Research Staff. -Environmental Services Steven G. Richardson. -Beneficiation Robert S. Akins

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2 The Florida Institute of Phosphate Research was created in 1978 by the Florida Legislature (Chapter , Florida Statutes) and empowered to conduct research supportive to the responsible development of the state's phosphate resources. The Institute has targeted areas of research responsibility. These are: reclamation alternatives in mining and processing, including wetlands reclamation, phosphogypsum storage areas and phosphatic clay containment areas; methods for more efficient, economical and environmentally balanced phosphate recovery and processing; disposal and utilization of phosphatic clay; and environmental effects involving the health and welfare of the people, including those effects related to radiation and water consumption. FIPR is located in Polk County, in the heart of the central Florida phosphate district. The Institute seeks to serve as an information center on phosphate-related topics and welcomes information requests made in person, by mail, or by telephone. Executive Director Richard F. McFarlin Research Staff Research Directors G. Michael Lloyd Jr. -Chemical Processing Gordon D. Nifong -Environmental Services Steven G. Richardson -Reclamation Hassan El-Shall -Beneficiation Robert S. Akins -Mining Florida Institute of Phosphate Research 1855 West Main Street Bartow, Florida (863)

3 EVALUATION AND PHOSPHATIC CLAY DISPOSAL AND RECLAMATION METHODS Volume 8: Predictive Methodology Applied to Case Histories Research Project FIPR Final Report, December 1991 Prepared by Ardaman & Associates, Inc South Orange Avenue Orlando, Florida Principal Investigators Anwar E.Z. Wissa Nadim F. Fuleihan Thomas S. lngra Mohamed M. Alawi Prepared for Florida Institute of Phosphate Research 1855 West Main Street Post Office Box 877 Bartow, Florida FIPR Program Manager Hassan El-Shall

4 DISCLAIMER The contents of this report are reproduced herein as received from the contractor. The opinions, findings and conclusions expressed herein are not necessarily those of the Florida Institute of Phosphate Research, nor does mention of company names or products constitute endorsement by the Florida Institute of Phosphate Research. ii

5 EVALUATION OF PHOSPHATIC CLAY DISPOSAL AND RECLAMATION METHODS Research Projects FIPR , FIPR and FIPR PREFACE As part of a Florida Institute of Phosphate Research project FIPR titled Evaluation of Phosphatic Clay Disposal and Reclamation Methods, Ardaman & Associates, Inc. performed a comprehensive study to evaluate the engineering properties of a wide range of phosphatic clays and sand-clay mixes, and developed a methodology for forecasting the performance of phosphatic clay settling areas during disposal and reclamation. The findings of the Phase I study (Research Project FIPR ) were presented in a series of six complementary volumes. Subsequently, a Phase II study was initiated (Research Projects FIPR and ) to evaluate the engineering properties of flocculated phosphatic clays, refine the predictive capability, and compare and evaluate the relative merits of various disposal methods on the basis of parametric studies illustrated via an example of a model mine. Findings from the Phase II study are presented in a series of four additional volumes. Laboratory evaluations of the engineering properties of phosphatic clays and sand-clay mixes were performed on phosphatic clays from twelve different mine sites. Volumes 1, 2 and 3 titled Index Properties of Phosphatic Clays, Mineralogy of Phosphatic Clays, and Sedimentation Behavior of Phosphatic Clays, respectively, present extensive data on the twelve clay sources selected in the study. The findings were used to screen the samples and select six clays covering the full range of anticipated behavioral characteristics. The selected clays were subjected to a comprehensive testing program for determining engineering parameters pertaining to consolidation and strength. Extensive sophisticated testing of three of the six phosphatic clays and corresponding sandclay mixes was subsequently undertaken. The results are presented in Volumes 4 and 5 titled Consolidation Behavior of Phosphatic Clays and Shear Strength Characteristics of Phosphatic Clays, respectively. Concurrent with the laboratory evaluation of phosphatic clay engineering properties, a theoretical model to evaluate disposal systems was developed. The finite difference program SLURRY can also be used in reclamation planning. In an attempt to verify and refine the prediction modeling technique, a preliminary field investigation program at six phosphatic clay settling areas ranging from retired to active sites was undertaken. Volume 6 titled Predictive Methodology for Evaluating Disposal Methods discusses the theoretical model and presents a comparison of predictions based on laboratory data and field measurements. To allow evaluation of flocculated clay disposal methods in addition to the conventional and sand-clay mix disposal methods, a laboratory testing program iii

6 was undertaken in the Phase II study on three flocculated phosphatic clays. The three clays were selected based upon the Phase I findings to cover the full range of anticipated behavioral characteristics. Laboratory testing consisting of evaluations of index properties, settling and consolidation behavior, and shear strength properties were subsequently undertaken. The results are presented in Volume 7 titled Engineering Properties of Flocculated Phosphatic Clays. Results from the Phase II field testing program performed to refine the predictive capability for three selected sites utilizing conventional, sand-clay mix, and flocculated clay disposal methods are presented in Volume 8 titled Predictive Methodology Applied to Case Histories. A simplified empirical model for estimating the consolidation properties of phosphatic clays and sand-clay mixes is presented. Predictions from laboratory data and empirical correlations using program SLURRY are compared with field measurements. A user s manual for a refined version of program SLURRY is presented in Volume 9 titled Program SLURRY-User s Manual. The theoretical background, layout and algorithm for the model is discussed, a program listing is given, and several sample problems illustrating the capabilities of the program are presented. Evaluation of the relative merits of various phosphatic clay disposal and reclamation methods is made in Volume 10 on the basis of parametric studies illustrated via an example of a model mine. iv

7 EVALUATION OF PHOSPHATIC CLAY DISPOSAL AND RECLAMATION METHODS Volume 8: Predictive Methodology Applied to Case Histories Research Project FIPR Final Report, December 1991 A simplified and practical empirical model for estimating the consolidation properties of phosphatic clays and sand-clay mixes, based on index properties, is presented. The recommended semi-empirical relations are based on previous findings and may be used in the absence of site specific consolidation data. Finally, the predictive methodology was calibrated based on predictions made from laboratory test data and program SLURRY, and comparison of the predictions to field measurements. In order to accurately model field conditions, it is concluded that the coefficient of permeability determined from laboratory tests, and/or from empirical correlations, may have to be increased to account for the fact that the in situ coefficient of permeability is often greater than backfigured from laboratory tests. V

8 ACKNOWLEDGEMENTS The support and sponsorship provided by the Florida Institute of Phosphate Research to conduct this investigation is gratefully acknowledged. The cooperation and assistance provided by the following mining companies and their representatives during collection of clay samples and during the field investigation is greatly appreciated, namely: Agrico Chemical Company AMAX Phosphate, Inc. Beker Phosphates Brewster Phosphates CF Mining Corporation Estech, Inc. Gardinier, Inc. Hopewell Land Corporation International Minerals and Chemicals Corporation Mobil Chemical Company Occidental Chemical Company United States Steel Corporation Agrichemicals Division W.R. Grace & Company vii

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15 Section 1 INTRODUCTION 1.1 Predictive Methodology To comprehensively evaluate phosphatic clay disposal and reclamation methods, a predictive methodology was developed based on: (i) laboratory determined engineering properties of phosphatic clays, sand-clay mixes and flocculated or thickened clays; and (ii) a theoretical model, Program SLURRY, for predicting performance during filling of settling areas, post-filling quiescent consolidation, and reclamation. The objective of this investigation was to use established correlations between the consolidation properties of phosphatic clays and to verify the applicability of the predictive methodology to a range of disposal alternatives (i.e., conventional, sand-clay mix and thickened clay disposal methods). Accordingly, the predictive methodology was applied to a case history that involved a sand-clay mix disposal method, at the CF Mining- Hardee site. Field investigations were also undertaken at each of three test sites to allow evaluation of in situ conditions (i.e., clay thickness, solids contents, effective stresses, etc.). The results of the field investigations, predicted performance from laboratory properties and Program SLURRY, and comparisons of predicted performance with in situ conditions are presented herein, 1.2 Scope of Investigation The scope of investigation included performing a field investigation to document in situ conditions, and making predictions of the performance using Program SLURRY and laboratory and empirically determined engineering properties. The disposal areas selected for investigation included: USSAC-Rockland Settling Area S-6; CF Mining-Hardee Reclamation Area R-2; and the Gardinier-Fort Meade Flocculated Clay Test Pit. The three selected areas use different methods available for disposal of phosphatic clays. USSAC-Rockland Settling Area S-6 is a conventional settling area; CF Mining-Hardee Reclamation Area R-2 is a sand-clay mix disposal area; and the Gardinier-Fort Meade Test Pit is an experimental flocculated clay area. The field measurements of engineering properties of phosphatic clay within each area are presented in Section 2. The in situ clay thickness, solids contents, pore pressures, boundary drainage conditions, effective stresses, field compressibility curve, cone tip resistance and undrained shear strength are considered. Empirical relations for preliminary estimates of consolidation properties of phosphatic clays and sand-clay mixes are presented and discussed in Section 3. In Section 4, performance predictions using Program SLURRY are presented and compared with in situ conditions for the CF Mining-Hardee sand-clay mix disposal method. 1-1

16 Section 2 FIELD MEASUREMENTS OF ENGINEERING PROPERTIES OF PHOSPHATIC CLAYS FOR THREE DIFFERENT DISPOSAL METHODS 2.1 Introduction The results of field measurements of the engineering properties of phosphatic clays disposed of via conventional, sand-clay mix and flocculated clay disposal methods are presented in this section. The objective of the field investigation was to determine the thickness, in situ solids contents, effective stresses and boundary drainage conditions existing at each site to allow a comparison with predicted thickness, solids contents, and effective stresses obtained via Program SLURRY* using laboratory measured properties. A comparison between in situ and predicted conditions will allow calibration of the predictive methodology to improve the model if need be. Accordingly, the field investigation included measurements of the in situ solids content with depth for estimating total stress and clay consistency; pore pressures at the lower boundary and within the clay with depth for estimating boundary drainage conditions, excess pore pressures and, hence, effective stresses; and undrained shear strength with depth as inferred from continuous cone soundings for estimating deposit uniformity, in situ shear strength and the locations of desiccated or overconsolidated zones. The field sampling methods used for obtaining specimens for solids content determinations, and in situ measurement techniques utilized for pore pressure probes and electric piezocone soundings were previously described in Volume Site Selection Three disposal areas were selected for the field investigation. The sites selected included: USSAC-Rockland Settling Area S-6 CF Mining-Hardee Reclamation Area R-2 Gardinier-Fort Meade Flocculated Clay Test Pit The three selected areas cover different methods presently used for disposal of phosphatic clays. USSAC-Rockland Settling Area S-6 is a conventional clay settling area; CF Mining-Hardee Reclamation Area R-2 is a sand-clay mix disposal area; and the Gardinier-Fort Meade Test Pit is an experimental flocculated clay disposal area. Due to the different disposal methods utilized at the three selected areas, it is anticipated that the field investigation will allow checking and calibration of the predictive methodology for the range of disposal alternatives available to mine planners. * Refer to Volume 6 Predictive Methodology for Evaluating Disposal Methods ; and Volume 9 Program SLURRY-User s Manual. 2-1

17 2.3 USSAC-Rockland Settling Area S Site Description The USSAC-Rockland mine was selected for investigation based on laboratory data previously presented in Volumes 1 through 5 which indicated that phosphatic clays produced at the mine displayed relatively average properties. The area selected for investigation, Settling Area S-6, is a relatively mature active pond which has been in operation since March The western compartment of the area which was sampled comprises about 370 acres and, except in the southwestern portion, was largely constructed within mined lands. The crest elevation of the dam is 170 feet (MSL) and adjacent natural ground surface elevations range from 130 to 140 feet (MSL). Hence, the embankment crest is 30 to 40 feet above adjacent natural ground. Postmining topographic maps of the area indicate that the pit bottom may extend to below Elevation 120 feet (MSL), potentially resulting in up to 45 feet of stored phosphatic clay based on a maximum fluid level of 165 feet (MSL) Field Testing Program The Phase II field investigation program at Settling Area S-6 consisted of six test holes advanced to pit bottom at the locations schematically illustrated in Figure 2-1. The locations of the test holes previously performed during the Phase I field investigation in March and April, 1983 are shown. At each of the Phase II field investigation test holes, samples were obtained with depth for solids content determinations. A total of 64 samples were recovered from within the settling area. Pore pressure probes were installed at various depths at each test hole location for the measurement of in situ pore pressures and for the determination of boundary drainage conditions, A minimum of four pore pressure probes were installed at each location. A total of 32 pore pressure probes were used within the settling area. One hydraulic piezometer was also installed at each of TH-2, TH-3 and TH-5 for measurement of in situ pore pressures. Electric piezocone soundings were performed at TH-1, TH-3, and TH-5. The Phase II sampling and pore pressure probe installations were completed in September and October, The electric piezocone soundings were completed in January, During the investigation, the water surface within the settling area varied from Elevation to feet (MSL) as recorded at the spillway within the south wall. The water depths above the surface of the settled clays at the test hole locations varied from 0.5 to 3.0 feet and averaged 2.0 feet, Since the area was ponded, all sampling and in situ testing was performed from an airboat equipped with a tripod for advancing field sampling equipment and instrumentation. The settling area was generally overgrown with cattails, except at the northwest corner. The depth locations of the recovered samples and the pore pressure probes (or piezoprobes) at each test hole are schematically illustrated in Figures 2-2 through 2-7. The Atterberg limits, solids contents, total unit weights, measured and hydrostatic pore pressures, and total stress profile developed at each test location are also presented on the figures. 2-2

18 2.3.3 Plasticity Characteristics The plasticity characteristics of phosphatic clays within Settling Area S-6 determined from sixteen Atterberg limits are presented in Figure 2-8. As shown, the liquid limit varied from 180% to 240% and the plasticity index varied from 145% to 195%. The average liquid limit and plasticity index were 211% and 173%, respectively, which are representative of relatively average phosphatic clay plasticity characteristics. Further, as shown in Figure 2-9, the plasticity characteristics were relatively uniform with depth displaying no monotonic trends or apparent layering. At individual test holes, the average plasticity index varied from 159% to 183% and generally increased in an easterly direction across the pond. The average liquid limit and plasticity index, as obtained in the Phase II field investigation, are also consistent with the average liquid limits of 195% to 206% and average plasticity indices of 160% to 156% previously found during the laboratory investigation (Volume 1) and Phase I field investigation (Volume 6). Accordingly, it appears that the plasticity characteristics of the settling Area S-6 phosphatic clays are relatively uniform, and based on the combined results of the laboratory, Phase I and Phase II field investigations, the Atterberg limits can be characterized by a representative liquid limit of 209% and plasticity index of 168% Clay Thickness To determine the clay thickness, each test hole was terminated within the cast spoil underlying the settling area. The clay thickness encountered at each test hole location is summarized below: As shown, the clay thickness was found to vary from 25.5 to 39.0 feet with an average of 30.5 feet for the six locations. Three test holes previously completed during the Phase I field investigation were terminated at a depth of 28 feet without encountering pit bottom. Hence, a clay thickness in excess of 30 feet appears typical for the settling area Solids Content The measured solids content versus depth at each test hole location are presented in Figures 2-2 through 2-7 and summarized in Figure 2-9. As shown, the solids content increases from 11% to 18% near the surface to over 40% at the lower boundary with the underlying cast spoil. The overall average solids content for samples recovered during the Phase II field investigation was 24.8%±8.5%.* The data also clearly illustrate (Figure 2-9) that the solids content at test holes * Plus and minus one standard deviation. 2-3

19 of shallower depths (i.e., TH-1, TH-3 and TH-4) are consistently slightly higher than those at test holes of deeper depths (i.e., TH-5 and TH-6). A comparison of solids contents measured during the Phase I and Phase II field investigations is presented in Figure In order to allow comparison at similar locations within the settling area, the solids contents measured at TH-3 and TH-4 during the Phase I field investigation were compared with the solids. contents measured at TH-4 and TH-6, respectively, during the Phase II field investigation (Figure 2-1). As shown, there are no significant differences between solids contents at the two locations despite the 17-month period between sampling dates, Also note that the solids contents are consistently lower with depth at the center location (i.e., Phase I TH-4 and Phase II TH-6) than at the location near the pond edge (i.e., Phase I TH-3 and Phase II TH-4) Pore Pressures and Boundary Drainage Conditions The measured pore pressures, u,, versus depth and estimated total vertical stress, uv,, versus depth from total unit weight determinations are presented for each test hole in Figures 2-2 through 2-7 and summarized in Figure 2-9. The hydrostatic pore pressure, u, with depth is included for comparison with measured values. As shown and expected, the measured pore pressures are slightly greater than hydrostatic pore pressures within the upper 64% to 90% of the clay thickness. At the lower boundary with the underlying cast spoil, the measured pore pressures are consistently lower than hydrostatic pore pressures by equivalent hydraulic heads of 2 to 19 feet of water. Hence, the underlying cast spoil is serving as a drainage boundary and downward seepage through the clay is occurring. Further, the measured pore pressures within the cast spoil were found to be influenced, as expected, by adjacent surface elevations and land use. At TH-1 and TH-5, the piezometric water elevations within the cast spoil were approximately 160 feet (MSL) and 159 feet (MSL) or 2 and 3 feet, respectively, below hydrostatic pore pressures, These relatively high boundary pore pressures may be explained by the presence of another USSAC-Rockland settling area immediately to the north of TH-1 and the Gardinier settling area immediately adjacent to the east wall of Settling Area S-6. Due to the presence of these adjacent settling areas, there is no nearby relief point for the underlying cast spoil at these locations and, hence, the pore pressures approach hydrostatic pond water level conditions, At TH-2, TH-3, TH-4 and TH-6 the piezometric water elevations within the cast spoil generally progressively decrease from the pond center towards the pond edge from approximately 155 feet (MSL) to 143 feet (MSL) or 7 to 19 feet below the hydrostatic head. Around the south and east walls of Settling Area S-6 the ground surface elevations range from 130 feet (MSL) to 140 feet (MSL) providing a relief point for the cast spoil foundation which is reflected in the lower measured boundary pore pressures Effective Stresses and Comparison of Laboratory and Field Compressibility Curves Effective vertical stresses, a,,, inferred from estimated total stresses o;,, and measured pore pressures, u,, are tabulated on Figures 2-2 through 2-7 for the location of each pore pressure measurement (where TV, = Q~;,-u,) and are graphically summarized with depth in Figure 2-9. As shown, in the upper 15 feet the effective vertical stresses are very low, typically in the range of 2-4

20 0 to 50 psf. Near the lower drainage boundary the effective vertical stresses increase to about 750 to 1250 psf. A comparison of the void ratio versus effective vertical consolidation stress relationships determined by in situ measurements of solids content* and pore pressure and by laboratory measurements from slurry consolidation tests is presented in Figure 2-11, The void ratio versus effective vertical consolidation stress laboratory compressibility curve previously established for the USSAC-Rockland phosphatic clay (Volume 4, Section ) is included on the figure, The void ratio and effective stress inferred at each pore pressure measurement point used to develop Figure 2-11 are tabulated on Figures 2-2 through 2-7. As shown, the agreement between field and laboratory data is excellent. Accordingly, the use of laboratory determined compressibility properties to predict the in situ compressibility behavior of the USSAC-Rockland phosphatic clay appears justified Cone Tip Resistance and SHANSEP Undrained Shear Strenqth Theoretical Background Estimates of the undrained shear strength, s,, of a clay from the cone tip resistance, q,, have been based on bearing capacity and cavity expansion theories. Three widely used theories include: where: N, is the bearing capacity factor for deep circular foundations; N*, and N, are bearing capacity type factors that depend on the rigidity index, I, = G/s, = E,/3s,; o, is the total vertical stress; 2 the effective vertical stress; o,, the total horizontal stress; K, the coefficient of lateral earth pressure at rest; G the undrained shear modulus; and E, Young s undrained secant modulus. Since the value of the various bearing capacity factors and required stresses can vary significantly, and since it is preferable not to require detailed information about the soil deposit (such as stress history) for interpretation of cone tip resistance, a simpler equation of the form s, = q,/c, where C is a constant, was originally proposed and utilized during the Phase I field investigation. The derivation of the constant C in this procedure, however, required the assumption of 100% consolidation at the existing in situ effective stresses which is typical for natural clay deposits but is clearly a simplifying assumption for the phosphatic clays, Instead, it appears that a more appropriate approach for the phosphatic clays is to reduce the bearing capacity and refined cavity expansion theories into the form: * Solids content, S, and void ratio, e, are related by the expression: e = ~(1 -S)/S, where p is the specific gravity of clay solids. 2-5

21 where yw is the unit weight of water and Z is the depth at which the cone tip resistance is obtained. Based on the bearing capacity and refined cavity expansion theories, values of C of 13 to 20 are expected. This form of equation relating q, and s, to a constant does not require the simplifying assumption of 100% consolidation under the existing in situ stresses and hence, is more applicable to phosphatic clays undergoing consolidation. Given the variability of the two theories and potential sensitivity of C to solids content, the selection of a unique C value based on theoretical considerations is not recommended. Hence, field comparisons of cone tip resistance and in situ undrained shear strength are necessary to establish empirical C values Comparison of Cone Tip Resistance and SHANSEP Undrained Shear Strength The cone tip resistance, q,, and pore pressure during penetration, u,, measured at test location 1 are presented in Figure As shown, both the cone tip resistance and pore pressure during shear increase uniformly with depth as expected for a uniform deposit of normally consolidated clay. The effective vertical stress, a,,, inferred from measured solids contents and the laboratory compressibility relationship, and the SHANSEP* undrained shear strength s,(dss)* obtained from the inferred effective vertical stresses are plotted versus depth on Figure The (q,- y,z)/s,(dss) ratio is also presented. As shown, the (q, - y,z)ls,(dss) ratio varies from 34 at a depth of 5 feet to 19 at a depth of 25 feet. Accordingly, the (q, - y,z)ls,(dss) ratio is not a constant, but apparently increases with decreasing solids content. A similar finding was noted during the Phase I field investigation. Hence, it appears piezocone soundings can be advantageously used to estimate in situ undrained shear strengths and effective stresses for normally consolidated clays. 2.4 Gardinier-Fort Meade Flocculated Clay Test Pit Site Description The Gardinier-Fort Meade test pit was selected for investigation because it is one of the few existing controlled flocculated clay disposal areas. The disposal area was located between two parallel cast spoil piles and comprised an area of about 4.6 acres. The bottom of the disposal area was initially filled with a layer of sand tailings to form an underdrain. A series of perforated pipes and a sump were used to remove water from the underdrain via a pumping station. The storage depth above the sand tailings underdrain reportedly varied from about 16 to 28 feet. Flocculated clay was initially deposited within the test pit in March, 1982 and continued for a 19- month period through September, During this time, 42,026 tons of flocculated clay were deposited within the test pit. The flocculent loading rate was proprietary, but was reportedly varied during the filling period. A total of 29 different flocculents were reportedly used. (4) * Refer to Volume 5 Shear Strength Characteristics of Phosphatic Clays. 2-6

22 2.4.2 Field Testing Program The field investigation program consisted of six test holes advanced to pit bottom at the locations schematically illustrated in Figure At each test hole, samples were obtained with depth for solids content determinations. A total of 60 samples were recovered from within the test pit, Pore pressure probes were installed at various depths at five of the test locations for the measurement of in situ pore pressures and for the determination of boundary pore pressures, A total of 18 pore pressure probes were used within the settling area. Four shallow observation wells were also installed for determining the depth to the water table. Electric piezocone soundings were performed at TH-1, TH-3 and TH-5. The sampling and pore pressure probe installations were completed during the period of May through August, The electric piezocone soundings were completed in January At the time of the investigation, the test pit was not ponded except for a relatively small portion along the south end. Most of the test pit also had a desiccated surface crust and was overgrown with cattails. Due to the presence of the desiccated crust which was sufficiently strong to support light traffic, access to the test locations was achieved by walking along sections of conveyor belt and/or plywood boards. The depth locations of the recovered soil samples and pore pressure probes (or piezoprobes) at each test hole are schematically illustrated in Figures 2-14 through The Atterberg limits, solids contents, total unit weights, measured and hydrostatic pore pressures and total stress profile developed at each test location are also presented on these figures Plasticity Characteristics The plasticity characteristics of flocculated phosphatic clays within the test pit determined from 22 Atterberg limits are presented in Figure As shown, the liquid limit varied from 148% to 200% and the plasticity index varied from 115% to 170%. The average liquid limit and plasticity index were 170% and 139%, respectively, which are representative of relatively average phosphatic clay plasticity characteristics. Further, as shown in Figure 2-21, the plasticity characteristics were relatively uniform with depth displaying no monotonic trends or apparent layering. At individual test holes, the average plasticity index varied from 129% to 156% and generally decreased in a southerly direction across the pond except at TH-6. The Atterberg limits of the parent phosphatic clay used to fill the test pit were determined on a sample of unflocced clay obtained from Settling Area E. The liquid limit and plasticity index of the parent clay were 145% and 112%, respectively. Accordingly, it appears that flocculation may have increased the liquid limit and plasticity index on the average by 25%. Although flocculation may be expected to increase the liquid limit ad plasticity index, part of the increase may also be potentially due to differences between the actual phosphatic clays used in the test pit and the presumed parent clays obtained from Settling Area E Clay Thickness To determine the clay thickness, each test hole was terminated within the sand tailings underdrain installed at the bottom of the test pit. The clay thickness encountered at each test hole location on the specified sampling dates is summarized below: 2-7

23 As shown, the clay thickness varied from 8.5 feet to 22.5 feet. At TH-2 and TH-4 within the eastern portion of the test pit, the pit bottom was encountered at depths of only 8.5 to 9.0 feet, At TH-1, TH-3, TH-5 and TH-6 within the central and western portions of the test pit, the pit bottom was encountered at depths of 17.5 feet to 22.5 feet with an average of about 19 feet. Hence, a clay thickness of 19 feet appears typical for most of the test pit, except for a strip along the eastern edge of the pit where the depth is only about 9 feet Solids Content The measured solids contents versus depth at each test hole location are presented in Figures 2-14 through 2-19 and summarized in Figure As shown, the solids content increases from the range of 24% to 33% near the surface to over 40% at the lower boundary with the sand tailings underdrain. The overall average solids content for samples recovered from the deeper test holes (i.e., TH-1, TH-3, TH-5 and TH-6) was 31.1%±8.4%.* If the solids contents determined at the two shallow test holes are included, the average solids content for the test pit increases to 32.4%±8.9%. * The measured solids contents at the shallow locations (i.e., TH-2 and TH-4) are consistently higher than those determined at the deeper locations and display an average solids content of about 36% compared to about 31% Pore Pressures and Boundary Drainage Conditions The measured pore pressures, u,, with depth and estimated total vertical stress, o;,, with depth from total unit weight determinations for each test hole where pore pressure probes were installed are presented in Figures 2-14 through 2-19 and are summarized in Figure 2-21, The hydrostatic pore pressure, u, with depth is included for comparison with measured values. As shown and expected, the measured pore pressures are slightly greater than hydrostatic pore pressures within most of the clay, except for the lower several feet adjacent to the boundary with the sand tailings underdrain. At the lower boundary with the underdrain, the measured pore pressures are at or slightly lower than hydrostatic pore pressure. The boundary pore pressures, however, will vary depending on whether the underdrain pump is operating: they will be lower when the pump is on and eventually rise to about hydrostatic pressure when the pump is off, Hence, the underdrain is serving as a drainage boundary and downward seepage through the clay is occurring. * Plus and minus one standard deviation. 2-8

24 2.4.7 Effective Stresses and Comparison of Laboratory and Field Compressibility Curves Effective vertical stresses, TV,, inferred from estimated total stresses, o,,, and measured pore pressures, u,, for the location of each pore pressure measurement (where a,, = a;, - u,) are tabulated in Figures 2-14 through 2-19 and are graphically summarized with depth in Figure As shown, in the upper 10 to 12 feet, the effective vertical stresses are very low, typically in the range of 0 to 50 psf. Near the lower drainage boundary, the effective vertical stresses increase to the range of 250 to 450 psf. The field void ratio versus effective vertical consolidation stress relationship determined via in situ measurements of solids content and pore pressure is presented in Figure The void ratio and effective stress inferred at each pore pressure measurement point used to develop Figure 2-22 are tabulated on Figures 2-14 through The results from a laboratory constant rate of strain consolidation test (CRSC test) performed on an undisturbed sample of the flocculated clay are also included. As shown, the field and laboratory data agree well, and yield an estimated field compressibility curve using a linear log void ratio versus log effective stress relationship of e = 2J(jg;jvc -o.226 (TV, in units of kg/cm*) Cone Tip Resistance and SHANSEP Undrained Shear Strength The cone tip resistance and pore pressure during penetration measured at test location 3 are presented in Figure As shown, the cone tip resistance, q,, clearly indicates a desiccated crust within the upper few feet of the deposit as evidenced by increasing q, values with decreasing depth. The general shape of the q, line with depth and the relatively low measured pore pressures during penetration also suggest that the clay may be slightly overconsolidated due to the effects of either a water table lowering or downward seepage gradient when the underdrain is pumped. The effective vertical stress, a,,, inferred from both measured solids contents and the field compressibility relationship (Figure 2-22), and from the estimated in situ total stresses and measured pore pressures are plotted versus depth on Figure The SHANSEP undrained shear strengths, s,(dss), for normally consolidated clays obtained from the effective vertical stresses and the (q, - y,z)/s,(dss) ratio are also presented. As shown, the (q, - yj)/s,(dss) ratio is very high and varies from 90 to over 110. These apparently high values are attributed to the effect of overconsolidation which implies that the shear strength of the clay is actually higher than indicated in Figure CF Mining-Hardee Reclamation Area R Site Description The CF Mining-Hardee Reclamation Area R-2 was selected for investigation because it was one of the first areas where experimental sand-clay mix disposal was being evaluated on a production scale, Further, laboratory test data presented in Volumes 1 through 5 indicated that phosphatic clay at the site (from Settling Area N-1) was characterized by a relatively low plasticity, high settled solids content, and low compressibility. Phosphatic clay from the benefication plant enters Initial Settling Area N-1 where it is allowed to thicken through natural settling and consolidation to solids contents of 12 to 18%. A dredge is then used to excavate the thickened clay and pump it to a mixing station where it is mixed with sand tailings. The sand tailings are 2-9

25 pumped from the flotation plant to the mixing tank where they are dewatered by cyclone prior to mixing with the clay. After the dewatered sand tailings and thickened clays are mixed they are pumped to Reclamation Area R-2. Approximately 1,010,000 tons of sand tailings and 534,000 tons of clay at an average sand-clay ratio of 1.89±0.32:1.00 and initial total solids content of 36.2±5.8% were deposited within the reclamation area during a filling period from October, 1980 through January, Reclamation Area R-2 covers approximately 110 acres of mined land and has a total storage capacity of 3,600 acre-feet. At the maximum design fluid level, the sand-clay mix would average 35 feet in thickness and would approach 50 feet in thickness in deeper mined-out pits Field Testing Program The Phase II field investigation program at Reclamation Area R-2 consisted of four test holes completed to pit bottom at the locations schematically illustrated in Figure 2-24.* At each of the Phase II field investigation test holes, samples were obtained continuously with depth for determination of sand-clay ratio and solids content. A total of 109 samples were recovered within the area. Four pore pressure probes were installed at different depths at each test location for the measurement of in situ pore pressures and for determination of boundary drainage conditions. Electric piezocone soundings were also performed at each test location. The field sampling and pore pressure probe installations were completed in July, The electric piezocone soundings were performed in late December, 1984 to early January, At the time of the investigation, the area was not ponded, had a desiccated surface crust, and was overgrown with cattails. Due to the desiccated crust which was sufficiently strong to support light traffic, access to the test locations was achieved by either walking directly on the crust or on plywood boards placed on the crust. The depth locations of the recovered samples and pore pressure probes (or piezoprobes) at each test hole are schematically illustrated in Figures 2-25 through The sand-clay ratio, solids contents, total unit weights, measured and hydrostatic pore pressures, and total stress profile developed at each test hole location are presented on these figures. The electric piezocone tip resistance, q,, and pore pressure during penetration u,, are also included Sand-Clay Mix Thickness To determine the sand-clay mix thickness each test hole was terminated within cast spoil underlying the settling area. The sand-clay mix thickness encountered at each test hole is summarized below: * The field investigation of this area was performed jointly for Florida Institute of Phosphate Research Contracts Evaluation of Phosphatic Clay Disposal and Reclamation Methods and Field Evaluation of Sand-Clay Mix Reclamation. 2-10

26 * The sampled hole and piezocone sounding were performed within the same general location but differences in thickness are expected due to variations in the pit bottom cast spoil. As shown, the thickness of the sand-clay mix was found to vary from 22.0 feet to 28.0 feet with an overall average for the sampled holes and piezocone soundings locations of about 26.2±2.1 feet Solids Contents and Sand-Clay Ratio The solids contents measured with depth at each test location are presented in Figures 2-25 through 2-28 and summarized in Figure The total solids content, S t, and sand-clay ratio, SCR, of each sample were determined and the clay solids content, S,, and effective clay solids content, S,, were calculated from the expressions: s, = l/[((l +SCR)/S,) - SCR] S, = li[(scr/ p,)+(lis,)] where the sand-clay ratio is defined as the ratio by dry weight of material retained on the United States Standard No. 140 sieve to the material passing the United States Standard No. 140 sieve, and ps is the specific gravity of the sand tailings. As shown in Figure 2-29, there is considerable variability in the total solids content and sand-clay ratio with depth among the four test locations. The clay solids content, however, generally displays little variability. For determination of the representative trends with depth shown in Figure 2-29, representative trends with depth were first selected for the measured total solids contents and sand-clay ratios. The representative trends with depth for the clay solids content and effective clay solids content were then calculated from Equations 5 and 6. Although there is considerable variability in the solids contents and sand-clay ratios, the following trends are apparent in the data: Although a desiccated crust existed at each test location as evidenced by visual observations and piezocone soundings, measurements do not clearly indicate an increase in solids content with decreasing depth near the surface as would be expected within a mature crust. The total solids content, clay solids content, and sand-clay ratio appear to increase slightly with depth, whereas the effective clay solids content appears to remain relatively uniform 2-11

27 with depth. Representative average values of each parameter determined for 5-foot depth intervals are summarized below: Representative average values (± one standard deviation) for the entire disposal area for total solids content (SJ, sand-clay ratio (SCR), clay solids content (SJ, and effective clay solids content (3,) based on measurements at the four test locations are presented below: Representative Value A trend of decreasing total solids content, sand-clay ratio, clay solids content and effective clay solids content from TH-5 towards TH-B (Figure 2-24) is apparent when the average values are compared at each test location as shown below: Pore Pressures and Boundary Drainage Conditions The measured pore pressures, u,, with depth and estimated total vertical stress, o;,, with depth from total unit weight determinations for each test hole are presented in Figures 2-25 through 2-28 and are summarized in Figure The hydrostatic pore pressure, u, with depth is included for comparison with the measured values. As shown and expected, the measured pore pressures are slightly greater than hydrostatic pore pressures within the upper 90% of the sandclay mix thickness. At the lower boundary with the underlying cast spoil, the measured pore pressures are consistently lower than hydrostatic pore pressures by 1.9 to 10.6 feet of water head with an 2-12

28 average of 6.8 feet of water. Hence, the underlying cast spoil is serving as a drainage boundary and downward seepage through the sand-clay mix is occurring Effective Stresses and Comparison of Laboratory and Field Compressibility Curves Effective vertical stresses, a,,, inferred from estimated total stresses, o,,, and measured pore pressures, u,, for the location of each pore pressure measurement (where ZVO = o,,-u,) are tabulated on Figures 2-25 through 2-28 and are graphically summarized with depth in Figure As shown, the effective vertical stresses vary considerably among the four test locations. At a depth of 15 feet, the effective vertical stress varies from a minimum of 100 psf at TH-B to a maximum of 500 psf at TH-5. Near the lower drainage boundary the effective vertical stresses increase to the range of 1300 to 1800 psf. Comparisons of the total void ratio and clay void ratio versus effective vertical consolidation stress determined from in situ measurements of solids content and pore pressure and from laboratory slurry consolidation test measurements are presented in Figures 2-31 and Previously established void ratio versus effective stress data from the Phase I field investigation are shown. The total void ratio, et, and clay void ratio, e,, are defined, respectively, in terms of solids contents by the equations: e t = p(1 - S t )/S t e, = e,u +WCRIPJ) (8) where: p is the effective specific gravity of the solids in the sand-clay mix; pc is the specific gravity of clay; and ps is the specific gravity of sand. Specific gravities of 2.81 for the clay and 2.70 for the sand tailings are characteristic for the CF Mining-Hardee site. The effective specific gravity of the solids in the sand-clay mix is given by the following relation: P = ps P,U +SCW@SR+p,) (9) Moreover, if the volume of sand is accounted for as an equivalent displaced water phase, then the effective clay void ratio, e,, is related to the clay void ratio, e,, by the expression: -EC = e, + C~,/P,FCR As shown, the field compressibility curves generally yield higher total and clay void ratios at a given effective stress than predicted from the laboratory compressibility curve for sand-clay mix with an overall average sand-clay ratio of 1.3:1.0 (i.e., the average sand-clay ratio for the area). Although it appears that the in situ compressibility of the sand-clay mix is slightly different than predicted from laboratory tests, part of this discrepancy is attributed to differences in sand-clay ratio and differences in plasticity of the in situ clay and the clay used in laboratory evaluations Cone Tip Resistance and SHANSEP Undrained Shear Strength The cone tip resistance and pore pressure during penetration are presented on Figures 2-25 through As shown, the cone tip resistance indicates a surface desiccated crust within the upper 1 to 2 feet of the sand-clay mix. Below the surface crust, the sand-clay mix generally 2-13

29 appears normally consolidated except for the presence of some additional overconsolidated zones from previous surface exposures during filling (i.e., TH-A from 13 to 15 feet, TH-B from 9.5 to 11.5 feet and TH-4 from 16.5 to 18 feet). Further, at TH-5 the presence of relatively sandy layers between depths of 2.5 to 3.5 feet and 9.0 to 14.0 feet are indicated by cone tip resistances in excess of 5.0 kg/cm 2. The (q, - y,z)/s,(dss) ratio was determined for each point at which the in situ pore pressures and, hence, effective vertical stresses were measured. Using a normalized undrained shear strength ratio of as recommended in Volume 5 for a sand-clay mix with a sand-clay ratio of 2.0:1.0, the (q, - y,z)ls,(dss) ratios were found to vary from 12 to 29 with an average of Summary and Practical Implications Based on the Phase II field measurement results and comparison to laboratory test data, the following relevant implications on the predictive capability can be made: Underlying cast spoil material generally serves as a lower drainage boundary during consolidation of phosphatic clays and sand-clay mixes. The use of laboratory determined compressibility properties to predict the in situ compressibility of phosphatic clays is justified based on the good agreement between field and laboratory void ratios and effective stresses. Piezocone soundings can be advantageously used to estimate the in situ undrained shear strength and, hence, effective stresses of normally consolidated phosphatic clays and sand-clay mixes. 2-14

30 FIGWE 2-1

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37 Figure 2-8

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39 Figure 2-10

40 Figure 2-11

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55 92-z 3klrlDIzl

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62 Section 3 SIMPLIFIED EMPIRICAL MODEL FOR CONSOLIDATION PROPERTIES OF PHOSPHATIC CLAYS AND SAND-CLAY MIXES 3.1 Introduction The consolidation properties of phosphatic clays and sand-clay mixes are of interest to mine planners since they dictate the relationship of height of slurry and/or solids content versus time during filling of a settling area at a given loading rate. For predicting the storage capacity and life of a settling area, selection of representative consolidation properties often constitutes the most relevant geotechnical engineering task. Accordingly, as part of research project FIPR Evaluation of Phosphatic Clay Disposal and Reclamation Methods performed for the Florida Institute of Phosphate Research, twelve phosphatic clays, sampled from various mine sites, were investigated via conventional, constant rate of strain and slurry consolidation tests to determine the range of consolidation behavior of Florida phosphatic clays. Consolidation tests were also performed on sand-clay mixes prepared from three selected phosphatic clays with widely varying properties, as outlined in Volume 4 titled Consolidation Behavior of Phosphatic Clays by Wissa et al. (1983). The mine sites were selected to provide a range of geographic locations and mining concerns. Based on this information, the consolidation properties of phosphatic clays and empirical correlations between index properties and consolidation parameters of phosphatic clay and sandclay mixes were developed and presented in Volumes 4 and 6, Wissa et al. (1983). The consolidation properties of interest include: Once compressibility and rate of consolidation relationships (e - a,,, and k - e) are well defined for a wide range of phosphatic clays and sand-clay mixes, correlations can be established with index properties, settling characteristics, and/or mineralogic composition. The range of * Void ratio, e, is the volume of water in a saturated soil element divided by the volume of solids; the solids content, S, is the weight of solids divided by the total weight of solids and water; these two terms are interrelated by the relationship S = p/b+e) where p is the specific gravity of solids. The two terms can be used interchangeably. 3-1

63 properties established in Volume 4, Wissa et al. (1983), and empirical correlations with index properties presented therein and in Volume 6, Wissa et al. (1983), are, therefore, invaluable to mine planners and can be used to comprehensively evaluate phosphatic clay disposal and reclamation methods. The following sections summarize and extend previous findings and present a simplified and practical empirical model for obtaining an estimate of the consolidation properties of phosphatic clays and sand-clay mixes based on index properties. These semi-empirical correlations do provide a reasonable fit to the consolidation behavior of phosphatic clays and sand-clay mixes, The simplified relationships are, therefore, of practical significance and can be advantageously used in evaluating phosphatic clay disposal and reclamation methods. 3.2 Phosphatic Clays Sedimentation of Phosphatic Clays Volume 3, Wissa et al. (1983), presented a comprehensive evaluation of the sedimentation behavior of phosphatic clays, The final solids content (and/or final void ratio*) of phosphatic clays after deposition and gravity settling within an impoundment is of interest in modeling the consolidation behavior of phosphatic clays because it corresponds to the initial solids content (or initial void ratio) used in consolidation predictions. Typically, the initial void ratio (or solids content) in consolidation analyses is selected as the 30-day settled void ratio (or solids content) determined from laboratory settling tests starting from the initial solids content of the slurry (generally 3 to 4%). Figures 3-1 and 3-2 illustrate correlations found between the liquid limit, LL, of the clay and the final settled solids content and void ratio after 30 days of settling, respectively, starting from an initial slurry solids content of 3%. As shown in Figure 3-1, there is some variability in the data for the 19 phosphatic clays evaluated, but a reasonable correlation between settled solids content and liquid limit is found, which can be expressed as follows: S i = 1/( x10-4 LL) (1) where Si is the initial settled solids content for use in consolidation predictions** and LL is the liquid limit, both expressed as a percent. The correlation coefficient, r, of indicates that 61% of the variability in the settled solids content data can be explained by variations in the liquid limit. The following somewhat improved correlation was found between the liquid limit and settled void ratio, as illustrated in Figure 3-2: ei = LL (2) * Solids content, S, and void ratio, e, are related by the expression: e = ~(1 -S)/S, where p is the specific gravity of the solids. ** S i in Equation 1 is essentially identical to S, in Figure 3-1, (Note that the linear regression previously presented in Volumes 3 and 6 incorporated a typographical error in the exponent of Equation 1, which has been corrected in this volume.) 3-2

64 where e i is the initial settled void ratio for use in consolidation analyses. This expression exhibits a correlation coefficient of which is only slightly improved from the value of found for the liquid limit versus settled solids content correlation. Either regression equation, therefore, is equally satisfactory for predicting the settling behavior of phosphatic clays from the liquid limit, in the absence of site specific settling test data Void Ratio Versus Effective Stress Relationship As discussed in Volume 4 and summarized in Volume 6, Wissa et al. (1983), the compressibility of phosphatic clays can be modeled by an equation of the form: e=azpyc (3) where a and /I are material property constants. The coefficient a represents the intercept of the log-log linear relationship at an effective consolidation stress, a,,, of unity and, hence, a depends on the units of stress used. The coefficient p is the slope of the log-log linear relationship characterizing compressibility. There is no theoretical basis for selection of a log-log linear relation, only empirical evidence that such a relation statistically yields a best fit curve for a series of data. The above form of compressibility equation is generally adopted due to its mathematical simplicity for use in consolidation models. As detailed in Volume 4, Wissa et al. (1983), the compressibility parameters a and /3 are correlated to the plasticity index of the clay, in accordance with the following relationships: a = PI (4) /3 = (12.868/Pl) (5) As indicated by Equation 4, the a parameter increases linearly with the plasticity index, PI. PI in Equation 4 is expressed as a percentage, and a corresponds to the value of the corresponding coefficient in Equation 3, when Oyc is expressed in units of kg/cm. Equation 4 exhibited a correlation coefficient of indicating that the linear regression provides a good statistical fit to the data, and that the variability in the plasticity index explains 95.1% of the variability in the a parameter for phosphatic clays. Moreover, as indicated by Equation 5, the 16 parameter increases in absolute magnitude with increasing plasticity, but the change is much more subtle. A hyperbolic relationship yielded the best statistical fit to the P-PI correlation. There is slightly more variability in the /? parameter than in the a parameter at a given plasticity index as indicated by a lower correlation coefficient of compared to for the a-pi relationship. Note, however, that for the range of plasticity indices of phosphatic clays (PI = 100% to 210%), the p parameter varies within a relatively narrow range of to In the absence of site specific consolidation data, Equations 3, 4 and 5 are recommended for predicting the compressibility relationship of phosphatic clays based on the plasticity index of the clay. A somewhat less accurate yet reasonable preliminary estimate of the a and /3 compressibility parameters may also be obtained from empirical correlations between the liquidity index, LI, and effective stress, Zvc. Carrier et al. (1981) and Volume 4, Wissa et al. (1983), present empirical procedures for preliminary estimates of the void ratio versus effective stress relationship based 3-3

65 on the plasticity index of phosphatic clays. The relationship between void ratio and effective stress is based on an empirical correlation between liquidity index, LI*, and effective stress as shown in Figure 3-3. The correlation includes data for a wide range of phosphatic clays. A least squares log-log linear regression presented in Volume 4, Wissa et al. (1983), was found to yield the best fit curve to the liquidity index versus effective stress data, with a correlation coefficient, r, of 0.988; the equation is in the form: LI = yc-o.338 (6) where ;;;, is expressed in units of kg/cm 2 and LI in decimal units. As shown in Figure 3-4, the predicted liquidity index from Equation 1 is within 10% of measured values for 53% of the data and within 20% of measured values for 84% of the data. Hence, once the specific gravity and Atterberg limits of a phosphatic clay are known, Equation 6 can be used to estimate the liquidity index within ±20%, thus providing a reasonable preliminary estimate for the void ratio versus effective stress compressibility curve. The liquidity index LI, can also be expressed as: where: LI = (WC-PL)/PI LI = Liquidity index in decimal units WC = Water content in percent PL = Plastic limit in percent PI = Plasticity index in percent For a fully saturated soil element, the above equation can be expressed in terms of void ratio, e, as: e = p(li 9 PI + PL)/100 (8) where p is the specific gravity of the solids. expression: Combining Equations 6 and 8 yields the following e = p(0.459 Zvc-o.338. PI + PL)/100 (9) Preliminary estimates of the coefficients CL and /3 (in Equation 3) can be determined from the empirical correlation between effective stress and Atterberg limits (Equation 9) once the plasticity of the clay is established. This may be achieved by calculating a set of void ratios, e, using Equation 9, for corresponding values of effective stress, &,, and then performing least square linear regression analyses between log e and log O,,, For each of twelve samples of phosphatic clays with varying plasticity indices ranging from low to high, which cover the anticipated range in the plasticity of phosphatic clays, a set of void ratios were generated via Equation 9 for values of effective stresses, O;,, of 0.001, 0.01, 0.1, 1 and 10 kg/cm 2, respectively, Least square log-log * Liquidity index is defined as: LI = (e-e,)/(e,-e,), where e = void ratio, ep = void ratio at plastic limit, and e, = void ratio at liquid limit; or LI = 100 (e-e,)&)pi, where PI = plasticity index in percent, and p = specific gravity of the solids. 3-4

66 linear regression analyses between e and a,, were than performed for each clay to determine the a and,!3 compressibility parameters, as summarized in Table 3-1. Finally, values of a and /I were plotted versus the corresponding PI, as shown in Figure 3-5, and least square regression analyses were used once again to establish the relationship between a-pi and B-PI. As a result, the following empirical correlations with plasticity index were developed: a = PI (10) /3 = /Pl (11) where PI is expressed as a percentage, and the parameters a and p are the coefficients of Equation 3 with a,, expressed in units of kg/cm. Equations 10 and 11 were found to yield correlation coefficients, r of 0.98 and 0.90, respectively.* It is evident from Equation 11 and Figure 3-5 that the parameter p is confined to a relatively small range of variation with respect to PI, such that p = -0.26±0.02. Equations 10 and 11 provide a good fit to the data, as shown in Figure 3-5, if one assumes the validity of a unique relationship between liquidity index and effective stress. These empirical correlations, however, do not account for the variability depicted in Figures 3-3 and 3-4. Equations 4 and 5, on the other hand, yield an excellent fit to actual data from slurry consolidation tests, as illustrated in Figure 3-6. Hence, use of Equations 4 and 5 will provide a more reliable estimate of the compressibility parameters of phosphatic clays. Nevertheless, as illustrated in Figure 3-6, the differences between the two sets of empirical correlations are rather subtle, particularly at plasticity index values up to 180%. Figure 3-7 compares the a-pi and P-PI empirical correlations with parameters determined from settling tests and conventional consolidation test data, a rather simplified test procedure suggested in Volume 4, Wissa et al. (1983), for use in the absence of slurry consolidation test data. There is no doubt that strong correlations between the plasticity index of phosphatic clays and the compressibility parameters a and /3 exist, These empirical correlations can be advantageously used in the absence of site specific test data Void Ratio Versus Coefficient of Permeability Relationship A log void ratio, e, versus log coefficient of permeability, k, relationship was found to adequately model the behavior of phosphatic clays during consolidation (Carrier et al., 1981; Volumes 4 and 6 of Wissa et al., 1983). This relationship often statistically yields the best fit curve to a series of data, and is characterized by only two coefficients, in accordance with the following equation: k = ye (12) where y and 6 are parameters that one hopes to correlate with the plasticity of the clay. * Equations 10 and 11 were obtained by regressing 10 of the 12 data points in Table 3-1, i.e., excluding data for the Estech-Watson and Hopewell-Hillsborough clays. By including all twelve data points in the regression analyses, the following empirical correlations with plasticity index would result: a = PI; and p = /PI. The latter relationships yield somewhat lower correlation coefficients, r, of 0.96 and 0.70, respectively. 3-5

67 As detailed in Volume 4, Wissa et al. (1983), the y and 6 parameters for phosphatic clays generally vary within the following ranges: y = 2x1 o-l0 to 2x1 o-g (13) 6=3to4 (14) Note that the y parameter depends on the units adopted for the coefficient of permeability, k, in Equation 12, and that the range indicated by Equation 13 corresponds to values of k expressed in units of cm/sec. Unfortunately, as noted in Volume 4, Wissa et al. (1983), no correlation is evident between the y and 6 parameters and the plasticity index of phosphatic clays. This underscores the need to perform tests on site-specific phosphatic clays to accurately determine the y and 6 parameters for use in consolidation analyses. In an attempt to correlate the y and 6 parameters to plasticity index, PI, for use in preliminary predictions in the absence of site-specific data, one can make use of empirical correlations between liquidity index, LI, and coefficient of permeability, k. This methodology is based on normalizing the void ratio with respect to the plasticity of the clay, as described in conjunction with compressibility in Section The results obtained by this procedure were presented in Volume 4, Wissa et al. (1983), and are illustrated in Figure 3-8. A least squares log-log linear regression was found to yield the best fit curve to the liquidity index versus coefficient of permeability data, with a correlation coefficient of 0.962; the equation is in the form: LI = k,344 (15) where k is expressed in units of cm/sec and LI in decimal units. Equation 15 yields predicted liquidity indices within 10% of measured values for 26% of the data, within 25% of measured values for 66% of the data, and within 50% for 94% of the data. The correlation is not as reliable as that obtained in the preceding section between liquidity index and effective stress, but can be used to estimate the liquidity index within about ±25% once the specific gravity and Atterberg limits of a phosphatic clay are known. Following the same procedure as outlined in Section 3.2.2, Equation 15 may be expressed as follows: e = p[l 74,32(Pl)k PL]/l00 (16) where is the specific gravity of the solids; PI and PL are the plasticity index and plastic limit of the clay, respectively, both expressed as a percentage; and k is the coefficient of permeability expressed in units of cm/sec. Preliminary estimates of the coefficients y and 6 (in Equation 12) can be determined from Equation 16 by calculating a set of void ratios, e, for corresponding values of coefficient of permeability, k, for a given clay characterized by specific Atterberg limits, and then performing least square linear regression analyses between log e and log k. For each of twelve samples of phosphatic clays with varying plasticity indices ranging from low to high, which cover the anticipated range in the plasticity of phosphatic clays, a set of void ratios were generated via Equation 16 for values of the coefficient of permeability, k, of 10-5, 1 Oe6, IO-, lo-* and 10. cm/sec, respectively. Least square log-log linear regression analyses between e and k were then 3-6

68 performed for each clay to determine the y and 6 parameters, as summarized in Table 3-2. Finally, values of y and 6 were plotted versus the corresponding PI, as shown in Figure 3-9, and least square regression analyses were used once again to establish the relationship between y-pi and &PI. As a result, the following empirical correlations with plasticity index were developed: y = (40 Pl) (17) 6 = 4.OkO.25 (18) Note that the 6 parameter is not correlated to plasticity, but varies within a relatively narrow range. Equation 17, on the other hand, exhibits a correlation coefficient, r, of about 0.70 between y and PI. Equations 17 and 18 provide a fair fit to the data, as shown in Figure 3-9, if one assumes a unique relationship between liquidity index and coefficient of permeability and if one assumes that Equation 16 can be expressed in the form of Equation 12 over a wide range of values, which, strictly speaking, is not mathematically correct. Because of the variability and lack of a very good correlation, preliminary estimates of the y and 6 parameters may be made using the following simplified relationships which essentially provide as good a fit to the data, as illustrated in Figure 3-9: y = (9 Pl)-3 (19) 6 = 4.0 (20) where PI is expressed in percentage, and k in Equation 12 is assigned units of cm/sec. As noted previously, the empirical relationships presented above assume that Equation 16 can be expressed in the form of Equation 12 over a wide range of values, which, strictly speaking is not mathematically correct. Note that Equations 19 and 20 were developed based on empirical correlations with liquidity index by performing regression analyses for phosphatic clays with varying plasticity over a wide range of coefficients of permeability ranging from IO5 to lo- cm/sec. The effect of void ratio and/or coefficient of permeability regression range on the resulting y and 6 parameters is quite significant as shown by the data in Table 3-3, and as illustrated in Figure 3-10, for three clays that cover the range of plasticity of all phosphatic clays evaluated. Note that the 18 cm/sec permeability range corresponds to very high effective stresses, and is of little practical significance to mine planners. Hence, the y parameter is expected to be greater than determined from Equation 19 which is based on a 1 OW5 to IO cm/sec regression range. The 6 parameter on the other hand is anticipated to be lower than indicated by Equation 20, but not significantly lower. It is, therefore, recommended that for preliminary estimates of the y parameter, one of the following modified relationships be used: y = 5(40 Pl) ; where f = 1.0 to 3.5 (21) y = 5 (9 Pl) 3 ; where c = 1.0 to 3.0 (22) The factor 5 is expected to vary from about 1.0 to 3.5, depending on the regression range of interest. In the absence of relevant information, values of 1.5 and 2.0 may be selected for preliminary estimates using Equations 21 or 22, respectively. A 6 value on the order of 3.7 may also be used, i.e., a somewhat lower value than inferred from Equation 20 above. 3-7

69 It should be emphasized that the data in Figures 3-9 and 3-10 and the empirical correlations y-pi and &PI presented above do not account for the variability depicted in Figure 3-8, because they assume a unique relationship between liquidity index and coefficient of permeability. y and 6 parameters backfigured from actual slurry consolidation test data using the fifitting procedure are presented in Figure 3-11 and compared to y and 6 values predicted from the empirical relationships (i.e., Equations 20 and 22). As previously noted, no correlation is evident between the y and 6 parameters determined directly from actual test data and the plasticity index of phosphatic clays. This underscores the importance of performing slurry consolidation tests on site specific phosphatic clays. Nevertheless, y and 6 relationships backfigured from empirical correlations with liquidity index (and, hence, plasticity index) generally bracket the range of the directly measured values. The empirical correlations can, therefore, be used in preliminary predictions of the consolidation properties of phosphatic clays. As such, they are invaluable to mine planners in evaluating the merits and disadvantages of various phosphatic clay disposal methods. Figure 3-12 compares the y-pi and &PI empirical correlations with corresponding parameters determined from conventional consolidation test data, a rather simplified test procedure suggested in Volume 4, Wissa et al. (1983). The simplified test procedure relies heavily on data at relatively high stress levels and should, therefore, only be used in the absence of slurry consolidation test data. Again, it is evident that the empirical correlations yield y and 6 values that fall within the observed range of test data. It should also be emphasized that even y values backfigured from slurry consolidation test data often underpredict the in situ coefficient of permeability of phosphatic clays. As detailed in Volume 4, Wissa et al. (1983) the coefficient of permeability backfigured from consolidation tests, k(&) or k(log t), is a function of the curve fitting procedure adopted for evaluating the test data (Car log t).* Moreover, direct measurements of the coefficient of permeability, k,, often yield results that are greater than values backfigured from curve fitting procedures of consolidation test data. Even the recommended square root of time (6) curve fitting procedure occasionally underestimates the coefficient of permeability such that k,/kflcould range from about 1.0 to as much as 2.0 (Volume 4, Wissa et al., 1983). It is, therefore, recommended that y values backfigured from consolidation test data be adjusted prior to use in consolidation models, in accordance with the following equation: Yfield = CYlab ; where 5; = 1.0 to 3.0 (23) where the factor c is expected to vary from 1 to 3, depending on site specific conditions. Equation 23 also accounts for the fact that the in situ coefficient of permeability is often greater than backfigured from laboratory tests on phosphatic clay samples. The factor 5 should be determined at each mine site by calibrating the laboratory test data with field performance measurements. In the absence of field data, 5 = 2.0 may be used and a sensitivity study performed to evaluate the impact of higher and lower 5 values. On the other hand, 6 values backfigured from consolidation test data need not be adjusted, as indicted by the following expression: Note that since they and 6 coefficients backfigured from consolidation tests need to be adjusted to accurately model field conditions (in accordance with Equations 23 and 24), the use of the *In general k(6) = k(log t). 3-8

70 empirical relationships between the y and 6 parameters and plasticity index (Equations 20, 21 and 22) is justified for use in preliminary evaluations in spite of the fact that the y and 6 parameters are not strongly correlated to plasticity. 3.3 Sand-Clay Mixes The consolidation behavior of sand-clay mixes was investigated in Volume 4, Wissa et al. (1983), at sand-clay ratios (by dry weight) ranging from 1:1 to 3:1. The three clays selected for the sandclay mix study bracketed the range of plasticity and behavior reported for phosphatic clays, and sand tailings was used in preparing the mixes. Detailed results were presented in Volume 4, Wissa et al. (1983), and summarized in Volume 6, Wissa et al. (1983). This section summarizes and extends previous findings. In evaluating the behavior of sand-clay mixes, it is useful to define both the solids content of the combined sand and clay, or total solids content, and that of the clay phase alone (assuming all free water is within the clay phase), or clay solids content. The total solids content, S,, and clay solids content, S,, are related by the expressions: s, = (1 +SCR)/((1/S c ) + SCR) (25) S, = 1/(((1 +SCR)/S,) - SCR) (26) where SCR is the sand-clay ratio by dry weight. Likewise, the total void ratio, e,, and clay phase void ratio, ec, are related to the total solids content and to each other by the following equations: e, = ~(1 -SJ/S, (27) e, = e,(l + Go,sW~d (28) where p, the effective specific gravity of the solids in the sand-clay mix, is related to pc, the specific gravity of clay, and ps, the specific gravity of sand by the following equation: P = P* AU + SW4 PSR + PJ (29) Sedimentation of Sand-Clay Mixes Settling tests on sand-clay mixes indicated that clays which settle to a relatively low final clay solids content without sand also settle to a relatively low final clay solids content with sand. The addition of sand tailings does little to increase the final clay solids content of a relatively poor settling clay relative to that of a relatively good settling clay. Table 3-4 summarizes results from 30-day settling tests on sand clay mixes prepared with the Agrico-Saddle Creek (PI=222%), CF Mining-Hardee (PI=113%), and USSAC-Rockland (PI=160%) phosphatic clays. The results in Table 3-3 were manipulated to determine a relationship between the settled solids content of the sand-clay mix as a function of sand-clay ratio. This relationship was found to be dependent on the plasticity of the clay and the initial solids content of the clay in the mix, prior to settling. The following equation may be used in the absence of settling test data: S, = A + K (SCR) (30) 3-9

71 where Sti is the initial settled solids content of the mix for use in consolidation predictions (S,, in Table 3-4), expressed as a percentage, SCR is the sand-clay ratio by dry weight, and L and IC are constants that are a function of the plasticity of the clay and the initial solids content of the clay in the mix prior to settling. The following expressions may be used to determine ;Z and K: where PI is the plasticity index of the clay and S, is the initial solids content of the clay in the mix prior to settling, both expressed as a percent. Equations 30 through 32 yield estimates of the settled solids content of the mix generally within ±1% for the mixes listed in Table 3-4. Hence, these equations may be used for predicting the settling behavior of sand-clay mixes in the absence of site specific settling test data.* Void Ratio Versus Effective Stress Relationship As detailed in Volume 4 and summarized in Volume 6, Wissa et al. (1983), the assumption of a unique relationship between clay void ratio and effective stress appears reasonable for sand-clay mixes particularly for sand-clay ratios up to 2:1. The degree of deviation from the unique relationship can be significant at sand-clay ratios greater than 2:1 especially for relatively low plasticity phosphatic clays and/or at high effective stresses, due to sand grain contacts. For preliminary evaluations, however, the assumption of a unique relationship is appropriate, and empirical correlations developed for the phosphatic clays may be used. The effect of sand-clay ratio on compressibility can be evaluated by considering the effect of sand-clay ratio on the a and p compressibility parameters for relationships of the form: et = a2 t C 4 where a, and a, are the a parameters for clay void ratio, ecr and total void ratio, et, relationships with effective stress, respectively. The assumption of a unique relationship between clay void ratio and effective stress requires that the /? parameter be independent of sand-clay ratio, SCR, such that: Pt = PC = P Using the above assumption, a relationship between at and a, can be developed via Equations 33 and 34, as follows: a,fa, = eje, (36) * A multivariate regression analysis of the data in Table 3-4 yields the following equation, which exhibits a correlation coefficient of 0.98, and which may be used in lieu of Equations 30 to 32: Sti = S, SCR PI. 3-10

72 Noting that e, is related to e, via Equation 28, one can show that: where pc and ps are the specific gravity of the clay and sand, respectively. Assuming pjps=l, Equation 37 becomes: a t = aj(1 +SCR) (38) Data in Figure 3-13 show excellent agreement between Equation 38, which was derived theoretically, and data generated from consolidation tests on sand-clay mixes. Moreover, data in Figure 3-14 indicate that a more reliable estimate of the /It parameter for sand-clay mixes may be obtained from the following expression: /tit = /3,( SCR) (39) Equation 39 is not significantly different from Equation 35 which assumes a unique relationship between clay void ratio and effective stress, particularly for low sand-clay ratios. Accordingly, Equations 38 and 39 should provide good estimates for the at and /3, parameters characterizing the compressibility of sand-clay mixes, in conjunction with Equation Void Ratio Versus Coefficient of Permeability Relationship. As with the clay void ratio versus effective stress relationship, the assumption of a unique relationship between clay void ratio and coefficient of permeability appears reasonable at least for sand-clay ratios up to 2:1. For preliminary evaluations, therefore, the assumption of a unique relationship is appropriate, and empirical correlations developed for the phosphatic clays may be used. The effect of sand-clay ratio on the coefficient of permeability can be evaluated by considering the effect of sand-clay ratio on the y and 6 parameters for relationships of the form: k = Yte:t (41) where Yc and Yt are the Y parameters for clay void ratio, e,, and total void ratio, e,, relationships with the coefficient of permeability, respectively. Following a similar derivation as in the preceding section, the assumption of a unique relationship between clay void ratio and coefficient of permeability leads to the following equations: 6, = 6, = 6 YJY, = (e,le*)* (42) (43) Equation 43 yields the following (based on the relationship between e, and e,): l/t = YJ +CohsWW* (44) 3-11

73 which may be approximated by the following: yt = ~~(1 +SCR)* (45) Note that, in the absence of test data, the 6 parameter for phosphatic clays may be estimated from Equation 18 as: 6 = 4.OkO.25 (46) Hence, Equation 45 becomes: yr = y,(l +SCR)4 (47) Data in Figures 3-15 and 3-16 support the validity of Equations 42 and 47 for use in conjunction with Equation 41 to determine the void ratio and coefficient of permeability relationship of a given sand-clay mix, in the absence of site specific test data. 3-12

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94 Section 4 PERFORMANCE PREDICTION USING PROGRAM SLURRY Of the three disposal areas selected for the Phase II field exploration (Section 2), only the CF Mining-Hardee Reclamation Area R-2 had detailed and reliable filling records. Reclamation Area R-2 was, therefore, selected for calibrating the predictive methodology. Settling Area R-2 was not only investigated in conjunction with this FIPR-sponsored research project, but even more so as part of another FIPR-sponsored study titled Field Evaluation of Sand-Clay Reclamation by Garlanger and Babcock (1988). Modeling of the filling and consolidation of CF Mining s Hardee Reclamation Area R-2 was made for both the east and west compartments, using laboratory measured properties as well as empirically established correlations with plasticity index. The predictions were made using a refined version of the computer program SLURRY (Volume 6; Wissa et al, 1983). The estimated input quantities of sand and clay based on plant production figures were matched at each point in time with surface elevations and corresponding storage volumes to calculate the average percent clay solids and sand-clay ratio as a function of time for the west and east compartments. The data is presented in Tables 4-1 and 4-2 along with results of direct sampling from each sampling trip. As shown, the field sampling data are in good agreement with values based on mass-volume calculations, indicating that the field data collected are representative for the area. 4.1 Filling Schedule and Initial Slurried Conditions Table 4-3 summarizes data obtained on random samples of sand-clay mix pumped from the mixer to the reclamation area. As indicated by the data, the average initial solids content of the mix equalled 37.0%, and the initial clay solids content, Si,, equalled 15.2%. The corresponding sand-clay ratio, SCR, equalled Filling data in Table 4-4 reflect a somewhat lower SCR on the order of 1.9. Figure 4-1 shows the actual filling schedule in terms of tons of clay compared to the filling schedule modeled in program SLURRY for each of the two compartments. The following initial sand-clay mix conditions were used in predictions: 4.2 Effect of Sedimentation Manipulation of the laboratory settling test results summarized in Table 3-4 for the CF Mining- Hardee sand-clay mixes yields the following relationship between the settled solids content, Stir* *Denoted as S,, in Table 3-4.

95 of the sand-clay mix, the sand-clay ratio, SCR, and the initial clay solids content prior to settling, sic: Sti = ( (S,,-8)) + ( O.O563(S,-8)) SCR where Sti and Si, are both expressed as a percentage. Substituting values of SCR and S, of 1.9 and 15.2%, respectively, results in the following value for the settled solids content of the mix for use in consolidation predictions: Settled Total Solids Content, Sti = 46.6%. Note that use of the generalized equations presented in Section 3 (Equations 30, 31a and 32) in conjunction with the plasticity index of the CF Mining-Hardee clay (PI=113%) yields a slightly higher value of S&, on the order of 47.4%. 4.3 Consolidation Properties Predictions were made using three sets of consolidation properties. One set was based on regression of actual laboratory test data on CF Mining-Hardee sand-clay mixes (Case 1). The second was based on regression of actual laboratory test data on CF Mining-Hardee phosphatic clay, and on the assumption of unique relationships between clay void ratio and effective stress as well as between clay void ratio and coefficient of permeability (Case 2). The third was based on empirical relationships established in Section 3 based on empirical correlations with liquidity index, using a plasticity index value PI of 113% for the CF Mining-Hardee phosphatic clay (Case 3). All consolidation parameters presented herein correspond to effective stress values expressed in units of kg/cm 2, and coefficient of permeability values expressed in units of cm/sec Case 1: Based on Sand-Clay Mix Data Figures 4-2 and 4-3 present actual test data obtained on the CF Mining-Hardee phosphatic clay, as well as on sand-clay mixes, from settling tests, and conventional, slurry and constant rate of strain consolidation tests. Volume 4, Wissa et al (1983), presents and discusses these data in detail. Regression of the actual test data obtained at sand-clay ratios, SCR, of 0, 1 and 3 yielded relationships for the compressibility parameters a, and pt as a function of SCR presented in Figures 3-13 and 3-14, respectively, and relationships for the coefficient of permeability parameters yt and 6, as a function of SCR presented in Figures 3-15 and 3-16, respectively. Interpolation of the specific relationships for the CF Mining-Hardee sand-clay mixes yields the following consolidation parameters at an SCR of 1.9: yt = 0.78 pi = yt = 7x1 o-8 6, = 3.0 Note that the linear regression used to determine the coefficient of permeability parameters, yt and Pt, for the sand-clay mix is poor and unreliable because it is based on limited test data obtained from conventional consolidation tests at relatively high stresses, beyond the stress 4-2

96 range of interest (see data in Figure 4-3 for SCR s of 1 and 3). Hence, in the absence of additional test data at lower stresses, Case 1 is not expected to reliably model field conditions Case 2: Based on Unique Relationships with Clay Void Ratio As shown on Figures 4-2 and 4-3, the assumptions of unique relationships between clay void ratio and effective stress and between clay void ratio and coefficient of permeability provide a good fit to the test data, particularly at sand-clay ratios up to 2. For the CF Mining-Hardee phosphatic clay, the following consolidation parameters were determined based on regression of actual test data: yc = PC = yc = 5.44x1 o-l0 6, = 4.08 For the CF Mining-Hardee sand-clay mix, the specific gravity of the clay and sand, pc and ps, equal and 2.700, respectively. The specific gravity of the mix, p, at a sand-clay ratio SCR of 1.9 is calculated to be (Equation 29; Section 3). The compressibility parameters a, and,& of the mix can be determined at an SCR of 1.9 from Equations 35 and 37 (Section 3), and the coefficient of permeability parameters can be similarly obtained from Equations 42 and 44 (Section 3). The assumptions of unique relationships between clay void ratio and effective stress and clay void ratio and coefficient of permeability yield, therefore, the following consolidation parameters at an SCR of 1.9: yt = 0.80 pt = yt = 4.7x1 o-8 6, = Case 3: Based on Empirical Correlations For a clay plasticity index, PI, of 113%, empirical relationships (Equations 10, 11, 19 and 20) established in Section 3 based on regression of empirical correlations with liquidity index yield the following consolidation parameters for the phosphatic clay: yc = 2.46 PC = yc = 9.5x1 o-l0 6, = 4.0 Using the approximate equations presented in Section 3 for the sand-clay mix (Equations 38, 39, 42 and 47) at an SCR of 1.9, the following consolidation parameters are estimated for the mix based on empirical correlations: yt = 0.85 pt = yt = 6.7x1 O-* 6, =

97 4.4 Predictions Based on the clay filling quantities presented in Figure 4-1, a sand-clay ratio, SCR, of 1.9 and an initial settled total solids content, Si, of 46.6%, program SLURRY was used to predict the height of fill and clay solids content as a function of time, for both the west and east compartments of CF s Reclamation Area R-2. Predictions were made using three sets of consolidation properties corresponding to the three cases presented in Section 4.3. Other input parameters included the specific gravity of the mix b=2.736), and double-drainage conditions as substantiated by field piezometric data (Section 2.5). At the end of filling, the water table was lowered by 2.5 feet to simulate the effects of a desiccated crust (substantiated by field water level and elevated surficial solids content data). Figures 4-4 and 4-5 present results of predictions for the height of the settled mix and the average clay solids content for the west and east compartments, respectively, and compare the predictions to average measured field data as a function of time. As shown, and as expected in light of the discussion in Section 4.3.1, Case 1 did not reliably model the field behavior and significantly underpredicted the rate of consolidation. Both Cases 2 and 3 provided a somewhat better fit to the data, yet still underpredicted the rate of consolidation. 4.5 Model Calibration As detailed in Section 3, the permeability parameter, y, often needs to be multiplied by a correction factor c to account for the fact that the in situ coefficient of permeability is often greater than backfigured from laboratory tests (Equation 23; Section 3), and/or to account for the effect of the regression range on the y value backfigured from empirical correlations (Equation 22; Section 3). As a result, additional predictions were performed for the following cases: Case 2A: Case 2 but with 5 = 2.0, i.e., yt = 9.4x10-@ Case 3A: Case 3 but with < = 1.5, i.e., yt = 1.01x1 0 Figures 4-6 and 4-7 compare predictions of height of fill and average solids content for these two cases with corresponding field data. As shown, the predictions closely match the actual field data. It is perhaps fortuitous that the model based on empirical correlations (Case 3A) provided as good a fit to the field data in this case as site-specific consolidation properties determined from laboratory tests (Case 2A). For illustration purposes, Appendices A and B present the input and output data files for the west and east compartments, respectively, generated via program SLURRY for Case 2A. Figure 4-8 compares predicted and measured clay solids contents versus depth in the west compartment, as of May Note that the predicted values based on a SCR of 1.9 are in good agreement with measured values for samples exhibiting sand-clay ratios ranging from 1.5 to 2.5 (except within the desiccated crust). Differences between predicted and measured data are attributed to desiccation effects, variations in sand-clay ratio in the reclamation area, non-homogeneity of the in situ sand-clay mix, and/or variations in clay plasticity. 4-4

98 4.6 Practical Implications Comparison of field performance with predictions made using program SLURRY and laboratory as well as empirically determined consolidation properties confirms the validity of the predictive methodology developed as part of this FIPR-sponsored research program. In order to get a good fit to the field data, however, the coefficient of permeability (and/or y parameter in a log void ratio versus log coefficient of permeability relationship) determined from laboratory tests needs to be multiplied by a correction factor 5 which is expected to vary from 1.0 to 3.0, depending on site specific conditions. 4-5

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111 Section 5 REFERENCES Baligh, M.M. (1975). Theory of Deep Site Static Cone Penetration Resistance. Research Project R75-56, No. 517, Department of Civil Engineering, Massachusetts institute of Technology. Carrier, W.D., Bromwell, L.G. and Somogyi, F. (1981). Slurried Mineral Wastes: Physical Properties Pertinent to Disposal. Waste Management: Engineering Solutions, ASCE National Convention, St. Louis. Garlanger, J.E. and Babcock, J.W. (1988). Field Evaluation of Sand-Clay Mix Reclamation-Phase II. Florida Institute of Phosphate Research, Research Project FIPR Vesic, A.S. (1975). Principles of Pile Design. Lecture Series on Deep Foundations Sponsored by the Geotechnical Group, BSCES/ASCE in cooperation with Massachusetts Institute of Technology. Wissa, A.E.Z., Fuleihan, N.F. and Ingra, T.S. (1982). Evaluation of Phosphatic Clay Disposal and Reclamation Methods, Volume 1, Index Properties of Phosphatic Clays. Florida Institute of Phosphate Research, Research Project FIPR Wissa, A.E.Z., Fuleihan, N.F. and Ingra, T.S. (1983). Evaluation of Phosphatic Clay Disposal and Reclamation Methods, Volume 2, Mineralogy of Phosphatic Clays. Florida Institute of Phosphate Research, Research Project FIPR Wissa, A.E.Z., Fuleihan, N.F. and Ingra, T.S. (1983). Evaluation of Phosphatic Clay Disposal and Reclamation Methods, Volume 3, Sedimentation Behavior of Phosphatic Clays. Florida Institute of Phosphate Research, Research Project FIPR Wissa, A.E.Z., Fuleihan, N.F. and Ingra, T.S. (1983). Evaluation of Phosphatic Clay Disposal and Reclamation Methods, Volume 4, Consolidation Behavior of Phosphatic Clays. Florida Institute of Phosphate Research, Research Project FIPR Wissa, A.E.Z., Fuleihan, N.F. and Ingra, T.S. (1983). Evaluation of Phosphatic Clay Disposal and Reclamation Methods, Volume 5, Shear Strength Characteristic of Phosphatic Clays. Florida Institute of Phosphate Research, Research Project FIPR Wissa, A.E.Z., Fuleihan, N.F. and Ingra, T.S. (1983). Evaluation of Phosphatic Clay Disposal and Reclamation Methods, Volume 6, Predictive Methodology for Evaluating Disposal Methods. Florida Institute of Phosphate Research, Research Project FIPR Wissa, A.E.Z., Fuleihan, N.F. and Ingra, T.S. (1985). Evaluation of Phosphatic Clay Disposal and Reclamation Methods, Volume 7, Engineering Properties of Flocculated Clays. Florida Institute of Phosphate Research, Research Project FIPR

112 Appendix A Program SLURRY Input/Output Data West Compartment Reclamation Area R-2 Case 2A

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124 Appendix B Program SLURRY Input/Output Data East Compartment Reclamation Area R-2 Case 2A

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