56 400 South at 37.5-39.5 feet 0.5 0.7 0.9 1.1 1.3 1.5 1.7 1.9 2.1 2.3 2.5 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf 0.5 0.7 0.9 1.1 1.3 1.5 1.7 1.9 2.1 2.3 2.5 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 150 A57 400 South at 37.5-39.5 feet 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 151 A58 400 South at 37.5-39.5 feet 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 152 A59 400 South at 37.5-39.5 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 153 A60 400 South at 37.5-39.5 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 154 A61 400 South at 37.5-39.5 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0 0.5 1 1.5 2 2.5 3 3.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 155 A62 400 South at 37.5-39.5 feet 0 1 2 3 4 5 6 7 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 16.0 tsf 0 1 2 3 4 5 6 7 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 16.0 tsf 156 A63 400 South at 37.5-39.5 feet -0.45 -0.40 -0.35 -0.30 -0.25 -0.20 -0.15 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf -0.45 -0.40 -0.35 -0.30 -0.25 -0.20 -0.15 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 157 A64 400 South at 37.5-39.5 feet -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 158 A65 400 South at 37.5-39.5 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 159 A66 400 South at 37.5-39.5 feet -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 160 A67 400 South at 37.5-39.5 feet -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 161 A68 400 South at 37.5-39.5 feet -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 162 A69 400 South at 40-42 feet 0.81 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.10 1.00 10.00 Strain (ΔH/H) Stress (tsf) Log of Stress vs. Strain 0.83 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0 1 2 3 4 5 6 7 8 9 Work (ton/ft3) Stress (tsf) 163 A70 400 South at 40-42 feet 0.0 1.0 2.0 3.0 4.0 5.0 6.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf 0.0 1.0 2.0 3.0 4.0 5.0 6.0 0 2 4 6 8 10 12 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 164 A71 400 South at 40-42 feet 0.0 0.5 1.0 1.5 2.0 2.5 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf 0.0 0.5 1.0 1.5 2.0 2.5 0 2 4 6 8 10 12 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 165 A72 400 South at 40-42 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 0.1 1 10 100 1000 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 0 2 4 6 8 10 12 14 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 166 A73 400 South at 40-42 feet 0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf 0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 167 A 74 400 South at 40-42 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0 1 2 3 4 5 6 7 8 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 168 A 75 400 South at 40-42 feet 0.0 1.0 2.0 3.0 4.0 5.0 6.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf 0.0 1.0 2.0 3.0 4.0 5.0 6.0 0 1 2 3 4 5 6 7 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 169 A 76 400 South at 40-42 feet -0.18 -0.16 -0.14 -0.12 -0.10 -0.08 -0.06 -0.04 -0.02 0.00 0.1 1 10 100 1000 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf -0.18 -0.16 -0.14 -0.12 -0.10 -0.08 -0.06 -0.04 -0.02 0.00 0 5 10 15 20 25 30 35 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 170 A 77 400 South at 40-42 feet -0.40 -0.35 -0.30 -0.25 -0.20 -0.15 -0.10 -0.05 0.00 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf -0.40 -0.35 -0.30 -0.25 -0.20 -0.15 -0.10 -0.05 0.00 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 171 A 78 400 South at 40-42 feet -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 172 A 79 400 South at 40-42 feet -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.5 tsf -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.5 tsf 173 A 80 400 South at 40-42 feet -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 174 A81 400 South at 45-47 feet 2.13 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.10 1.00 10.00 100.00 Strain (ΔH/H) Stress (tsf) Log of Stress vs. Strain 2.12 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 0 2 4 6 8 10 12 14 16 18 Work (ton/ft3) Stress (tsf) 175 A 82 400 South at 45-47 feet 1.0 1.2 1.4 1.6 1.8 2.0 2.2 2.4 2.6 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf 1.0 1.2 1.4 1.6 1.8 2.0 2.2 2.4 2.6 0 0.5 1 1.5 2 2.5 3 3.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 176 A 83 400 South at 45-47 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0 1 2 3 4 5 6 7 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 177 A 84 400 South at 45-47 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 178 A 85 400 South at 45-47 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 179 A 86 400 South at 45-47 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 180 A 87 400 South at 45-47 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 181 A 88 400 South at 45-47 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 16.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 0 0.5 1 1.5 2 2.5 3 Strain (%) Root of Time (√min) Root of Time vs. Strain for 16.0 tsf 182 A 89 400 South at 45-47 feet -0.16 -0.14 -0.12 -0.10 -0.08 -0.06 -0.04 -0.02 0.00 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf -0.16 -0.14 -0.12 -0.10 -0.08 -0.06 -0.04 -0.02 0.00 0 1 2 3 4 5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 183 A 90 400 South at 45-47 feet -0.6 -0.6 -0.5 -0.5 -0.4 -0.4 -0.3 -0.3 -0.2 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf -0.6 -0.6 -0.5 -0.5 -0.4 -0.4 -0.3 -0.3 -0.2 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 184 A 91 400 South at 45-47 feet -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 185 A 92 400 South at 45-47 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 186 A 93 400 South at 45-47 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 187 A 94 400 South at 45-47 feet -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf -1.2 -1.0 -0.8 -0.6 -0.4 -0.2 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 188 A95 400 South at 50-52 feet 1.18 0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.10 1.00 10.00 Strain (ΔH/H) Stress (tsf) Log of Stress vs. Strain 1.06 0 0.1 0.2 0.3 0.4 0.5 0.6 0 1 2 3 4 5 6 7 8 9 Work (ton/ft3) Stress (tsf) 189 A96 400 South at 50-52 feet 1.4 1.6 1.8 2.0 2.2 2.4 2.6 0.1 1 10 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.25 tsf 1.4 1.6 1.8 2.0 2.2 2.4 2.6 0 0.5 1 1.5 2 2.5 3 3.5 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.25 tsf 190 A97 400 South at 50-52 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 191 A98 400 South at 50-52 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0 2 4 6 8 10 12 14 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 192 A99 400 South at 50-52 feet 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 0 2 4 6 8 10 12 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 193 A100 400 South at 50-52 feet 0 1 2 3 4 5 6 7 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf 0 1 2 3 4 5 6 7 0 2 4 6 8 10 12 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 194 A101 400 South at 50-52 feet 0 1 2 3 4 5 6 7 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 8.0 tsf 0 1 2 3 4 5 6 7 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 8.0 tsf 195 A102 400 South at 50-52 feet -0.35 -0.30 -0.25 -0.20 -0.15 -0.10 -0.05 0.00 0.1 1 10 100 1000 10000 Strain (%) Log of Time (min) Log of Time vs. Strain for 4.0 tsf -0.35 -0.30 -0.25 -0.20 -0.15 -0.10 -0.05 0.00 0 10 20 30 40 50 60 70 Strain (%) Root of Time (√min) Root of Time vs. Strain for 4.0 tsf 196 A103 400 South at 50-52 feet -0.45 -0.40 -0.35 -0.30 -0.25 -0.20 -0.15 -0.10 -0.05 0.00 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 2.0 tsf -0.45 -0.40 -0.35 -0.30 -0.25 -0.20 -0.15 -0.10 -0.05 0.00 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 2.0 tsf 197 A104 400 South at 50-52 feet -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 1.0 tsf -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 1 2 3 4 5 6 Strain (%) Root of Time (√min) Root of Time vs. Strain for 1.0 tsf 198 A105 400 South at 50-52 feet -1.0 -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Time vs. Strain for 0.50 tsf -1.0 -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0 2 4 6 8 10 Strain (%) Root of Time (√min) Root of Time vs. Strain for 0.50 tsf 199 A106 400 South at 50-52 feet -0.9 -0.8 -0.7 -0.6 -0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 1 10 100 Strain (%) Log of Time (min) Log of Ti