UNIVERSITY OF CINCINNATI

Transcription

1 UNIVERSITY OF CINCINNATI Date:08/04/06 I, Nishant Dayal, hereby submit this work as part of the requirements for the degree of: M.S. in: Civil Engineering It is entitled: Consolidation Analyses of Greater Cincinnati Soils This work and its defense approved by: Chair: Dr. Mark T. Bowers Dr. Anastasios M. Ioannides George C. Webb

2 Consolidation Analyses of Greater Cincinnati Soils Cincinnati, Ohio A thesis submitted to the Division of Research and Advanced Studies of the University of Cincinnati in partial fulfillment of the requirements for the degree of MASTER OF SCIENCE in the Department of Civil and Environmental Engineering of the College of Engineering 2006 by Nishant Dayal B.S., Regional Institute of Technology, Jamshedpur, India, 2000 (now, National Institute of Technology) Committee Chair: Dr. Mark T. Bowers

3 Abstract The compression analysis of argillaceous material under a proposed superstructure is one of the primary tasks required of a geotechnical engineer. The accuracy of such an analysis is dubious in the absence of proper geotechnical investigation. A proper geotechnical investigation inlves site reconnaissance, determination of subsurface conditions, in situ and laboratory tests, and settlement monitoring, if required. This research fulfills several objectives, i.e. analyses of primary consolidation, proposes new correlations, and compares observed and predicted values of settlement. The current paper introduces several new correlations between the compression index and soil parameters such as natural water content, initial id ratio, liquid limit, plasticity index, specific gravity of soil solids, and dry unit weight for the soils of Greater Cincinnati based on regression analysis. Previously available correlations are also analyzed. Most of the new equations obtained show statistical improvement over their predecessors. The data used to obtain these new correlations were provided by local consulting firms including H. C. Nutting Company and Thelen Associates, as well as the Ohio Department of Transportation. Several case studies are presented which compare predicted settlements (by local engineers) and measured settlements due to consolidation of clays. This work then validates the new equations. These case studies are based on project reports provided by H. C. Nutting Company and are intended to help local engineers in predicting consolidation settlements more accurately. iii

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5 ACKNOWLEDGEMENTS I am indebted to many people who have helped directly and indirectly in making this paper a success. First, I would like to thank my thesis advisor Dr. Mark T. Bowers for his continuous support, guidance, and encouragement through my graduate years. I would also like to extend my gratitude to other members of my committee for their time and commitment: Dr. Anastasios M. Ioannides and Mr. George C. Webb. I indeed appreciate Mr. Webb s help for opening the doors of the H.C. Nutting Company for this research. Local geotechnical engineering firms were instrumental in this research by providing existing data and project reports for the Greater Cincinnati area. I would like to thank Mr. Ted Vogelpohl of Thelen Associates; Mr. Joe Smithson of ODOT; and Mr. Robert Lennertz and Dr. Andrew Bodocsi of the H.C. Nutting Company. All my friends were very helpful, especially Ashish Kumar for being there for my parents when I wasn t able to and James Swindler for his encouragement throughout my graduate years. I would finally like to thank my family for their love and support, especially my cousin Mukul Anand and my parents, Sudhir and Usha Dayal. iv

6 DISCLAIMER Reference herein to any specific commercial products, process, or service by trade name, trademark, manufacturer, or otherwise, does not necessarily constitute or imply its endorsement, recommendation, or faring by the author of this paper. The views and opinions of the author expressed herein shall not be used for advertising or product endorsement purposes. v

7 Table of Contents 1. Introduction Settlement in Clays One-Dimensional Consolidation Important Factors that Influences Primary Consolidation Literature Review Introduction Previous Correlations Theory of Consolidation Laboratory Test Statistics General Definitions Methods and Data Introduction Characteristics of Data Compression Index using a single soil parameter Compression Index using multiple soil parameters Case Studies Introduction Case A- The Expansion of Warehouse and Distribution Facility in Claryville, KY Project Summary Geology and Subsurface Conditions Conclusions and Recommendations by the Project Engineer vi

8 Analysis Case B- Large One-Story Structure in Newport, KY Project Summary Geology and Subsurface Conditions Conclusions and Recommendations by the Project Engineer Analysis Case C- Five Story Steel Frame Structure, Milford, OH Project Summary Geology and Subsurface Conditions Conclusions and Recommendations by the Project Engineer Analysis Case D- Two Story Steel Frame Structure. Springdale, OH Project Summary Geology and Subsurface Conditions Conclusions and Recommendations by the Project Engineer Analysis Case E- Aeration Basins. North Bend, OH Project Summary Geology and Subsurface Conditions Conclusions and Recommendations by the Project Engineer Analysis Case F- Large One-Story Structure. Cincinnati, OH Project Summary Geology and Subsurface Conditions vii

9 Conclusions and Recommendations by the Project Engineer Analysis Conclusions and Recommendations References viii

10 List of Figures Figure 2.1. Void ratio versus log pressure curve illustrating deposition, sampling (unloading) and reconsolidation in the consolidation test apparatus.. 6 Figure 2.2. Direct determination methods of preconsolidation pressure Figure 2.2. Direct determination methods of preconsolidation pressure (contd.)...11 Figure 2.3. Data on sedimentation compression curves - with two laboratory compression curves for comparison (Skempton, 1944).13 Figure 3.1. Classification of soil samples.. 27 Figure 3.2. Relationship between compression index and initial id ratio.. 29 Figure 3.3. Relationship between compression index and liquid limit.. 30 Figure 3.4. Relationship between compression index and natural water content.. 31 Figure 3.5. Relationship between compression index and dry unit weight 32 Figure 3.6. Relationship between compression index and plasticity index Figure 4.1. Location of warehouse and distribution facility in Claryville, KY...39 Figure 4.2. Subsurface profile based on the borehole logs.41 Figure 4.3. Test boring location and settlement plates plan...45 Figure 4.4. Test boring location plan Figure 4.5. Assumed soil profile for settlement calculation Figure 4.6. Location of large one-story structure in Newport, KY Figure 4.7. Boring locations Figure 4.8. Subsurface investigation.. 60 Figure 4.9. Legend for figure 4.8 (re-drawn)..61 ix

11 Figure Assumed soil profile for settlement calculation for the location of plate no Figure Generalized representation of settlement plate readings (SP 1 to SP 8) 67 Figure Generalized representation of settlement plate locations (SP 1 to SP 8).68 Figure Location of the five story steel frame structure, Milford, OH 69 Figure Assumed soil profile for settlement calculation for the location of plate # 1 near Test boring hole Figure Test boring number Figure Settlement recorded at plate # Figure Settlement recorded at plate # Figure Settlement plate locations Figure Summary of geotechnical data Figure Summary of geotechnical data (contd.) Figure Legend for Figures 4.19 and Figure Location of two story steel frame structure, Springdale, OH.. 81 Figure 4.23 Test boring location plan.. 83 Figure Subsurface profile for test borings A, B, E, and F.. 84 Figure Subsurface profile for test borings D, G, and C.. 85 Figure Assumed profile for settlement estimation.. 89 Figure Settlements plate locations Figure Recorded settlement for plate Figure Recorded settlement for plates 1 to Figure Location of aeration basins, North Bend, OH. 95 Figure Simplified subsurface profile 1 under Aeration basin.. 98 x

12 Figure Simplified subsurface profile 1 under Aeration basin. 100 Figure Test boring and settlement plate location plan Figure Settlement plate readings for aeration basin # Figure Generalized representation of settlement plate readings for aeration basin # Figure Location of large one-story structure, Cincinnati, OH 105 Figure Simplified subsurface profile for B Figure Simplified subsurface profile for B Figure Settlement plates location plan xi

13 List of Tables Table 2.1. Mechanisms causing a maximum past pressure 9 Table 2.2. Optimized regression equations inlving a single soil parameter for the compression index Table 2.3. Optimized regression equations inlving multiple soil parameters for the compression index Table 3.1. Summary of statistical parameters for different soil properties of all samples Table 3.2. Summary of statistical parameters for different soil properties of all samples for the current research Table 3.3. Ratio of predicted to measured compression indices using single and multiple soil parameters Table 3.4. Ratio of predicted to measured compression indices using single and multiple soil parameters for Cincinnati and Northern Kentucky Table 4.1. Predicted floor slab settlement for case A- warehouse, Claryville, KY Table 4.2. SP-1 Settlement plate readings 46 Table 4.3. SP-2 Settlement plate readings 47 Table 4.4. SP-3 Settlement plate readings 48 Table 4.5. SP-4 Settlement plate readings 49 Table 4.6. SP-5 Settlement plate readings 50 Table 4.7. Limiting angular distortion as recommended by Bjerrum Table 4.8. Allowable settlement criteria: 1955 U.S.S.R. Building Code.. 87 Table 4.9. Allowable average settlement for different building types.. 88 Table Recorded depth of fill xii

14 Table Recorded settlements Table Case study summary xiii

15 List of Symbols σ, p = Preconsolidation Pressure p c σ, = Effective Overburden Pressure vc σ e = Void Ratio e = Initial Void Ratio 0 LL = Liquid Limit PI = Plasticity Index C = Compression Index c C = Recompression Index r w = Natural Water Content n w = Initial Water Content 0 G = Specific Gravity of Soil Solids s γ = Dry Unit Weight d γ = Wet Unit Weight m σ = Change in Effective Stress n = Initial Porosity 0 S = Undrained Shear Strength u µ = Mean x σ = Variance 2 x σ = Standard Deviation x c = Coefficient of Consolidation v 2 r = Coefficient of Determination - xiv -

16 1. INTRODUCTION This research is directed to fulfill several objectives. One objective is to obtain new correlations for compression index and soil parameters inlving soils in Greater Cincinnati area. Another important objective is to dissect the thought process of local engineers as they deal with consolidation problems. This will be investigated through several case studies based on actual settlement reports. 1.1 Settlement in Clays Any application of load either by a structure or fill leads to deformations of a soil deposit. The magnitude of this deformation is known as settlement. Engineers are interested in obtaining an estimate of settlement and the time for such settlement or a given percentage of it to occur. Total settlement can be divided into three basic components: 1. Initial Settlement (Immediate or Elastic Settlement) The application of load causes some settlement immediately due to distortion of soil structure in saturated and unsaturated clays. Immediate settlement also occurs in unsaturated soils due to lume change on the application of load. 2. Primary Consolidation Primary consolidation inlves the gradual dissipation of excess pore water pressure due the application of external load and subsequent drainage. The rate of lume change in the clays is a function of the drainage under a hydraulic gradient. 3. Secondary Compression The end of the primary consolidation phase usually marks the beginning of a secondary compression phase without any increase in the effective stress. The deformation continues in the soil skeleton even after the primary phase is over - 1 -

17 albeit at a much slower rate. This is because soil structure realigns to more stable arrangements due to creep under the influence of a constant stress. The rate by which the structure can deform controls the rate of secondary compression (Mitchell, 1993). It is difficult to separate the primary and secondary processes in the field if the layer undergoing the primary phase is relatively thick (Holtz and Kovacs, 1981). 1.2 One-Dimensional Consolidation Settlement analysis of saturated clays in a one-dimensional environment (Terzaghi, 1925) can be accomplished by carrying out drained consolidation tests. The variability between field and oedometer test results is minimal when the following conditions prevail: 1. A relatively thin layer of clay lying between incompressible layers; 2. A large loaded area of which the horizontal extent is great compared with the thickness. 1.3 Important Factors that Influences Primary Consolidation Every natural soil is characterized by its structure and composition. The structure, in turn, is dependent on time. A change in the effective stress experienced by the soil changes the structure in time. Any further change in the structure is dependent on the composition of the soil and the increase in effective stresses. Some soils far more structural change than others based on their composition. Therefore, the factors that influence consolidation significantly can be summarized as: 1. The magnitude of applied load

18 2. The geometry of the loaded area. 3. Type of soil. 4. The stress history of the soil. 5. The rate of load application. Another important parameter that governs the amount and rate of consolidation in soils is temperature. A change in temperature leads to lume and/or effective stress changes in saturated soils (Mitchell, 1993). Under drained conditions, an increase in temperature causes thermal expansion of mineral solids and pore water. The change in lume due to temperature changes should be equal to the change in lume due to pressure changes under undrained conditions. Soils in the Greater Cincinnati typically consist of glacial till, outwash, loess, and silty and clayey lacustrine material underlain by interbedded shale and limestone bedrock and was formed mostly during Wisconsinan, Illinoan, or Kansan age (Soil Survey-Hamilton County, 1981). Data from more than a hundred consolidation tests were obtained from various local firms in Cincinnati/Northern Kentucky. These data were screened and analyzed to obtain new correlations. New correlations were required as most of the existing correlations are region specific and even the most widely used equations have a high variability. Most of these consolidation tests were carried out on samples obtained from various places in the Greater Cincinnati and Northern Kentucky region. A few settlement analysis reports were - 3 -

19 also obtained from the H. C. Nutting Company. These reports were used in the comparison of predicted and observed magnitudes of settlement and its time rate. The new correlations obtained from the collected consolidation test data were used by the author to predict settlement values using the settlement analysis reports. Thus, a comparison of the predicted values based on previous correlations and new correlations was made possible

20 2. Literature Review 2.1 Introduction Terzaghi s experiments on clay paved the way for the birth of modern soil mechanics (Terzaghi, 1922). Some of these experiments marked the beginning of Terzaghi s consolidation theory. This theory explained the load transfer from pore water in the clay strata to the mineral skeleton of the soil. Terzaghi presented the basic differential equation for the consolidation process in a paper published in It was possible to compute the rate of settlement with the use of this equation by conducting laboratory tests. Terzaghi concluded that the relationship between id ratio and pressure for the virgin section of the compression curve could be illustrated by a logarithmic curve (Casagrande, 1936, Figure 2.1). Preconsolidation pressure (Table 2.1) is one of the most important components of clay used in the accurate evaluation of settlement due to primary consolidation. Preconsolidation pressure is defined as the maximum pressure experienced by a soil profile. A soil is said to be overconsolidated when its preconsolidation pressure is greater than the present effective overburden stress and it is said to be normally consolidated when its maximum past pressure is at equilibrium with the present effective overburden pressure. Over Consolidation Ratio (OCR) is defined as the ratio of present effective overburden pressure to preconsolidation pressure

21 Figure 2.1: Void ratio versus log pressure curve illustrating deposition, sampling (unloading) and reconsolidation in the consolidation test apparatus. (Holtz and Kovacs, 1981) - 6 -

22 One of the greatest difficulties in obtaining an accurate magnitude of consolidation is the ability to simulate field conditions in the laboratory. Much research has been carried out and scientists have developed nine empirical methods to estimate the preconsolidation pressure. The most popular of these are Casagrande s graphical method and Schmertmann s method (1955, Figure2.2d). According to Senol and Saglamer (2000): The Tavenas method (1979, Figure 2.2b) gives better results than Casagrande method (1936, Figure 2.2c). The value of the correlation coefficient (84-90%) is an acceptable result. The correlation coefficient (77-82%) of the Casagrande method is also good (Şenol, 1997). Other methods suggested for the determination of maximum past pressure include Burmister (1951, Figure 2.2f), Janbu (1967, Figure 2.2a), Butterfield (1979, Figure 2.2e), Senol (1996, Figure 2.2g), and more recently, the dissipated strain energy method (Wang and Frost, 2004). According to Butterfield (1979), Sridharan (1991), and Den Haan (1992), log( 1+ e) log p correlation is better than e log p correlation for several reasons, such as a better log( 1+ e) log p linear relation, the possibility of yielding a negative id ratio by the e log p method, and the physical meaning of the reference lume of ( 1+ e) as the total lume. However, the second reason is questionable as it has never been practically observed to yield a negative id ratio. The greatest hurdle in the determination of maximum past pressure is sample disturbance. While soil disturbance is still not understood properly, it lowers the magnitude of preconsolidation pressure and the lume of ids for any given value of effective overburden pressure. This results in an increase in compressibility with stresses less than - 7 -

23 the preconsolidation pressure, and a decrease in compressibility with stresses greater than the maximum past pressure. This is illustrated in the change in shape of the consolidation curve (Figure 2.1). Terzaghi (1941) concluded that every clay passes from a solid to a partially lubricated state during sampling operations. This leads to a loss of information regarding the physical properties of clays in the solid state. Hence, the only way this information can be obtained is through proper field observations. These observations include but are not limited to 1. Subsurface investigation through drilled holes. 2. Geological considerations. Sample disturbance obscures the stress history of a soil profile which, in turn, makes an accurate determination of preconsolidation pressure impossible (Mitchell, 1993). It has been observed that the slope of the recompression curve is less than that of the field virgin compression curve even with high quality sampling and testing (Holtz and Kovacs, 1981). The evaluation of the slope of the field virgin compression curve (known as compression index, C c ) was made possible by Schmertmann (1955) with his graphical method (Figure 2.2d). The procedure inlves a typical e log p plot. The magnitude of compression index for any soil is very important in the determination of settlement under the application of load. Since this is purely empirical, scientists have previously developed correlations between compression index and other soil parameters based on the available data and other assumptions

24 Table 2-1 Mechanisms causing a maximum past pressure (Ladd, 1971) Item Remarks/References A. Change in total stress due to: 1. Removal of overburden 2. Past structures 3. Glaciation B. Change in pore water pressure due to: 1. Change in water table elevation 2. Artesian pressures 3. Deep pumping 4. Desiccation due to drying 5. Desiccation due to plant life 1. See Kenney (1964) for sea-level changes 2. Common is glaciated areas 3. Common in many cities 4. May have occurred during deposition 5. May have occurred during deposition C. Changes in soil structure due to: 1. Secondary compression (aging) 2. Changes in environment, such as ph, temperature, salt concentration, etc. 3. Chemical alternations due to: weathering, precipitation of cementing agents, ion exchange, etc. 4. Change of strain rate on loading 1. Raju (1956), Leonards and Ramiah (1959), Leonards and Altschaeffl (1964)* 2. Lambe (1958) Bjerrum (1967) Leonards and Altschaeffl (1964) 3. Bjerrum (1967), Cox (1968) 4. Lowe (1974) *The magnitude of OCR due to secondary compression for mature natural deposits of highly plastic clays may reach values as high as 1.6 ±

25 (a) Janbu Methods (b) Tavenas Method (c) Casagrande Method Figure 2.2: Direct Determination Methods of Preconsolidation Pressure (Senol and Saglamer, 2000)

26 (d) Schmertmann Method (e) Butterfield Method (f) Burmister Method (g) New Method (Senol, 1997) Figure 2.2: Direct Determination Methods of Preconsolidation Pressure (contd., Senol and Saglamer, 2000)

27 2.2 Previous Correlations Consolidation tests conducted by Skempton (1944) on remolded clays with their initial moisture content equal to their liquid limit resulted in the following equation: C c = (LL - 10) (2.1) Where: C c = Compression index = Slope of compression curve in virgin region of the consolidation curve for remolded cohesive soil. LL = Liquid limit (%) Skempton (1944) concluded: 1. The type of clay, as obtained from its liquid limit or hydrometer analysis, is an important factor which governs both the laboratory and sedimentation compression curves. 2. Generally, a compression curve obtained by laboratory tests for a particular type of clay is similar to its sedimentation compression curve but falls below the sedimentation curve. 3. Sedimentation curves and compression curves obtained by laboratory tests on undisturbed samples lie closer to each other (Figure 2.3). Nishida (1956) correlated the compression index with initial id ratio by introducing the fundamental relation given below: C c = 1.15 (e 0.35).(2.2a) The above correlation was obtained after assuming the soil grains to be uniform rigid spheres in a state of closest packing

28 Sedimentation curve Laboratory curve Suggested typical sedimentation curve for clays with LL = 60 to 80 Average liquid limits are given in brackets Figure 2.3: Data on sedimentation compression curves - with two laboratory compression curves for comparison (Skempton, 1944)

29 Nishida (1956) obtained a second correlation due to a loosely packed soil grain structure: C c = 1.15 (e 0.91)....(2.2b) A third correlation assuming deformable but incompressible soil grains was also obtained: C c = 1.15 e...(2.2c) Where: e = in situ id ratio under the present effective overburden stress. Nishida expressed the relation between compressive pressure and the id ratio as e = e o - C c (log p log p o )...(2.2d) where p o is the pressure corresponding to the initial id ratio e o and p is the pressure corresponding to the id ratio e. A fourth correlation between initial id ratio and compression index was obtained by substituting the value of C c from equation 2.2d in equation 2.2a: C c = 0.54 (e o 0.35)....(2.2e) If it is assumed that the ids of a soil mass are saturated with water and that the specific gravity of the soil solids is about 2.60, equation 2.2e becomes: C c = 0.54 (2.6w o 0.35)...(2.2f) Where: w o = initial water content. Nishida (1956) concluded that equation 2.2e agrees well with test results for any kind of soil

30 Hough (1957) introduced another correlation inlving compression index (C c ) and initial id ratio (e o ): C c = 0.30 (e o 0.27).(2.3a) Hough s equation is applicable primarily for inorganic cohesive soil, silt, silty clay, and clay. He observed a converging pattern among the virgin compression curves of many different types of soil specimens. Hough s observation confirmed Schmertmann s findings (1955) which showed the id ratio at the point of convergence to be about 0.4e o. Hough (1957) conducted tests on remolded specimens of different types of sand and found the following linear relationship: C c = a (e o b)...(2.3b) Where a represents slope and is dependent mainly on particle shape, size, and gradation. b represents a close approximation of minimum attainable id ratio under normal circumstances. Terzaghi and Peck (1967) gave a correlation (based on Skempton s work) between compression index and liquid limit for undisturbed normally consolidated clays. The resulting equation was: C c = (LL 10).....(2.4) Where: C c = Compression index for undisturbed normally consolidated clay. LL = Liquid limit (%)

31 Nacci et al. (1975) proposed a correlation between compression index and plasticity index for North Atlantic clay: C c = (PI).(2.5) Where: PI = Plasticity index (%) Azzouz et al. (1976) carried out extensive regression analyses (on the results of approximately 700 consolidation tests) to introduce relations between compression index, natural moisture content, initial id ratio, and liquid limit for Chicago clay, Brazilian clay, Motley clays for Sao Paulo City, and organic soil: Chicago clay: C c = 0.01 w n...(2.6a) Brazilian clay: C c = (LL 9) (2.6b) Motley clays from Sao Paulo city: C c = (e o 1.87).(2.6c) Where: e o = initial id ratio; w n = natural moisture content (%) Chicago clay: C c = 0.208e o (2.6d) Where: e o = initial id ratio Organic soil, peat: C c = w n...(2.6e) Where: w n = natural moisture content (%) All clays: C c = (LL 9)..(2.6f) Where:

32 LL = Liquid Limit (%) Azzouz et al. (1976) concluded, that both compression index and compression ratio are best expressed in terms of the initial id ratio by means of simple linear regression models, and the introduction of other independent variables did not significantly improve the accuracy of the resulting regression models. Mayne (1980) obtained a correlation between compression index and liquid limit using clay samples from all over the world (56 data points with a correlation coefficient of 0.813): C c = (LL 13)/109 (2.7) Rendon-Herrero (1983) obtained a relation between the compression index, specific gravity of soil solids, and initial id ratio based on data from 94 consolidation test results. Approximately 85% of these were carried out on specimens with compression indices between 0.04 and 0.5: C c = 0.141G + e Gs s (2.8) Where: G s = specific gravity of soil solids e o = in situ id ratio Another correlation inlving compression index, specific gravity, and liquid limit was obtained by Nagaraj and Murty (1985):

33 C c = LL Gs.(2.9) 100 Where: G s = specific gravity of soil solids LL = liquid limit (%) Park and Koumoto (2004) added to the list of correlations by introducing compression index equations of remolded clays and undisturbed Ariake (Japan) clay: Remolded clay: C c /n o = C c Where: n o = initial porosity for correlation coefficient, r = and number of samples, Ns = 66 Undisturbed clay: C c /n o = C c Where: n o = initial porosity for correlation coefficient, r = and number of samples, Ns = 83 Yoon et al. (Table 2.2 and 2.3, 2004) proposed a set of equations correlating various index properties for marine clay from the east, south, and west coasts of Korea with the help of linear and multiple regression analysis based on data from 1200 consolidation tests

34 Table 2.2: Optimized regression equations inlving a single soil parameter for the compression index (Yoon et al., 2004) Equation For C c = f (LL) Correlation Coefficient, R Applicability C c = 0.012( LL ) 0.64 South coast C c = 0.011( LL 6.36) 0.64 East coast C c = 0.01( LL 10.9) 0.67 West coast For C = f w ) c ( n C 0.013( w 3.85) 0.73 South coast c = n C 0.01( w ) 0.54 East coast c = n C 0.011( w 11.22) 0.67 West coast c = n C c = f ( e 0 For ) C 0.54( e 0.37) 0.77 South coast c = o C 0.39( e 0.13) 0.54 East coast c = o C 0.37( e 0.28) 0.65 West coast c = o For C c = f (PI) C c = PI 0.61 East coast For C = f γ ) c ( d C 1.6γ South coast c = d C 0.66γ West coast c = d Note: Plasticity Index (PI), liquid limit (LL), and natural water content ( w n ) are in percentages, and dry unit weight ( γ ) is in t/m 3 (metric) d Table 2.3: Optimized regression equations inlving multiple soil parameters for the compression index (Yoon et al., 2004) Equation Correlation Applicability coefficient, R Cc = wn eo LL South coast Cc = LL eo PI East coast Cc = wn eo LL West coast

35 2.3 Theory of Consolidation Terzaghi (1925) proposed the first theory to evaluate the rate of consolidation for saturated clay soils. Terzaghi s treatment, however, was restricted to the one-dimensional problem of a column under a constant load. Biot (1941) presented a more complete treatment of the theory resulting in more general results. Biot (1941) made some basic assumptions: 1. The soil medium is isotropic. 2. The soil medium is elastic. 3. The stress-strain relation is linear. 4. The strains are small. 5. The water in the pores is incompressible but the water may contain compressible air bubbles. 6. Darcy s law is valid. Although assumptions (2) and (3) have been most criticized, they were also the assumptions made by Terzaghi (1925) and have been quite satisfactory for practical purposes. Accurate predictions of settlement and their time rates have been rare and the disparity between the predicted and actual values exists due to the following reasons (Crooks et al., 1984; Becker et al., 1984; Tse, 1985; Mitchell, 1986); 1. Variation in the magnitude of the predicted and observed initial pore pressure developed upon the application of load; 2. Variation in the predicted rate of consolidation based on laboratory tests and observed values in the field;

36 3. Variable rates of pore pressure dissipation during and after construction; 4. The pore pressure build up even after completion of loading. 5. Apparent lack of strength gain with consolidation following load application. 2.5 Laboratory Test The standard one-dimensional consolidation test procedure (ASTM D2435), also known as the oedometer test, inlves strains and drainage only in the vertical direction with the load increment ratio being unity. This method is most widely used in the industry. The load-deformation data is either presented as vertical strain versus the effective consolidation stress or id ratio to the effective consolidation stress. Other than the standard one-dimensional consolidation test procedure, at least two more test procedures have been developed. These other one-dimensional consolidation tests are less time consuming with reasonably good results. The constant rate of strain method (Smith and Wahls, 1969) requires the operator to choose a suitable strain rate. Smith and Wahls (1969), based on the results of several tests conducted on two clays Massena clay and calcium montmorillonite, concluded that the e log p curves obtained at higher strain rates may deviate considerably from e log p curves obtained by standard one- dimensional consolidation tests. The constant gradient consolidation test was developed by Lowe et al. (1969). In such tests, the load is constantly changing with pore pressure. These tests have several advantages over the standard one (Ladd, 1971): 1. The time required to perform the complete test is significantly reduced. 2. A continuous compression curve can be obtained. 3. Continuous values of c v can be obtained

37 2.6 Statistics General Definitions Population Population is the representation of all likely observations of a random variable. Sample A sample represents the population. It is used for statistical analysis. Mean This is also known as expected value. It is a measure of the central tendency in the data and is also known as the first moment. Mean is generally represented by E(X) or µ x, and can be calculated for n observations as Mean = E(X) = µ x = 1 n x i n i= 1 Variance Variance is the spread of data about its mean or expected value and is also 2 known as second moment. It is generally denoted by σ X and is evaluated for n observations as Variance = Var(X) = σ = ( µ ) 2 X 1 1 n n i= 1 x i x Standard Deviaton It is the square root of variance and is denoted byσ. Standard deviation is also measured about its mean. Regression Analysis Regression analysis is a statistical tool that provides us with the ability to investigate relationships between variables. Simple regression inlves a dependent and an independent variable while multiple regression has several dependent variables. 2 X Algebraically, y = mx + c represents a straight line with x as the independent and y as the dependent variables. m is the slope and c is the y intercept. Multiple regression can be described as y = m1 x1 + m2 x mn xn + c

38 Microsoft Excel calculates the squared difference between the y-value estimated for each point and its actual y-value in regression analysis. The sum of these squared differences is called the residual sum of squares, ss resid. Microsoft Excel then evaluates the total sum of squares, ss total. When c is calculated normally, the total sum of squares is the sum of the squared differences between the actual y-values and the average of the y-values. When c is set equal to 0 and the m-values are adjusted to fit y = mx, the total sum of squares is the sum of the squares of the actual y-values (without subtracting the average y-value from each individual y-value). Then regression sum of squares, ss reg, can be found from: ss reg = ss total - ss resid The smaller the residual sum of squares is, compared with the total sum of squares, the larger the value of the coefficient of determination, r 2. r 2 = ss reg /ss total. The coefficient of determination compares estimated and actual y-values, and ranges in value from 0 to 1. If it is 1, there is a perfect correlation in the sample there is no difference between the estimated y-value and the actual y-value. On the other hand, if the coefficient of determination is 0, the regression equation is not helpful in predicting a y-value ( It is always possible to increase the value of r 2 by adding more regressor variables or increasing their range, therefore, a higher r 2 value does not always mean a better correlation (Haldar and Mahadevan, 1999)

39 3. Methods and Data 3.1 Introduction Cincinnati and its vicinity are underlain by an arrangement of horizontal beds of shale and limestone, the former exceeding the latter by a significant degree in thickness (Fenneman, 1916). These bedrock formations are primarily Ordovician in age ( million years ago). Colluvium, usually overlaying the bedrock, is characterized by high plasticity and low shear strength, and is formed by the weathering of shale. Glacial till, outwash sand and gravel, alluvium, lacustrine clays, and fill are among the other soils found in Cincinnati (Johnson, 2002). Glacial till consists of clay, silt, sand, gravel, and boulders and was deposited by melting glaciers. Outwash sand and gravel are the result of glacier retreat. Lacustrine sediments usually consist of fine sand, silt, and clay deposited on lake floors. These soils have low shear strengths. Alluvial sediments are carried by streams and rivers and deposited as these rivers slow down. The total settlement S for normally consolidated clay with a thickness H, initial id ratio ' e o, compression index C c, effective overburden pressure σ, and induced stress σ, may be expressed as S = C σ + σ ' c H log ' 1+ e0 σ C c can be interpreted as the slope of the virgin compression part of the e log p curve obtained from consolidation tests carried out on undisturbed soil samples. The current

40 study is based on selected 81 consolidation test results on undisturbed soil samples from Cincinnati and Northern Kentucky provided by Thelen Associates Inc. (70 test data), H.C. Nutting Company (20 test data), and ODOT (5 test data). Some of the test results were discarded due to dubious nature of their results. The Atterberg limits were available for 65 test data and only these were used for the analysis. The data used to evaluate the correlations exclude the consolidation test results present in the case studies later on in this paper. While the specifics of sampling and testing procedures are not covered exclusively, their effects may be significant. Hence, the results should be viewed with these limitations in mind. 3.2 Characteristics of Data Casagrande (1932) introduced correlations between Atterberg limits and many properties of silts and clays, such as their dry strength, compressibility, consistency, and their reaction to the shaking test by means of a plasticity chart (Figure 3.1). The chart is divided into six regions, three above the A-line and three below it. Figure 3.1 illustrates the number of samples that fall in each category based on the Unified Soil Classification System. This plasticity chart indicates that most of the soils are inorganic clays of either low plasticity or medium plasticity as the collected data. This is in agreement with the majority of soils found in the Cincinnati and Northern Kentucky area as most of the soils consist of glacial till, loess, colluvium, alluvium, and outwash sand and gravel on the inorganic side (Potter, 1996). A summary of statistical parameters of simple soil properties obtained by Azzouz et al. (1976) are given in Table 3.1. The data compiled for

41 this research includes the errors arising from the uncertainties in the operator skill and deviations from the standard testing procedures. Table 3.2 provides a summary of statistical parameters for different soil properties of the available data. The coefficient of variation for different soil properties is a measure of their dispersion and the statistical parameters of liquid limit (LL) and plasticity index (PI) show high dispersion tendencies. The liquid limit has a range of 102 and the plasticity index has a range of These values are very high. The dry unit weight has a range of 66.3 pcf. These values also indicate the vast range of the type of soils in the Cincinnati Northern Kentucky area. The average values obtained for water content, liquid limit, dry unit weight, and initial id ratio match very well with the average values for the soil samples obtained by Azzouz et al. (Table 3.1, 1976). Figure 3.2 to Figure 3.6 presents the plots illustrating the data scatter for various soil parameters. Table 3.3 presents some of the correlations introduced previously by various scientists. The ratio of C c predicted and C c measured, however, is evaluated by the author of this paper based on the data collected. Table 3.4 presents the new correlations obtained during the current research

42 LL = 50 PI = 0.83(LL - 17) R = Plasticity Index (%) Inorganic clays of low plasticity (CL) LL = 30 Inorganic clays of medium plasticity (CL) Inorganic clays of high plasticity (CH) PI = 0.73(LL - 20) A-Line Inorganic silts of high compressibility and Organic clays (MH-OH) Inorganic silts of low compressibility (ML) Inorganic silts of medium compressibility and Organic silts (ML-OL) Liquid Limit (%) Figure 3.1: Classification of soil samples (after Casagrande, 1932)

43 Table 3.1: Summary of statistical parameters for different soil properties of all samples (Azzouz et al., 1976) Soil Property Units Mean Median Standard Deviation Standard Deviation Of Mean Coefficient of Variation Maximum Value Minimum Value Range No. of Samples Liquid limit % Plastic limit % Plasticity Index % Water content % Dry unit weight g/cm Initial id ratio Percent clay Compression index Coeff. Of Consolidation Preconsolidation Pressure cm 2 /sec kg/cm

44 C c = 0.46 (e o ) R = Compression Index (Cc) Initial Void Ratio Figure 3.2: Relationship between compression index and initial id ratio

45 C c = (LL ) R = Compression Index (Cc) Liquid Limit (%) Figure 3.3: Relationship between compression index and liquid limit

46 C c = w n R = Compression Index (Cc) Natural Moisture Content, Wn (%) Figure 3.4: Relationship between compression index and natural water content

47 C c = Y d R = 0.81 Cc Dry Density (pcf) Figure 3.5: Relationship between compression index and dry unit weight

48 C c = PI R = Compression Index (Cc) Plasticity Index, PI (%) Figure 3.6: Relationship between compression index and plasticity index

49 Table 3.2: Summary of statistical parameters for different soil properties of all samples for the current research Soil Parameter Min Max Median Coefficient of Variation Average, µ x Standard Deviation, σ x No. of Samples Water Content, w n (%) Liquid Limit, LL (%) Dry Density, γ d (pcf) Initial Void ratio, e o Specific Gravity, G s Plasticity Index, PI (%) Compression Index, C c Table 3.3: Ratio of predicted and measured compression indices using single and multiple soil parameters Cc = (LL - 10) Equation Reference Applicability Skempton (1944) Remolded clays Cc predicted/cc measured* (0.512) Cc = 1.15 (e o ) Nishida (1956) All clays (0.844) Cc = 0.54 (e o ) Nishida (1956) All clays (0.396) Cc = 0.30 (e o ) Hough (1957) Inorganic cohesive soil; silt; some clay; silty clay; clay (0.257) Cc = (LL - 10) Terzaghi and Peck (1967) Undisturbed, Normally Consolidated clays (0.659) Cc = 0.014PI Cc = 0.01(w n - 5) Cc = 0.141G s 1.2 [(1 + e o )/G s ] 2.38 Cc = (LL/100)G s Nacci et al (1975) Azzouz et al. (1976) Rendon- Herrero (1983) Nagaraj and Murty (1985) North Atlantic clays All clays All clays All clays (0.881) (0.436) (0.319) (0.529) Cc = (LL - 13)/109 Mayne (1980) All clays (0.676) Koppula All clays Cc = 0.01W n (0.560) (1981) *Note: The values in the table are averages with standard deviation in parenthesis

50 Table 3.4: Ratio of predicted to measured compression indices using single and multiple soil parameters for Cincinnati and Northern Kentucky Correlation Cc New Equation (Cincinnati and Northern No. Coefficient, predicted/cc Kentucky) R measured* 1. Cc = 0.46 (e o ) (0.385) 2. Cc = (LL ) (0.475) 3. Cc = w n (0.480) 4. Cc = γ d (0.400) 5. Cc = PI (0.501) 6. Cc = 0.46e o G s (0.379) 7. Cc = e o w n (0.379) 8. Cc = e o LL+ 0.01w n (0.491) *Note: The values in the table are averages with standard deviation in parenthesis

51 3.3 Compression index using a single soil parameter Most of the available regression equations are linear in nature and generally relate to a single soil parameter as summarized in Table 3.3. In the present research, the relationships between compression index and soil parameters are determined for initial id ratio, natural water content, and dry unit weight with high correlation coefficients. The relationship between the compression index and the plasticity index gives the least value of correlation coefficient. This is expected due to a significant scatter present in the plot (Figure 3.6). A good correlation coefficient value of 0.88 is obtained between the compression index and initial id ratio. The average values and the standard deviations (in parenthesis) of the ratios of compression index derived from the regression equation to the measured compression index are (0.385), (0.480), and (0.400), respectively, where the initial id ratio, natural water content, and dry unit weight are a function of the equation. The new correlations (Table 3.4) were obtained with the help of MS Excel. The best fit curves between different soil properties were determined by the least squares method which provides an optimum fit for the data because it minimizes the sum of the squares of the ordinate differences. A summary of available correlations is provided in Table 3.3 with the average values of the ratio of predicted to measured compression index with standard deviation in parenthesis. It shows that most of the existing correlations either underpredict or overpredict the settlement. While it is more conservative to overpredict the settlement, it should not be significantly large. In this respect, Nishida s (1956), Mayne s (1980), and Azzouz et al. (1976) correlations provide good estimations for the consolidation settlement for Cincinnati and Northern Kentucky soils. All the new

52 equations inlving correlations between compression index and soil parameters overpredict the settlement by a small amount and most of these equations provide statistically significant results. 3.4 Compression index using multiple soil parameters Table 3.3 summarizes some of the equations inlving multiple soil parameters obtained by previous researchers. The regression equations, as proposed by Renden-Herrero (1983) which is a function of specific gravity of solids and initial id ratio, underestimates the compression index for Cincinnati and Northern Kentucky soils while the correlation obtained by Nagaraj and Murty (1985) overestimates it. Under such circumstances, it is more plausible to use the latter equation in order to predict the settlements, however, the new correlations obtained in this analysis provide lower averages and standard deviations for the ratio of compression index derived from the regression equation to the measured compression index for various combinations of initial id ratio, natural water content, and liquid limit. In the equations inlving multiple soil parameters (Table Equations 6, 7, and 8), standard errors of , , and have been obtained respectively. Therefore, prediction of the magnitude of primary consolidation using the currently derived equations in lieu of the correlations from previous studies is statistically improved and recommended for the soils of Cincinnati and Northern Kentucky

53 4. Case Studies 4.1 Introduction Six project reports were obtained from H. C. Nutting Company. These reports had most of the required information (including the monitoring data) to enable a complete settlement analysis to be undertaken. The reports have been studied and presented hereafter in the form of case studies by the author of this paper. These case studies start with a brief summary of the project, present the geologic and subsurface conditions based on the test borings, provide the recommendation of the project engineers inlved, and then an analysis of the settlement. This settlement analysis inlves the evaluation of settlements using the newly obtained correlations and their comparison with the estimated and the actual recorded values. There have been a few cases without all the required data for a proper analysis because some of the projects are 30 to 40 years old and therefore some information has been lost. The author of this paper made several assumptions based on the available information and his judgment to evaluate the settlement value. Further, the results of consolidation tests carried out in the reports used in the case studies were excluded in the formulation of new correlations

54 4.2 Case A* - The Expansion of Warehouse and Distribution Facility in Claryville, KY. Figure 4.1: Location of warehouse and distribution facility in Claryville, KY Project Summary The project site is located to the east of US 27 in Claryville, KY. The project inlved the expansion of an existing warehouse and distribution facility to the west, covering an area of approximately 155 ft x 295 ft in size. The floor slab of the proposed facility was to be at the same elevation as that of the existing structure. The maximum column loads were estimated to be approximately 80 kips, with floor loadings of approximately 2000 psf by the author of the project report. *This case study is based on the report written by Jerome B. Kenkel, P.E., of the H. C. Nutting Company, Cincinnati

55 4.2.2 Geology and Subsurface Conditions The area was formed due to the deposition of fine textured sediments in water with less energy often known as lake bed clays. As the glaciers melted, lakes were formed in the areas of low elevation surrounded by hills and caused the deposition of sediments in the ponded water. The natural soils consist of a stiff to hard sandy lean clay containing gravel, which is glacial till, stiff to very stiff fat plastic clays, and thick lakebed deposits consisting of stiff to very stiff varved clays. Alternating layers of medium gray inorganic silt and darker silty clay with low values of shear strength form varved clay. All borings terminated in very highly weathered soft shale bedrock of Ordovician age belonging to the Kope Formation. The Kope Formation is the thickest of all outcropping formations in the Cincinnati region being 200 to 250 feet thick with 65 to 80 percent shale (Potter, 1996). The existing structural fill, the fat clays and the varved lakebed clays are soils with relatively high plasticity. Soil moisture control is critical because of the potential damage to the foundation/structure in the event of shrink-swell. The three borings drilled for the expansion encountered stiff to very stiff soil conditions. Hole B-1 (Figure 4.2) indicated that the structural fill placed during the previous construction in 1982 generally had a very stiff consistency. For the new facility, 2 ft. to 10 ft. of structural fill was required to reach the existing floor. A substantial amount of this fill came from the surrounding hillsides and valleys as per the project engineer s recommendation. A large fraction of this soil included the fat clays and lakebed clays and they had high shrink-swell potential

56 Figure 4.2: Subsurface profile based on the borehole logs (1989)

57 4.2.3 Conclusions and Recommendations by the Project Engineer Site Preparation It was recommended that the site be graded in a manner to provide a suitably compacted fill for footing, floor slab, and pavement support. All the top soil was to be stripped from the site within building and pavement areas to at least 10 ft. beyond exterior building lines. The exposed surface was then compacted to ±2% of optimum moisture content to provide a suitably dense subgrade to the superstructure, and for the overlaying of fill. It was recommended for the fill to be compacted to at least 98% Standard Proctor Moisture- Density, ASTM D 698 after being placed in loose 8 inches lifts and with a least slope of 3H:1V. Foundations The project engineer recommended that the warehouse be borne by continuous and isolated spread footings. Due to the shrink-swell potential of the soil, it was recommended that all the exterior footings bear a minimum of 3.5 feet below the finished grades where the moisture fluctuation becomes progressively minimal. It was strictly suggested to the contractor not to expose the footing excavations to standing water during construction due to the lume change characteristics of the soils present on the site. Any water infiltration was to be blocked during concrete placement for the footings. It was recommended that the concrete placement be done without losing time after the footing excavations in order to protect the soils from being dried out or gain any moisture. A design of shallow footings was recommended using a maximum allowable soil pressure of 4 ksf

58 Settlement Calculations were made by the project engineer by assuming the best and the worst soil conditions encountered at the site. The settlements of the floor slab and column were anticipated to be the most where the least amount of fill was placed and vice versa. An estimate of approximately 1 inch (maximum) settlement was made by the engineer along the west building line due to the weight of 10± feet of newly placed structural fill. Settlement computations inlved a ft 2 footing, with a 4 ksf soil pressure under 80 kips column load. The computations indicated settlements in the range of 1 inch to 1.5 inches. A summary of the anticipated settlements is provided in Table 4.1 assuming a 35 feet x 33 feet floor bay. Surcharge Loading A recommendation was made by the engineer to place an 8 feet high soil surcharge over the entire Phase I building area (Figure 4.3), i.e., approximately 150 ft. x 295 ft. for about 2 to 4 weeks. This surcharge could then be used as a fill for the Phase II area (Figure 4.4). This surcharge was recommended to help attain most of the settlements and minimize differential settlements between adjacent columns and floor bays, however, in the opinion of the author, 2 to 4 weeks is not enough time for substantial settlements to occur. The computed settlement under the weight of an 8 feet surcharge load was approximately 1.3 to 1.6 inches as determined by the engineer. Installation of 3 to 5 settlement plates was recommended at the subgrade level before the placement of surcharge

59 Table 4.1: Predicted floor slab settlement Floor Loading (psf) Settlement (inches), with thickest amount of structural fill below slab and above natural soil Settlement (inches), with least amount of structural fill below slab and above natural soil

60 Figure 4.3: Test Boring Location and Settlement Plates plan

61 Table 4.2: SP-1 Settlement Plate Readings Date Top of Pipe Elevation (ft) /Day (ft) Total Cumulative (ft) Total Settlement (inches) Fill Elev. (ft) Total Fill Height (ft) April April April April May Pipe Extension Added May May May June June June June June June June

62 Table 4.3: SP-2 Settlement Plate Readings Date Top of Pipe Elevation (ft) /Day (ft) Total Cumulative (ft) Total Settlement (inches) Fill Elev. (ft) Total Fill Height (ft) April April Pipe Possibly Disturbed by Earth Moving Equipment April May May May Pipe Extension Added May May June June June Stopped Collecting Data on June 11. Settlement Plate is Located Near Interface of West Surcharge and East Surcharge Pipe is Leaning Downslope Towards East Surcharge Area. Readings from June 5 to June 11 are Considered Highly Questionable

63 Table 4.4: SP-3 Settlement Plate Readings Date Top of Pipe Elevation /Day (ft) Total Cumulative (ft) Total Settlement (inches) Fill Elev. (ft) Total Fill Height (ft) (ft) April April April Pipe Possibly Disturbed By Earth Moving Equipment Verify On 5/27/90 April 27 Pipe Destroyed By Earth Moving Equipment - SP-3 Replaced On April 27, New Datum Started. April May 9 Pipe Disturbed By Earth Moving Equipment. New Datum Set On May 9, May May May Pipe Extension Added. May May June June June June 22 No Reading Taken. Pipe Was Disturbed By Earth Moving Equipment Between June 19 June 22, Pipe is Leaning Sharply. No Future Readings To Be Taken SP-3 is Now Considered Void

64 Table 4.4: SP-4 Settlement Plate Readings Date Top of Pipe Elevation (ft) /Day (ft) Total Cumulative (ft) Total Settlement (inches) Fill Elev. (ft) Total Fill Height (ft) April April April May New Datum Set May 9, 1990 = ft. Stop Collecting Data SP-4 Will Be Relocated. June 19 New Datum Started: SP-4 Established In New Location June Pipe Possibly Disturbed By Earth Moving Equipment Top Of The Pipe Is Cracked. SP-4 Is Located In East Surcharge Area. June June

65 Table 4.5: SP-5 Settlement Plate Readings Date Top of Pipe Elevation (ft) /Day (ft) Total Cumulative (ft) Total Settlement (inches) Fill Elev. (ft) Total Fill Height (ft) May May May May June June June June June June June

66 Figure 4.4: Test Boring Location Plan

67 4.2.4 Analysis Table 4.2 to Table 4.5 presents the actual settlement readings registered by the settlement plates indicating a maximum settlement of 4.08 inches and a minimum settlement of 1.56 inches. Figure 4.2 illustrates the location of the settlement plates. Settlement plates SP-1 and SP-5 were located very close to each other and plates SP-2 and SP-3 at the south-east and south-west sides respectively of the Phase I building area (Figure 4.3). Plate SP-4 was installed along the building line at the eastern part of the Phase I area (Figure 4.3). The predicted value of the settlement under the weight of an 8 feet surcharge was approximately inches. The settlement plate readings provided from Tables 4.2 to 4.5 shows that maximum settlements corresponding to plates SP-1, SP-2, SP-3, SP-4, and SP-5 are 4.08" for approximately 15 feet of fill, 1.92" for 7.44 feet of fill, 1.56" for feet of fill, 2.04" for 3.85 feet of fill, and 2.76" for 7.51 feet of fill. Considering the fact that the prediction was made for 8 feet of fill, most of the observed settlements were significantly greater than the predicted range of 1.3" 1.6". These values also indicate that there is a need to review the current methods and/or equations being used to predict the settlements in regions of Cincinnati and Northern Kentucky or at least for this site. Nevertheless, the settlement plates are not a very reliable source of measurement because they are always prone to erroneous results in the event of improper installation, neglect, and inadequate maintenance. Further, the report acquired by the author of this paper didn t contain any consolidation test data therefore it was difficult to evaluate the settlements based on new equations (Table 3.3). However, the author of this paper did try to estimate the settlement due to the surcharge based on the following assumptions: 8 feet high surcharge with a wet unit weight of 125 pcf

68 The thickness of compressible layer was assumed to be 40 feet, divided into three sub-layers with different dry unit weights based on the three borehole reports available. A worst case scenario was considered (Figure 4.5). The first two layers were assumed to be overconsolidated (owing to the geologic history of glacier presence and its subsequent retreat) while the third layer was assumed to be normally consolidated based on the id ratio. The preconsolidation pressure was estimated from the following equation: S u p c = 0.23 ± 0.04 (Jamiolkowski et al., 1985, for lightly overconsolidated soils). S u = undrained shear strength p c = preconsolidation pressure. The dry unit weights and the initial id ratios were obtained from the unconfined compression test data from the project report. The estimated settlement ranged from 5.9 to 7.2 inches with the use of new equations inlving initial id ratio, dry unit weight, and the natural water content. The values obtained are greater than the actual values observed, however, it should be noted that significant settlements were still taking place when the settlement monitoring program was abandoned. In the last 6 days (Table 4.2 through 4.5), plate 5 experienced 0.48 inch of settlement, plate 4 recorded 0.84 inch of settlement, and plate 1 recorded 0.60 inch of settlement. Plates 2 and 3 recorded 1.2 and 0.36 inches of settlement in the last five days respectively. Hence it is the opinion of the author of this paper that the monitoring program may have been abandoned too soon

69 Sample Evaluation: Area of surcharge: LxB = 150 x 295 γ m = 126 pcf; e o = ; Overconsolidated γ = 102 pcf d Surcharge; γ m = 125 pcf γ m = 122 pcf; e o = ; Varved Clay Overconsolidated; γ d = 94.1 pcf γ m = 116 pcf; e o = ; Varved Clay Normally Consolidated; γ d = 82.1 pcf γ m = Wet Unit Weight γ = Dry Unit Weight d Figure 4.5: Assumed soil profile for settlement calculation The stress distribution by depth is assumed to be constant (= 1000 psf) throughout due to the large extent of the area when compared to the depth. S u p c = 0.23 ± 0.04 (Jamiolkowski et al., 1985, for lightly overconsolidated soils). S u = undrained shear strength p c = preconsolidation pressure. Average S u = 1.27 tsf = 2540 psf (from unconfined compression test results) p c = 9407 psf 1. Use, C c = 0.46 (e o ) Layer 1: C c = 0.46 (e o ) = 0.185; C r = 0.2 C c = Layer 2: C c = 0.46 (e o ) = 0.250; C r = 0.2 C c =

70 Layer 3: C c = 0.46 (e o ) = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = ft = 6.2 inches 2. Use, C c = γ d Layer 1: C c = γ d = 0.173; C r = 0.2 C c = Layer 2: C c = γ d = 0.245; C r = 0.2 C c = Layer 3: C c = γ d = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = ft = 5.9 inches 3. Use, C c = w n ; w n = natural water content Layer 1: C c = w = 0.201; C r = 0.2 C c = n Layer 2: C c = w = 0.285; C r = 0.2 C c = Layer 3: C c = w = n n

71 Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = ft = 7.2 inches σ + σ i * e = C c log for NC clays and σ σ + σ i * e = C r log for OC clays where σ + σ i p c σ e *S = H i 1+ e

72 4.3 Case B*- Large One-Story Structure in Newport, KY. Figure 4.6: Location of large one-story structure in Newport, KY Project Summary The project site is located to the east of I 471 and south of the Ohio River in Newport, KY (Figure 4.6). The project inlved a one story building covering an area of approximately 300 ft x 400 ft in size. Two methods of providing the roof were considered by the author of the project report: (a) The maximum column loads were estimated to be approximately 35 kips and uplift forces of 8 kips utilizing a flexible roof support on cables; (b) A conventional long-span steel joist spanning between a steel frame structure with loads of approximately 125 kips. *This case study is based on the report written by James J. Flaig, P.E., of the H. C. Nutting Company, Cincinnati

73 4.3.2 Geology and Subsurface Conditions The site is located within the present Ohio River Valley. It was observed that the bedrock exists 80 to 90 feet below the grade through the eight test borings drilled (Figures 4.7, 4.8, and 4.9). The bedrock is the typical local Cincinnati Series layered gray shale and limestone of Ordovician age. Five to ten feet of glacial till (stiff to very hard sandy silty clay with gravel and rock fragments) was encountered above the bedrock in borings 5, 6, and 8. Medium dense outwash sand and gravel was encountered above the glacial till and bedrock for the most part. At several of the test borings above the sand, stiff to very stiff silty clay was encountered. Rubble fill was encountered from 5 to 15 feet at test borings 1, 2, 6, 7, and 8. This fill consist of clay, sand, concrete and limestone slabs, cinders, brick, glass, metal, and wood. Ground water was encountered typically at a depth of 30 to 35 feet with the exception of test boring number 5 where it was encountered at a depth of 12 to 13 feet Conclusions and Recommendations by the Project Engineer The following (but not limited to) gives the conclusions and recommendations provided by the project engineer: Two methods of construction were proposed. The first one had a structural reinforced concrete deck supported on columns with piles or piers as foundations. The second method inlved the placement of structural fill at the site to support the proposed structure. The latter method was recommended due to economic reasons

74 Figure 4.7: Boring Locations

75 Figure 4.8: Subsurface Investigation

76 Figure 4.9: Legend for Figure 4.8 (Redrawn)

77 The existing rubble fill was to be removed from the building area to a distance of 15 to 20 feet outside of the area for the placement of structural fill. Approximately 15 feet of fill was recommended. Based on laboratory consolidation tests, 6 to 7 inches of settlement was estimated at test boring number 5 where there was approximately 26 feet of soft to medium stiff silty clay and 1 inch at test boring number 6. It was predicted that settlements would range from 1 to 6 inches depending on the compressibility of soil at various site locations within two to three months after the placement of surcharge. The following recommendations were made for the fill specification: o The weed and grass cover was to be stripped before the fill placement. o The top 6 inches of the flood plain was to be harrowed and compacted before the fill placement when the weather was dry. o The fill was to be placed in 8 inch loose lifts and then compacted with a 20 ton sheepsfoot roller or equivalent to at least 98% of maximum density determined by the Standard Proctor Method, ASTM D 698. o Settlement plates were to be set near the boring locations after the fill reached a height of 2 ft. above the prepared starting surface. o At least 4 feet of surcharge was to be placed after the placement of 15± feet of structural fill was complete. This surcharge fill was to be placed in 12 inch loose lifts and monitored for surface water infiltration. o It was recommended that the fill materials should be approved by a geotechnical engineer prior to placement and compaction based on laboratory moisture-density tests

78 o All the rubble fill material excavated from the site was to be wasted elsewhere. o All fill placement and compaction including preparation before filling was to be carried out under the supervision of a geotechnical engineer and continuous inspection of a qualified soil technician. All the footings were to be backfilled with either cohesive or granular material and then compacted to at least 95% of the Standard Proctor Value. It was recommended that finished grades be sloped away from the building. A ditch should be provided along the west side within the structural fill where it intersects the existing slope Analysis The consolidation tests were carried out on the soil samples obtained from boring number 5 at depths of 2 4 ft and ft. The tests indicated that the soil was overconsolidated at depth 2-4 ft. but normally consolidated at ft. The overconsolidation in the upper 2 to 4 ft was most likely desiccation. The author of this paper estimated the settlements to be ft at the location of test boring number 5 using the new correlations and based on the following assumptions (Figure 4.10): 15 feet high surcharge with a wet unit weight of 120 pcf. The thickness of the compressible layer was assumed to be 26 feet (as assumed by the project engineer of the actual report), divided into two sub-layers with ground water level at 12.5 feet as per boring number

79 The clay was normally consolidated as per the report of the consolidation test at a depth of ft. It was assumed to be normally consolidated for all but top 5 feet of the compressible layer. The top five feet was assumed to be overconsolidated. The dry unit weights and the initial id ratios were obtained from the consolidation test data from the project report and assumed as shown in the profile (Figure 4.10) for the total thickness except for the variation in unit weight caused by the ground water level. Figure 4.14 presents the settlement plate readings. Plate number 5, which was close to the location of boring number 5, gives the maximum value of settlement. It recorded settlement up to 1.2 ft as opposed to the estimated settlement of 6 7 inches by the project engineer. In fact, the settlement observed was more than 9 inches in three months time and reached about 1.2 ft in about 16 months. Sample Evaluation: Area of surcharge: BxL = 300 x 400 Surcharge; γ m = 125 pcf γ m = 115 pcf; e o = OC Clay γ m = 118 pcf; e o = NC Clay γ m = Wet Unit Weight Figure 4.10: Assumed soil profile for settlement calculation for the location of plate no

80 The stress distribution by depth is assumed to be constant (= 1875 psf) throughout due to the large extent of the area when compared to the depth. p c = preconsolidation pressure. p c = 0.75 tsf (from consolidation test on Boring 5, feet deep) p c = 0.80 tsf (from consolidation test on Boring 5, 2-4 feet deep) 1. Use, C c = 0.46 (e o ) Layer 1: C c = 0.46 (e o ) = 0.261; C r = 0.2 C c = Layer 3: C c = 0.46 (e o ) = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = 1.21 feet 2. Use, C c = γ d Layer 1: C c = γ d = 0.271; C r = 0.2 C c = Layer 2: C c = γ d = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = 1.24 feet 3. Use, C c = w n ; w n = natural water content

81 Layer 1: C c = w = 0.230; C r = 0.2 C c = Layer 2: C c = w = n n Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = 1.17 feet σ + σ i * e = C c log for NC clays and σ σ + σ i * e = C r log for OC clays where σ + σ i p c σ * e = C p C σ + σ c i r log + c log for OC clays where σ p c σ < p c < σ + σ i e *S = H i 1+ e

82 Settlement Plate Readings (Newport, KY) Months Apr-75 Jul-75 Nov-75 Feb-76 May-76 Aug-76 Dec-76 Mar Settlement (ft.) Figure 4.11: Generalized representation of settlement plate readings (SP 1 to SP 8)

83 Figure 4.12: Generalized representation of settlement plate locations (SP 1 to SP 8)

84 4.4 Case C*- Five Story Steel Frame Structure, Milford, OH Figure 4.13: Location of the five story steel frame structure, Milford, OH Project Summary The proposed building was a 4-story, steel frame construction with brick veneer without any basement. Measuring approximately 192 ft x 88 ft, exterior and interior column loads were estimated to be 650 kips and 500 kips respectively. Ground floor level was proposed approximately 4 to 23 feet above the existing grades (elevation 607) at the beginning of construction. *This case study is based on the report written by George C. Webb, P.E., of the H. C. Nutting Company, Cincinnati

85 4.4.2 Geology and Subsurface Conditions The construction site was located at the base of the Illinoian till terrace (~300,000 years). This is characterized by minor sand and gravel at higher elevations overlying interbedded till, sand, and gravel and commonly thickest in the preglaciated tributaries of the Ohio River, Mill Creek, and East Fork of the Little Miami River (Potter, 1996). Ten Standard Penetration Test borings were performed (ASTM D 1586) and subsoils consisting of topsoil, glacial till, alluvial, glacial outwash, and lakebed deposits were encountered. Lakebed deposits were encountered only in test boring 17-1 and B-1 (1976). The project engineer observed that the basement outwash and lakebed soils were eroded by glacial meltwater and replaced by alluvium. The SPT blows for alluvium ranged from 2 to 20 and the unconfined compressive strength from 0.5 to 1.9 tsf. The consolidation test carried out on a sample recovered from boring 17-5 (Figure 4.14) indicated that this alluvium was overconsolidated Conclusions and Recommendations by the Project Engineer It was decided to place structural fill at the site to bring the grade up to the required floor elevation and to account for most of the settlement. The fill was to uniformly compacted to at least 98% Standard Proctor (ASTM D 698) and the subgrade be compacted to 100% Standard Proctor. The ground was to be stripped of the top soil and rolled with a Caterpillar 815 self-propelled sheepsfoot roller before placing the fill. Placement of a 7-8 ft high surcharge was recommended over the compacted fill from one corner of the proposed building to another

86 Two settlement plates were set up to monitor the amount and rate of settlement resulting from embankment and surcharge fill. These plates were installed on the original grade prior to placement of any fill Analysis The author of this paper estimated the settlement to be 5.6 to 5.8 inches based on the following assumptions (Figure 4.14) at the location of test boring number 17-5 (Figure 4.15): 33 feet high surcharge with a wet unit weight of 125 pcf. The thickness of the compressible layer was assumed to be 26 feet as per test boring number 17-5 (Figure 4.15). The clay was overconsolidated as per the report of the consolidation test at a depth of ft. The dry unit weights and the initial id ratios were obtained from the consolidation test data from the project report and assumed to be the same for the total thickness of the compressible layer. The results of settlement monitoring program are illustrated in Figures 4.16 and These show settlements of 6.8 inches due to a 33 feet high surcharge and 11.1 inches due to a 24 feet high surcharge at different locations. Thus, the estimated value of settlement obtained by using new correlations compares well with the actual recorded value. Other settlement plate locations and geotechnical summary data are presented in Figures 4.18, 4.19, 4.20, and

87 Sample Evaluation: Area of surcharge: 192 x 88 Surcharge; γ m = 125 pcf 33 γ m = 123 pcf; e o = Compressible Layer 26 γ m = Wet Unit Weight Figure 4.14: Assumed soil profile for settlement calculation for the location of plate # 1 near Test boring hole 17-5 The stress distribution by depth is assumed to be constant (= 4125 psf) throughout due to the large extent of the area when compared to the depth. p c = preconsolidation pressure. p c = 2.5 tsf (from consolidation test on Boring 17-5, feet deep) 1. Use, C c = 0.46 (e o ) Layer 1: C c = 0.46 (e o ) = 0.201; C r = 0.2 C c = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = 0.48 ft = 5.76 inches

88 2. Use, C c = γ d Layer 1: C c = γ d = 0.195; C r = 0.2 C c = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = ft = 5.6 inches 3. Use, C c = w n ; w n = natural water content Layer 1: C c = w = 0.202; C r = 0.2 C c = n Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = ft = 5.78 inches σ + σ i * e = C c log for NC clays and σ σ + σ i * e = C r log for OC clays where σ + σ i p c σ * e = C p C σ + σ c i r log + c log for OC clays where σ p c σ < p c < σ + σ i e *S = H i 1+ e

89 Figure 4.15: Test Boring number

90 Figure 4.16: Settlement recorded at Plate #1-75 -

91 Figure 4.17: Settlement recorded at Plate #2-76 -

92 Figure 4.18: Settlement plate locations

93 Figure 4.19: Summary of Geotechnical data (contd. in the next page)

94 Figure 4.20: Summary of Geotechnical data (contd. from previous page)

95 Figure 4.21: Legend for Figures 4.19 and

96 4.5 Case D*- Two Story Steel Frame Structure. Springdale, OH Figure 4.22: Location of two story steel frame structure. Springdale, OH Project Summary The proposed structure was two storied with ground floor on grade, occupied a 230 x 420 ft area, and had a steel frame with maximum column loads of 154 kips and 30 x 30 ft bay spacing. The site was located in Springdale, north of Cincinnati (Figure 4.22). Previous grades within this portion of the tract ranged from approximately 851 to 863 and the proposed finished floor was to be at *This case study is based on the report written by Merle F. Nethero of the H. C. Nutting Company, Cincinnati

97 4.5.2 Geology and Subsurface Conditions Seven test borings were completed at locations selected by the project engineer (Figure 4.23). The subsurface profiles are presented in Figures 4.24 and They illustrate a complex and varied soil types. Geologically, the site is situated within the area abraded by both the Illinoian and Wisconsin glaciers. The soils at the site ranged from outwash sand and gravel to lakebed clays. The test borings exposed fill, top soil, and alluvium for the first several feet. About 6 feet of soft wet high silt content materials were found within the lower eastern swale. This material was removed for grading purposes. Below this layer existed glacial till with high N-values and shear strength in the west which gradually lessens in the east towards the swale. The above layer is underlain by varved silt, silty clay, and medium stiff unweathered till. At boring E, the lake bed is mixed with outwash sand, gravel, and silt. Layers of dark silt, moderate to highly plastic clays, sand and gravel, and buried glacial logs were also found in this zone. Compact sand and gravel outwash or stiff to hard weathered glacial till (presumably of older Illinoian age) was found around the lowest bottom elevation (804.5) explored. The bedrock was assumed to be at a depth of 70 feet below the average existing grade. The groundwater level was encountered at a depth of 3 to 4 feet below the existing grade in eastern swale areas and at increased depths elsewhere

98 Figure 4.23: Test Boring Location Plan

99 Figure 4.24: Subsurface profile for test borings A, B, E, and F

100 Figure 4.25: Subsurface profile for test borings D, G, and C

101 4.5.3 Conclusions and Recommendations by the Project Engineer Use of spread footing was recommended over piers for economic reasons. Site grading was to be performed under the full time inspection of a soil technician. Removal of trees and brush, stripping of top soil, and undercutting of saturated alluvium within the eastern swale areas were recommended. All structural fill was to be placed in lifts of 6 inch following compaction to a minimum field density of 98% Standard Proctor (ASTM D 698). A net soil pressure of 3-4 psf was recommended for adequate factor of safety against shear failures or excessive settlement caused by the immediate foundation material. A maximum settlement of 4 inches was predicted for the foundation soils on the eastern swale areas due to the superimposed load of the embankment plus building. The minimum predicted settlement in the central and western building areas was 2 inches. These settlements were expected to occur within 60 days of the application of loads. Six settlement plates were installed at locations specified by the project engineer. Placement of additional surcharge was recommended to expedite the settlement process. Consolidation tests carried out on the samples obtained from boring number C (25-27 ft), F (15-17 ft), and G (22-24 ft) indicated the soil to be overconsolidated. The natural soil was to be compacted and covered with a 1 inch thick layer of sand before the placement of the settlement plate. Approximately 3 ft of fill was to be placed and compacted around the plate. After bringing the fill to the finished grade, readings were to be recorded bi-weekly or as per the directions of the soil engineer until the time rate of settlements reached tolerable limits (Tables 4.7, 4.8, and 4.9)

102 Table 4.7: Limiting angular distortion as recommended by Bjerrum (1963) η = angular distortion Table 4.8: Allowable settlement criteria: 1955 U.S.S.R. Building Code (Wahls, 1981)

103 Table 4.9: Allowable average settlement for different building types (Wahls, 1981) Analysis The settlement was estimated for the location of test boring F at the south-east corner of the proposed building because maximum amount of settlement was expected at this location due to the placement of the greatest quantity of fill. The following assumptions were made in this evaluation (Figure 4.26): The wet unit weight of the fill was assumed to be 125 pcf. The ground water level was assumed to be at a depth of 3 feet below the existing grade based on test boring F. The total thickness of the compressible layer was assumed to be 36 feet based on test boring F. The soil was overconsolidated as indicated from the results of three consolidation tests carried out on test borings C, F, and G

104 New equations were used to evaluate the compressibility of soil. The settlement estimated by the author of this paper ranged from 4.7 to 6.7 inches at the location considered. The settlement plate locations, amount of fill placed, and recorded settlement values are provided in Figures 4.27, 4.28, 4.29, and Table The actual settlement recorded for settlement plate #6 illustrates the settlement to be around 6 inches in Figure 4.29 which is close to this author s estimated value. Thus the estimated settlements using the new correlations proved not only accurate but also slightly conservative. Sample Evaluation: Area of surcharge: 230 x 420 Fill; γ m = 125 pcf 24.2 γ m = 131 pcf e o = Compressible Layer 36 γ m = Wet Unit Weight Figure 4.26: Assumed profile for settlement estimation The stress distribution by depth is assumed to be constant (= 3025 psf) throughout due to the large extent of the area when compared to the depth. p c = preconsolidation pressure

105 p c = 1.9 tsf (a lower value chosen from consolidation tests on samples from Borings C, F, and G) σ = effective overburden pressure = 3 (131) + 15 ( ) = 1422 psf (at the mid point of compressible layer) 1. Use, C c = 0.46 (e o ) Layer 1: C c = 0.46 (e o ) = 0.128; C r = 0.2 C c = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = ft 5.5 inches. 2. Use, C c = γ d Layer 1: C c = γ d = 0.110; C r = 0.2 C c = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = 0.39 ft 4.7 inches 3. Use, C c = w n ; w n = natural water content Layer 1: C c = w = 0.157; C r = 0.2 C c = n Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = ft = 6.7 inches

106 σ + σ i * e = C c log for NC clays and σ σ + σ i * e = C r log for OC clays where σ + σ i p c σ * e = C p C σ + σ c i r log + c log for OC clays where σ p c σ < p c < σ + σ i e *S = H i 1+ e

107 Figure 4.27: Settlement plate locations

108 Table 4.10: Recorded depth of fill Figure 4.28: Recorded settlement for plate

109 Figure 4.29: Recorded settlement for plates 1 to 6

110 4.6 Case E*- Aeration Basins. North Bend, OH Figure 4.30: Location of aeration basins. North Bend, OH Project Summary The report inlves two aeration basins along the Ohio River shore in North Bend (Figure 4.30). Settlements and bearing capacity were predicted assuming 11 feet high soil surcharge. Settlement monitoring was conducted. The diameter of the basin was assumed to be 100 feet with 1 ksf design load for evaluation purposes. *This case study is based on the report written by Jess A. Schroeder, P. E. of the H. C. Nutting Company, Cincinnati

111 4.6.2 Geology and Subsurface Conditions The test borings indicated the presence of approximately 30 feet of alluvium underlain by dense sand and gravel. These are mostly believed to be glacial drift of Wisconsinan and Illinoian age. These test borings were probably conducted 10 to 30 years prior to this project. Previously (10 to 30 years earlier) conducted consolidation tests on similar soils and locations illustrated the top 30 feet to be overconsolidated, however, it should be noted that there were no consolidation tests carried out for this specific project Conclusions and Recommendations by the Project Engineer The construction of both the aeration basins was similar with some minor differences. During the excavation of the new aeration basin (# 1), it was recommended to excavate the area with some slope or crowning to provide drainage. Other solutions included installation of sump pumps and minimized traffic and disturbance to the exposed subgrade after final undercutting. The project engineer was to visually inspect the surface before proceeding. It was recommended that a woven geotextile such as MIRAFI 600X (or equivalent) be placed across the base under the ODOT Item 304 compacted stone. As an alternate option, the surface could also be stabilized by placing a layer of No. 2 stone and punching it into the subgrade. The woven fabric was to be placed relatively wrinkle free across the subgrade and unrolled with seam overlaps of 2 to 3 feet. A small dozer was recommended to spread the lifts of the ODOT 304 material and any traffic inlving dump trucks or rubber tired

112 vehicles was prohibited until the fabric and a sufficient amount of stone was in place to stop any noticeable deflection. Three settlement plates (Figure 4.33) were to be placed across the footprint of each of the aeration basins. These plates were roughly aligned perpendicular to the river, one at the center while the other two were to be placed 15 to 20 feet inside the perimeter of each of the basins. The settlement plate and the initial 5 feet section of PVC pipe was to be installed after the placement of the geotextile and no more than 1 foot of ODOT 304 compacted granular material. It was recommended that readings be taken at least every other day during the ODOT filling process and twice per week during and after the placement of surcharge Analysis Settlements were predicted for three cases by the project engineer to be: 1. Under design loading (1 ksf): 8.2 inches 2. Undercut 6 feet and replaced with ODOT 304 (design loading): 4.8 inches 3. Same as 2 with an additional soil surcharge 11 feet high: 7.0 inches Although the consolidation test results were not recent and they indicated an overconsolidated soil profile for the top 30 feet, the project engineer predicted the settlements by assuming the profile to be normally consolidated. This was probably a conservative estimate since there was no consolidation test carried out for this project but the available data belonged to tests that were conducted previously on similar locations. However, the author of this paper assumed only the bottom layer to be normally

113 consolidated in his computations (Figure 4.31 and 4.32). The dry unit weights used in the following computations are the same as the ones used in the original report. Initial id ratios are used from the consolidation test results referenced in the original report. Sample Evaluation: 2 Area of surcharge: π R = 7854 sq. ft. Profile 1: Under design loading (1.0 ksf) p = 1000 psf; Circularly loaded area (100 dia.) Lean Clay; Overconsolidated γ = 122 pcf ; e0 = m 13 Alluvium; Normally Consolidated γ = 120 pcf ; e0 = m 16 Figure 4.31: Simplified subsurface profile 1 under Aeration Basin p c = preconsolidation pressure. p c = 6400 psf.layer 1 p c = 8000 psf.layer 2 1. Use, Cc = 0.46 (e o ) Layer 1: Cc = 0.46 (e o ) = 0.161; Cr = 0.2 Cc =

114 Layer 2: Cc = 0.46 (e o ) = 0.150; Cr = 0.2 Cc = 0.03 Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = ft = 6.3 inches 2. Use, Cc = γ d Layer 1: Cc = γ d = 0.191; Cr = 0.2 Cc = Layer 2: Cc = γ d = 0.164; Cr = 0.2 Cc = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = 0.56 ft ~ 7.0 inches

115 Profile 2: Under 11 (1.3 ksf) feet surcharge and 6 feet undercut (ODOT 304) p = 1300 psf; Circularly loaded area (100 dia.) Layer 1 Lean Clay; Overconsolidated γ = 122 pcf ; e0 = m ODOT 304 γ m = 125pcf 6 7 Layer 2 Alluvium; Normally Consolidated γ = 120 pcf ; e0 = m 16 Figure 4.32: Simplified subsurface profile 2 under Aeration Basin 1. Use, Cc = 0.46 (e o ) Layer 1: Cc = 0.46 (e o ) = 0.161; Cr = 0.2 Cc = Layer 2: Cc = 0.46 (e o ) = 0.150; Cr = 0.2 Cc = 0.03 Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = ft = 6.15 inches 2. Use, Cc = γ d Layer 1: Cc = γ d = 0.191; Cr = 0.2 Cc = Layer 2: Cc = γ d = 0.164; Cr = 0.2 Cc =

116 Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = ft = 6.8 inches σ + σ i * e = C c log for NC clays and σ σ + σ i * e = C r log for OC clays where σ + σ i p c σ * e = C p C σ + σ c i r log + c log for OC clays where σ p c σ < p c < σ + σ i e *S = H i 1+ e 0 It is clear from the settlement plate readings (Figure 4.34 and 4.35) that the assumption about the soil being normally consolidated proves to be right. This is also one of the best examples of the use of engineering judgment found during this research. The predicted values of settlement compare well with the observed values but for plate E (Figure 4.34). This may be due to the presence of a soft spot under that settlement plate

117 Figure 4.33: Test boring and settlement plate location plan

118 Figure 4.34: Settlement plate readings for aeration basin #2 (Re-drawn)

119 Settlement Plate Monitoring Aeration Basin 1 20 Height of Fill Above Bottom-of-Undercut (ft.) Settlement (inches.) Days Plate D Plate E Plate F Fill Figure 4.35: Generalized representation of settlement plate readings for aeration basin #

120 4.7 Case F*- Large One-Story Structure. Cincinnati, OH Figure 4.36: Location of large one-story structure. Cincinnati, OH Project Summary The location (Figure 4.34) includes areas both east and west of Paddock Road, south of Seymour Avenue but this report addresses only a portion of the area east of Paddock Road. The proposed building was a single level, 600,000 sq. ft. structure with expected floor loading to be around 150 psf and interior column loads around 150 kips. The finished floor elevation was assumed to be at 551 with a subgrade at 550. *This case study is based on the report written by C. R. Lennertz, P. E. of the H. C. Nutting Company, Cincinnati

121 4.7.2 Geology and Subsurface Conditions The site was divided into three principal areas on the basis of topography and soil conditions: deep valley, buried slopes, and exposed slopes and upland area. According to C. R. Lennertz, The soils forming the upland area and underlying the slopes were deposited during the Illinoian glacial period and include till, outwash, and lake bed clays. The valley was probably formed during the recession of Illinoian ice. The uppermost soils encountered in the valley were basically fine textured alluvium. The typical profile encountered in the deep valley consisted of about 10 feet of alluvium underlain by about 30 feet of lake bed clays. The unconfined compressive strengths of the alluvium ranged from 2 to 4 ksf and that of the lake bed clays around 1.5 ksf. Two consolidation tests conducted on samples from this clay indicated that the clay was overconsolidated. However, with the increase in the depth and in the middle to south part of the site, the soil was found to be normally consolidated. The material underlying steep exposed hillsides and the upland area was compact glacial till, well graded material consisting of clay through gravel sizes, and at least one layer on a heavily overconsolidated Illinoian silty lake bed deposit. The groundwater level varied considerably in all the borings and was particularly shallow closer to the Bloody Run Channel. The sand layers beneath the buried slopes were free of water due to the high permeability of the sands

122 4.7.3 Conclusions and Recommendations by the Project Engineer It was recommended that the high ground be excavated within the northwest and southwest parts of the site and placed in a controlled manner in the valley area to develop a building pad and parking areas since the valley area is underlain by soft normally consolidated clay extending to about 40 feet below the grade. Settlement of the ground surface as the result of the weight of the compacted fill was estimated to be as much as 15 inches over the deep valley and less than 6 inches over the buried slopes adjacent to the valley. Removal of around 8 inches of top soil and any other unsuitable materials was recommended before grading. A 20% shrinkage factor was recommended for determining the quantity of compacted fill that could be obtained from a unit quantity of excavation. It was thought that the dissipation of excess pore pressure would require a minimum of 15 months if no other measure was taken to increase the rate of drainage from the compressible clay. A settlement of 4 inches was assumed in case there was no preloading undertaken after the dissipation of excess pore pressure due to the placement of compacted fill. A 6 feet high surcharge with an average load of 750 psf was recommended. In case of unavailability of sufficient earth fill required for a future expansion area and parking areas so that the temporary surcharge could be placed over the entire building area at one time, the surcharge could then be placed on one-half of the area and then moved to the other half

123 It was recommended that preliminary estimates for PV (wick) drains be based on 10 feet spacing within the deep valley area and an increased spacing in the adjacent buried slopes. Installation of wick drains was recommended to increase the rate of settlement. An allowable soil bearing pressure was recommended under full dead and live loads Analysis The subsurface profiles assumed for boring B-20 and B-25 is shown in Figure The settlements estimated with the use of newly obtained correlations were in the range of inches which is close to the actual recorded maximum settlement of 8.4 inches (Table 4.11) after six months of monitoring. The ground water level in B-20 was encountered at 3 feet below the surface while it was 5 feet deep for B-25. The predicted settlements for the valley area by the project engineer, based on a 1000 psf surcharge and consolidation tests conducted mostly on the samples from test borings B-20 and B-25 lie in the range of 9 15 inches. Sample Evaluations: Area of surcharge: LxB = 600,000 sq ft The stress distribution by depth is assumed to be constant (= 1000 psf) throughout the area with depth. p c = preconsolidation pressure. p c = 2000 psf

124 p = 1000 psf 3 Fill; Overconsolidated γ m = 120 pcf ; C 1+ e0 = 0.04 Alluvium; Overconsolidated γ = 99 pcf ; e0 = 1.78 m c Lake Bed Clay; Overconsolidated γ = 117 pcf ; e0 = 1.18 m Lake Bed Clay; Normally Consolidated γ = 108pcf ; e0 = 1.28 m Figure 4.37: Simplified subsurface profile for B Use, C c = 0.46 (e o ) Layer 1: The initial id ratio was not provided in the report. Only the C c /1+e o was given. Layer 2: C c = 0.46 (e o ) = 0.690; C r = 0.2 C c = Layer 3: C c = 0.46 (e o ) = 0.414; C r = 0.2 C c = Layer 4: C c = 0.46 (e o ) = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = 0.97 ft = 11.6 inches

125 2. Use, C c = γ d Layer 1: The dry unit weight was not provided in the report. Only the C c /1+e o was given. Layer 2: C c = γ d = 0.550; C r = 0.2 C c = Layer 3: C c = γ d = 0.379; C r = 0.2 C c = Layer 4: C c = γ d = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = ft = 10.3 inches p = 1000 psf 5 Fill; Overconsolidated γ m = 120 pcf ; C 1+ e0 = 0.04 Alluvium; Overconsolidated γ = 97 pcf ; e0 = m c Lake Bed Clay; Overconsolidated γ = 106 pcf ; e0 = m Lake Bed Clay; Normally Consolidated γ = pcf ; e0 = 1.17 m Figure 4.38: Simplified subsurface profile for B

126 1. Use, C c = 0.46 (e o ) Layer 1: The initial id ratio was not provided in the report. Only the C c /1+e o was given. Layer 2: C c = 0.46 (e o ) = 0.740; C r = 0.2 C c = Layer 3: C c = 0.46 (e o ) = 0.525; C r = 0.2 C c = Layer 4: C c = 0.46 (e o ) = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft) Therefore, S total = = 1.0 ft = 12 inches 2. Use, C c = γ d Layer 1: The dry unit weight was not provided in the report. Only the C c /1+e o was given. Layer 2: C c = γ d = 0.578; C r = 0.2 C c = Layer 3: C c = γ d = 0.453; C r = 0.2 C c = Layer 4: C c = γ d = Layer Hi (ft) σ (psf) σ i (psf) e* S* (ft)

127 Therefore, S total = = ft = 10.5 inches σ + σ i * e = C c log for NC clays and σ σ + σ i * e = C r log for OC clays where σ + σ i p c σ * e = C p C σ + σ c i r log + c log for OC clays where σ p c σ < p c < σ + σ i e *S = H i 1+ e 0 Table 4.11: Recorded settlements Settlement Plate Recorded Settlement (inches) SP1 1.7 SP2 1.7 SP3 1.7 SP4 5.8 SP5 4.7 SP6 8.4 Figure 4.39 illustrates the location of settlement plates 1 through 6 on the site. A figure containing the boring location plan was not available and therefore, not included. Plates 1, 2, and 3 were installed on January 22, 2001 and plates 4, 5, and 6 were installed on January 26, The final reading on each plate was obtained on June 30,

128 The recorded settlement of plates 1 through 3 was very uniform varying from 1.69 to 1.72 inches. Settlements observed at plates 4 through 6 were much greater varying from 4.7 to 8.4 inches. The observations also indicated that the settlements were still continuing albeit at a much smaller rate. The total amount of fill placed at the 6 locations varied from approximately 9 feet to 10.5 feet, the final 6 feet of this being temporary surcharge. The author of this paper assumed the profile below SP1-SP6 to be B-20/B-25 with the help of the calculations carried out in the original project report

129 Figure 4.39: Settlement plates location plan