Analysis and Interpretation of In Situ Rock Bolt Pull Tests in Hard Rock Mines. Luke Nicholson

Transcription

1 Analysis and Interpretation of In Situ Rock Bolt Pull Tests in Hard Rock Mines by Luke Nicholson A thesis submitted in conformity with the requirements for the degree of Masters of Applied Science Graduate Department of Civil Engineering University of Toronto Copyright 216 by Luke Nicholson

2 Abstract Analysis and Interpretation of In Situ Rock Bolt Pull Tests in Hard Rock Mines Luke Nicholson Masters of Applied Science Graduate Department of Civil Engineering University of Toronto 216 Rock bolts are the principal reinforcement element of many underground support systems. This thesis investigates and characterizes the behaviour and performance of rock bolts as measured by a pull test. A database composed of 985 pull tests from six mines in the Sudbury Basin was assembled. Procedures and apparatuses used to conduct these tests were compared to ASTM s standards and ISRM s suggested methods. The results from the pull tests were used to compare the behaviour of reinforcement elements with theoretical models and to quantify performance metrics and their distributions. The influence of bolt, installation and rock mass parameters on the performance of certain rock bolts was investigated, and distributions of expected behaviour were constructed. These may be used in the design of hard rock underground excavations using methodologies that incorporate both the load capacity and displacement behaviour of rock bolts. ii

3 Acknowledgements I would like to thank my supervisor, Professor John Hadjigeorgiou, for the guidance and educational experience I have had the pleasure to experience over the last two years. This project would not have been possible without the funding and active participation by Vale. In particular, I would like to thank Dr. Mike Yao, Lindsay Moreau-Verlaan, Derek Boucher and the rest of the ground support staff at Vale s Sudbury operations for the support they provided me. I also extend my gratitude to ground support suppliers who provided technical information on testing, including Mansour Mining Technologies Inc, Jennmar Canada, DSI, Normet and Atlas Copco. In particular, Francois Charette, Lynn Mainville-Beach and Bryan Lamothe. Thank you to my friends and research associates Marie-Helene Fillion, Philippe Morissette and Stratos Karampinos for their education, support and generally putting up with me. I am very grateful to all my friends home and abroad for being there for me when I needed it and generally improving my life. Last but certainly not least, I thank my family for their advice and unlimited support on every step of my journey. iii

4 Contents 1 Introduction Problem Definition Significance Objectives Methodology Structure of the Thesis Testing of Rock Bolts Rock Bolts Laboratory Testing Methods Tension Test Wedge Tension Test Bend Test Tests of Expandable Rock Bolts Laboratory Rock Anchor Capacity Pull Test Laboratory Drop Test In Situ Pull Test Discussion of Testing Methods Summary Composition of the Database Pull Test Setting Regional Geology Coleman Mine Copper Cliff Mine Creighton Mine Garson Mine Stobie Mine Totten Mine Reinforcement Elements in the Pull Test Database Friction Rock Stabilizers Rebar Rock Bolts Modified Cone Bolt D-Bolt iv

5 3.2.5 Expandable Bolts Database Database Description Comparison to Other Pull Test Databases Specific and General Limitations of the Pull Test Database Summary Review of Implemented Pull Test Methods Implementation of Pull Tests in Practice Deviations from ASTM Standard D and ISRM Suggested Methods for Rockbolt Testing Deviations in Apparatus Deviations in Procedure Practical Considerations in Pull Testing Pull Test Data Working Capacity and the Measurement of Load Recording Displacement During a Pull Test Rock Bolt Stiffness Limitations of the Metrics Measured in a Rock Bolt Pull Test Summary Summary Statistics and Interpretation of Pull Test Data Summary Statistics and Statistical Techniques Summary Statistics Statistical Techniques Friction Rock Stabilizers Theoretical Behaviour of a Frictional Rock Stabilizer Observed Behaviour of Friction Rock Stabilizers Characterization of Performance Metrics for Friction Rock Stabilizers Rebar Rock Bolts Theoretical Behaviour of a Rebar Rock Bolt Observed Behaviour of Rebar Rock Bolts Characterization of Performance Metrics for Rebar Rock Bolts Modified Cone Bolts Theoretical Behaviour of a Modified Cone Bolt Observed Behaviour of Modified Cone Bolts Proposed Interpretation of Modified Cone Bolt Behaviour Characterization of Performance Metrics for Modified Cone Bolts D-Bolts Theoretical Behaviour of a D-Bolt Observed Behaviour of D-Bolts Characterization of Performance Metrics for D-Bolts Expandable Bolts Theoretical Behaviour of an Expandable Bolt v

6 5.6.2 Observed Behaviour of Expandable Bolts Characterization of Performance Metrics for Expandable Bolts Other Reinforcement Elements Yield-Lok Fibreglass Rebar DS Bolt Other Expandable Bolts MD Bolt Summary Factors Influencing Pull Test Performance Friction Rock Stabilizers Influence of Length Installation Method Influence of Drive Time Influence of Drill Bit Diameter Influence of Bolt Diameter Geology Rock Mass Quality Summary of Investigation on FRS Pull Tests Rebar Rock Bolts Rebar Rock Bolt Length Encapsulation Length Spin Time Residence Time Geology Summary of Investigation on Rebar Rock Bolt Pull Tests Modified Cone Bolts Residence Time Geology Inter-variable Relationships Summary of Cone Bolt Findings Summary Characterization of Rock Bolt Behaviour Characterisations of Bolt Behaviour using Laboratory Pull Tests Friction Rock Stabilizers Characterization of FRS Performance Characterization of FRS Behaviour Rebar Rock Bolts Characterization of Rebar Rock Bolt Performance Characterisation of Rebar Rock Bolt Behaviour Modified Cone Bolt Characterisation of Modified Cone Bolt Performance vi

7 7.4.2 Characterisation of Modified Cone Bolt Behaviour D-Bolt Characterisation of D-Bolt Performance Characterisation of D-Bolt Behaviour Expandable bolts Characterisation of Expandable Bolt Performance Characterisation of Expandable Bolt Behaviour Summary Conclusions Contributions Load Capacities of Reinforcement Elements Limitations Recommendations Implications and Path Forward List of References 144 Appendices 151 A Pull Testing Forms 152 A.1 ASTM D Sample Form A.2 ISRM Suggested Method for Pull Testing Data Sheet A.3 Proposed Pull Test Information Sheets vii

8 List of Figures 2.1 Types of rock bolt (Hadjigeorgiou & Charette, 21) Tensile test specimen with a reduced section (ASTM A37, 212) Apparatus for a tensile test on an FRS (ASTM F432, 213) Bolt head configuration for a wedge tension test (ASTM F66, 213) Two alternative configurations for the ferrule test (ASTM F432, 213) Apparatus for the Laboratory Rock Anchor Capacity Pull Test (ASTM D741, 28) Apparatus for the Laboratory Drop Test (ASTM D741, 28) Apparatus for a rock bolt anchor pull test (ASTM D4435, 213) Conceptual load versus bolt head deflection curve for a rock bolt pull test (ASTM D4435, 213) Apparatus for a pull test (ISRM, 1981) Map of Sudbury area showing locations of relevant mines (after Eckstrand & Hulbert, 27) Longitudinal section of Coleman Mine (Morissette et al, 214) Longitudinal section of Copper Cliff Mine (Chinnasane et al, 214) Longitudinal section of Creighton Mine (Snelling et al, 213) Longitudinal section of Stobie Mine Longitudinal section of Totten Mine (After Sudbury Platinum Corporation, 215) FRS A schematic (Courtesy of Supplier A) FRS B schematic (Courtesy of Supplier B) Schematics of rebar manufactured by Supplier C (top; courtesy of Supplier C), Supplier A (middle; courtesy of Supplier A) and Supplier B (bottom; courtesy of Supplier B) MCB33 (Courtesy of Mansour) D-Bolt schematic (Normet, 214) VersaBolt schematic (Courtesy of Mansour) Schematic of a Python bolt, and cross sections before (1) and after (2) inflation (Courtesy of Jennmar) Normet s pull test apparatus mounted on a 22 mm D-Bolt Static laboratory pull test on a 22 mm D-Bolt 2.1 m in length (left; Doucet & Voyzelle, 212) and dynamic impact test on a 22 mm D-Bolt 1.5 m in length (right; Li & Doucet, 212) Methods of determining yield strength: halt of the pointer method (left) and offset method (right; ASTM E6, 29) viii

9 4.4 Determination of working capacity from a pull test Measurement of displacement for a pull test on a generic reinforcement element Generic reinforcement system (Thompson et al, 212) Measurement of displacement for axial loading of a point anchored rock bolt Design of support systems using the ground reaction curve (Brady & Brown, 26) Calculation of secant and tangent stiffness from a pull test Pull test campaign results for partially encapsulated Rebar A from November 16 th, 212 at Coleman Mine (Mainville et al, 212) Shear stress distribution along a frictionally coupled bolt subject to axial load (Li & Stillborg, 1999) Shear stress and axial load along a Swellex rock bolt (Li & Stillborg, 1999) Pull test performed on an FA Ultimate capacity per unit length distributions for FRSs with nominal diameters of 35, 39 and 46 mm Stiffness metrics for an FRS Stiffness distributions for the FA35 and FA Model of the shear stress profile in a grouted rock bolt (Li & Stillborg, 1999) Model of tensile load and shear stress profile for a rebar (Li & Stillborg, 1999) Pull test performed on a 2 mm Rebar B Pull test performed on a 2 mm Rebar A Working capacity distributions for rebar Secant stiffness for 2 mm rebar Tangent stiffness for 2 mm Rebar Comparison of Rebar A unloading stiffness to secant (a) and tangent (b) stiffness Laboratory pull test of an MCB33 (Simser et al, 26) Conceptual load displacement behaviour of a cone bolt subject to quasi-static loading (Simser et al, 26) Pull test performed on an MCB33 with plough Pull test performed on an MCB33 without a linear plough response Close-up of the bolt response shown in Figure Amended conceptual load displacement behaviour of a cone bolt Performance metrics measured from a cone bolt pull test Load metric distributions for the MCB Stiffness metric distributions for the MCB Apparatus for a simulated joint laboratory test on a D-Bolt (Li, 212) Results of simulated joint laboratory tests on 2 mm D-Bolts (Li, 212) Pull tests performed on 22 mm D-Bolts Pull tests performed on 2 mm D-Bolts Pull tests performed on 22mm D-Bolts Distributions of stiffness for 2 and 22 mm D-Bolts Pull tests performed on Pm12 and Mn12 expandable bolts Working capacities of Swellex Pm12 and Mn Secant stiffness of Swellex variants ix

10 5.33 Secant stiffness of Swellex sorted by installation medium Relationship between ultimate capacity and length for the FA Relationship between ultimate capacity and length for the FB Relationship between ultimate capacity and length for the FB Relationship between ultimate capacity and anchorage length for all FRS bolts Comparison of ultimate capacities between jackleg and bolter installations of the FA Comparison of ultimate capacities between jackleg and bolter installations of all FRS bolts Relationship between drive time and ultimate capacity for all FRS bolts Relationship between drive time and ultimate capacity for FB35, distinguishing between installation methods Relationship between drive time and absolute ultimate capacity for all FRS configurations installed using a bolter Relationship between installation time and ultimate capacity for all bolter-installed FRS fit with linear and power functions Relationship between drill bit diameter and ultimate capacity for all FRS configurations Relationship between drill bit diameter and ultimate capacity for two testing campaigns performed on the FB Relationship between drill bit diameter and ultimate capacity for FA35 and FB35, separated by installation method Bolt diameter to drill bit diameter ratios for all FRS variants Relationship between bolt diameter to drill bit diameter ratio and ultimate capacity Relationship between bolt diameter to drill bit diameter ratio and ultimate capacity for all FRS configurations by testing campaign FRS ultimate capacity by lithology Relationship between UCS and ultimate capacity Distributions of FRS bolts installed in ore and waste rock normalized to the campaign average ultimate capacity for bolts installed in waste rock Ultimate capacities recorded for FA46 bolts installed in poor and good quality ground Ultimate capacities recorded for FRSs installed in poor and good quality ground FA39s pulled at Garson, 9/2/ Rebar performance metrics by length Relationships between stiffness and grout length : rebar length for Suppliers A and B Unloading stiffness and grout length : rebar length for Rebar A Relationship between stiffness and resin spin time for Rebar A Stiffness comparison of Rebar A installed on the day of testing versus previously Stiffness comparison between lithologies for rebar Performance comparison of MCB33s installed prior to and on the day of testing Performance comparison of MCB33 bolts installed in ore, igneous/metaigneous and metasedimentary lithologies Relationship between plough stiffness and plough point for the MCB Average load displacement behaviours obtained by Stillborg in a laboratory setting (Stillborg, 1993; composited by Hoek et al, 1995) x

11 7.2 Distributions of secant stiffness and ultimate capacity for a pull test on an FRS with an anchorage length of 1.52 m Ultimate capacity per metre distributions for all FRS configurations Distribution of secant stiffness and ultimate capacity for all FRS bolts tested All FA35 and FA39 pull test load-displacement relationships Distribution of displacement measured during pull testing of FA35 and FA39 with anchorage lengths of 1.52 m Conceptual displacement distribution of FRSs with anchorage lengths of 1.52 m subject to a pull test Conceptual displacement distribution of pull tests performed on FRSs Distributions of secant stiffness and working capacity for a pull test with a pre-load of 17.8 kn on a rebar rock bolt All rebar pull test load-displacement relationships Distribution of displacement measured during pull testing of 2 mm rebar 1.8 m to 2.4 m in length with a pre-load of 17.8 kn Distribution of displacement measured during pull testing of 2 mm rebar 1.8 m to 2.4 m in length without a pre-load Conceptual distribution of displacement during pull testing of 2 mm rebar without a pre-load Distributions of secant stiffness and working capacity for a pull test with a pre-load of 17.8 kn on an MCB MCB33 pull tests collected from Vale s Sudbury operations Distribution of displacement measured during pull testing of a 2.4 m MCB33 with a pre-load of 17.8 kn Conceptual distribution of displacements for a pull test of a 2.4 m MCB33 without a pre-load Distribution of secant stiffness for a pull test with a pre-load of 17.8 kn on a 2 mm D-Bolt Distribution of secant stiffness for a pull test with a pre-load of 17.8 kn on a 22 mm D-Bolt mm D-Bolt pull tests collected from Vale s Sudbury operations mm D-Bolt pull tests collected from Vale s Sudbury operations Load displacement behaviour of a 2 mm D-Bolt with a pre-load of 17.8 kn Load displacement behaviour of a 22 mm D-Bolt with a pre-load of 17.8 kn Conceptual distribution of displacements for a pull test of a 2 mm D-Bolt without a pre-load Conceptual distribution of displacements for a pull test of a 22 mm D-Bolt without a pre-load Distributions of secant stiffness and working capacity for a pull test on Swellex Pm12 and Mn12 without a pre-load Swellex Pm12 and Mn12 pull tests collected from Vale s Sudbury operations Swellex Pm24 and Mn24 pull tests collected from Vale s Sudbury operations Load displacement behaviour of Pm12 and Mn12 bolts Load-displacement behaviour of Pm24 and Mn24 bolts Conceptual load displacement behaviour of Pm12 and Mn12 bolts subject to a pull test. 135 xi

12 7.32 Conceptual load displacement behaviour of Pm24 and Mn24 bolts subject to a pull test Conceptual load displacement behaviour extrapolated to load for various rock bolts subject to a pull test Median load displacement behaviour for all bolts with no pre-load A.1 Rock bolt pull test sample form (ASTM D4435, 213) A.2 Rock bolt pull test data sheet (After ISRM, 1981) A.3 Proposed pull test campaign information sheet A.4 Proposed pull test data sheet xii

13 List of Tables 3.1 Intact rock UCS values in MPa by mine site Comparison of FRS supplier specifications (Courtesy of Suppliers A and B) Steel properties for a 2 mm threaded rebar (Courtesy of Suppliers A, B and C) MCB33 mechanical properties (Courtesy of Mansour) VersaBolt and D-Bolt mechanical properties (Courtesy of Mansour, Normet) Summary of Swellex, Omega and Python bolt mechanical properties (Atlas Copco, 212; Courtesy of Jennmar, DSI) Number of pull tests by bolt type General One-Way ANOVA table (NIST-SEMATECH,23) Statistics regarding the ultimate capacities of FRSs FRS stiffness summary statistics Summary statistics for the working capacity of rebar Summary statistics for the stiffness of rebar Comparison of stiffness between partially and fully encapsulated rebar Summary statistics of MCB33 load metrics Summary statistics of MCB33 stiffness metrics D-Bolt stiffness by anchor stability Unloading stiffness calculated for pull tests performed on 22 mm D-Bolts Working capacities obtained from all D-Bolt pull tests D-Bolt summary statistics Swellex behaviour breakdown Coupling strength of partially embedded and slipped Swellex pull tests Summary statistics of Swellex Pm12 and Mn12 working capacity Swellex secant stiffness summary statistics Secant stiffness summary statistics on Swellex sorted by installation medium (2.44 m) Yield-Lok pull test result summary DS Bolt campaign summary Summary of expandable bolt campaigns not including Swellex Summary of working capacities for all bolts pull tested Statistics on the relationship between ultimate capacity and length for the FA Statistics on the relationship between ultimate capacity and length for the FB Statistics on the relationship between ultimate capacity and length for the FB xiii

14 6.4 Single factor ANOVA performed on the relationship between ultimate capacity and length for the FB Breakdown of bolts contributing to major sets of anchorage length Comparison of 1.52 m and 1.83 m of anchorage length for all FRS bolts Ultimate capacity statistics for jackleg and bolter installations of the FA Ultimate capacity statistics for jackleg and bolter installation of all FRS bolts Description of relationships between drive time and ultimate capacity for FRSs installed with a MacLean Bolter Summary statistics for bolt diameter to drill bit diameter ratios for all FRS variants Comparison of average ultimate capacities for pull tests performed in ore and waste rock in the same campaign Comparison of FRS bolts installed in ore and waste rock Comparison of FA46 bolts installed in poor and good quality ground Comparison of all FRSs installed in poor and good quality ground Comparison of partial and full encapsulation test statistics for Rebar A and B Statistics regarding residence time for Rebar A Comparison of stiffness across different lithologies Performance comparison of MCB33s installed prior to and on the day of testing Performance comparison of MCB33 bolts installed previously and on the day of testing Summary of observed relationships of between various factors and performance indicators for each rock bolt type FA35 and FA39 performance percentiles Distribution of ultimate capacity across all FRS configurations FRS performance compared to supplier specifications Percentiles of performance metrics for rebar rock bolts Rebar performance compared to supplier specifications Distributions of MCB33 performance metrics MCB33 performance compared to supplier specifications Secant stiffness distributions of 2 mm and 22 mm D-Bolts Swellex secant stiffness percentiles Swellex Pm12 and Mn12 performance compared to supplier specifications xiv

15 Chapter 1 Introduction 1.1 Problem Definition An effective ground support system has two primary functions: to strengthen the rock mass with the use of reinforcement elements, and to retain and hold fractured rock using surface support (McCreath & Kaiser, 1992). During the development of a rock support strategy, five basic questions should be addressed: where do stability problems exist, why do they exist, what reactions are necessary to alleviate them, when should support be installed, and how should the strategy be implemented (Thompson et al, 212). In order to determine the what and when of support system reactions, an understanding of the behaviour of a support system (and thus its constitutive elements) in terms of load capacity and the corresponding displacements is necessary. In its most elementary form, the design of a ground support system for an underground excavation can be seen as a force balance problem; an unstable portion of a rock mass must have its mass supported by a stable portion of the rock mass through load transfer along a reinforcement element (Windsor & Thompson, 1992). In theory, a design methodology may consider the characteristic force displacement relation of an excavation s failure mechanism (such as Windsor, 1997). Such a methodology incorporates the development of displacement with load, as well as the load capacity of the support system. Rock bolts are the principal reinforcement elements used in underground support systems (Hadjigeorgiou & Charette, 21). While mechanical properties such as tensile strength or elastic modulus of the materials used to manufacture rock bolts may be quantified in laboratory conditions, there exists an incomplete understanding of how the same rock bolts perform in situ. Various parameters that may influence how a bolt performs include those associated with the installation of the bolt, the nature of the rock mass the bolt is installed in, and the physical characteristics of the bolt itself. One method of quantifying bolt performance is the pull test, in which an increasing load is applied to the head of a rock bolt installed in a rock mass until either a predetermined load is reached or a certain behaviour (such as yield, slip or failure of the bolt) is observed. Pull testing campaigns generally include a small number of bolts (usually 5 to 1) installed and tested under very similar conditions. There is a lack of large-scale, in-depth analysis across different testing campaigns on the part of the mine sites to investigate variability in bolt performance. This limitation is addressed in this thesis by reviewing pull test results from six mines sites across the Sudbury Basin. 1

16 Chapter 1. Introduction Significance The appropriate use of rock bolts in underground mines represents a cost-benefit problem. Over design of ground support systems may lead to inefficiencies in labour and equipment usage, inflating costs and cycle times. Under design may result in serious safety concerns and production stoppages. An important step in the design and optimization of a ground support system is the selection of appropriate elements and systems for a set of conditions. The performance that may be expected from a particular element given a set of installation parameters and its associated variability should be recognised by those who design the support system. With improvements in the understanding, measurement, and quantification of performance, safer and more cost effective ground support systems may be developed. 1.3 Objectives This thesis aims to provide information that can be used for design purposes based on in situ pull tests rather than laboratory results. There are three principal objectives of this thesis. The first is to develop an understanding of the load displacement behaviour of various rock bolts, and use it to interpret pull test data. The second is to identify relationships between recorded pull test parameters and the performance of the rock bolt. The third is to develop input parameters in terms of bolt load capacity and displacement that can be used for design purposes in hard rock mines. 1.4 Methodology The progression of this thesis can be broadly summarized in the following sequence: 1. Data Collection: Pull test campaign reports from several rock bolt suppliers at six Vale mine sites in the Greater Sudbury area were collected, and their results and all recorded testing parameters input into a database. The majority of the pull tests used for this analysis were performed between 211 and Data Analysis: Statistical methods including linear regression and analysis of variance (ANOVA) amongst others were used to quantify the performance of the rock bolts tested, as well as investigate the influence of various factors on bolt performance. 3. Development of Performance Guidelines: Load displacement data from individual pull tests was analysed and combined to create performance envelopes for various rock bolts. 1.5 Structure of the Thesis The structure of the thesis is as follows: Chapter 1: Introduction - The need for data regarding rock bolt capacity is addressed. The methodology used to undertake the thesis is laid out, and the thesis structure stated. Chapter 2: Testing of rock bolts A review of rock bolts used in underground hard rock mines is presented. Standardized testing methods used to characterize rock bolt behaviour are identified, including

17 Chapter 1. Introduction 3 both laboratory and field tests. Chapter 3: Composition of the database The mines from which the pull tests were collected, the rock bolts tested, and the database used for analysis are described. Chapter 4: Review of implemented pull test methods Typical pull test methodologies and apparatuses used in the constructed database are described. Deviations from standards and suggested methods are identified. The information obtained from a pull test is discussed. Chapter 5: Summary statistics and interpretation of pull test data Methods of data analysis are described. A theoretical understanding of rock bolt behaviour is used to interpret pull tests, and performance metrics are quantified. Chapter 6: Factors influencing pull test performance - The influence of a variety of individual factors on rock bolt performance are examined and discussed. Chapter 7: Characterization of rock bolt behaviour The load displacement behaviours of several reinforcement elements subject to a pull test are quantified and described. Chapter 8: Conclusions The main conclusions of this thesis are presented. Contributions and limitations of the current work are identified, and recommendations are made for future work.

18 Chapter 2 Testing of Rock Bolts In this chapter, rock bolts are categorized, and various laboratory and field testing methods applicable to rock bolts are described. While the focus of this thesis is on in situ pull tests (described in Section 2.3), a comparison between testing methods targeting a well-defined set of material properties and a test more representative of in situ performance is of value. 2.1 Rock Bolts Rock bolts are the primary rock reinforcement element in most underground mines (Hadjigeorgiou & Charette, 21). Various mechanisms by which load is transferred from the element to the rock mass are used by different types of bolt. Rock bolts can be broadly categorized into three groups; mechanical, grouted (or resin), and friction bolts (Figure 2.1). Mechanical Bolts Resin Bolts Paddle bolt Posimix Solid deformed Wriggle bolt Chemical anchor Friction bolts Split set Swellex Figure 2.1: Types of rock bolt (Hadjigeorgiou & Charette, 21) 4

19 Chapter 2. Testing of Rock Bolts 5 Mechanical bolts are steel tendons fitted with an expansion shell that anchors the bolt at the toe of the hole, pressing against the sides of the hole as the bolt tendon is torqued and tensioned. Grouted bolts, such as rebar, may use either a cement or resin grout to bond the bolt to the rock mass continuously along its length. Changes in the pattern on the bolt, bolt shape (e.g. paddle bolts) and fixtures on the bolt (e.g. Posimix) are used to more effectively mix the grout and/or increase the strength of the bolt resin interface. A friction bolt also transmits load from the bolt to the rock mass continuously along its length, but does so using frictional resistance. An example is the friction rock stabilizer (FRS), composed of a steel tube with a gap in the circumference running along its length, tapered on one end and with a steel ring welded to the other (or alternatively the bolt s head is crimped). An FRS is installed in a drill hole with a diameter slightly smaller than that of the bolt, compressing the bolt radially and generating forces that result in frictional resistance to pull out. The Swellex bolt is another example of a friction bolt. In recent years since the expiration of Atlas Copco s patent, several similar bolts from different manufacturers have become available. These are commonly referred to as expandable or inflatable bolts. Expandable bolts are rock bolts that are expanded using pressure exerted by a water pump once the bolt is placed loosely in a hole. Resistance to pull out is not only frictional, but is also attributed to mechanical interlock between any undulations in the sides of the drill hole and the resulting shape of the inflated bolt (Hadjigeorgiou & Charette, 21). More recently, a distinction is made for yielding or energy-absorbing rock bolts. While these are often grouted bolts, they are specifically designed to address stability problems caused by high stress, such as rock bursting or squeezing (Li et al, 214). Yielding bolts generally have larger deformation and energy capacities than bolts only intended for use in static loading scenarios. This is usually achieved through one of two mechanisms. The first is an increase in uncoupled tendon length between two or more anchor points. An example of a bolt that uses this mechanism is the D-Bolt (available from Normet), where a smooth bar deforms between anchors set in grout. The second mechanism is movement of the bolt or a section of the bolt relative to an internal fixture or the drill hole; for example a Modified Cone Bolt (available from Mansour) ploughing through resin (Thompson et al, 212). 2.2 Laboratory Testing Methods There are several tests that examine bolt performance and behaviour in a laboratory setting, including tests used to determine the material properties of the bolts. Because of the wide variety in types of rock bolt, certain tests may only be applicable to or relevant for a specific set of bolts Tension Test The American Society for Testing and Materials (ASTM) Standard F is ASTM s primary rock bolt testing standard, describing the chemical, mechanical and dimensional requirements for roof and rock bolts and accessories (ASTM F432, 213). It references and specifies a range of testing methods, the first of which is the tension test. ASTM F dictates that test bars for the manufacture of bolts and threaded and threaded slotted bars (ASTM F432, 213) must be in accordance with ASTM A37-12, which describes procedures and definitions for the mechanical testing of steels, stainless steels, and related alloys (ASTM A37, 212). A full-sized or machined sample of known dimensions including cross-sectional area with or without a reduced test section (Figure 2.2) is gripped in two locations (designated as the grip sections ), and tensioned at a constant strain, stress or crosshead travel rate.

20 Chapter 2. Testing of Rock Bolts 6 The yield point and the yield strength of the test sample may be determined using a variety of methods during the test or post-processing. Subsequent to the yield of the bolt, its tensile strength is determined, as well as elongation, elongation at fracture and reduction of area. Figure 2.2: Tensile test specimen with a reduced section (ASTM A37, 212) Tensile Test of Friction Stabilizers This particular test is a recent addition and is applicable to what ASTM refers to as friction stabilizers, synonymous with FRS. An FRS bolt is installed in a test plate made flush against the head of the bolt while maintaining clearance between the circumference of the plate hole and the FRS (i.e. the bolt is loose in the plate). The other end of the bolt is gripped and plugged so as to be held in place (Figure 2.3). The bolt is then tensioned to the minimum ultimate load specified by the standard for the applicable nominal bolt diameter without failure of the plate or any visual destruction of the bolt head (ASTM F432, 213). Figure 2.3: Apparatus for a tensile test on an FRS (ASTM F432, 213) Wedge Tension Test The wedge tension test, described in ASTM Standard F66-13, is to be used to test bolts and threaded bars in conjunction with the nut intended for practical use. A 1 wedge is placed under the bolt head (Figure 2.4), and the unheaded end of the bolt gripped. The bolt is then tensioned until failure. The

21 Chapter 2. Testing of Rock Bolts 7 yield strength is obtained using the drop of the beam method (ASTM D4435, 213), which designates the yield strength as the point where either a drop or a plateau in load is registered while deformation of the bolt continues. Tensioning progresses in order to determine the wedge tensile strength of the sample, while also demonstrating head quality and ductility of the product (ASTM F66, 213). Figure 2.4: Bolt head configuration for a wedge tension test (ASTM F66, 213) Bend Test The bend test as described by ASTM Standard F is only applicable to notched bendable bolts. Bolts are subject to one bending cycle, i.e. are bent 9 in the area of the reduced cross-section with respect to the original position, and then bent back. The test is passed if there is no obvious visible evidence of fracturing. The same bolt is subsequently subject to a tensile test across the section that was bent, and must attain the minimum yield load specified in ASTM Standard A corresponding to the steel grade and bar diameter, plus 6 lbf (26.7 kn). After fracture of the bolt in the tensile test (which must occur above a load of 23 lbf; 12.3 kn), there should be no evidence of prior fracture induced by the bend test (ASTM F432, 213) Tests of Expandable Rock Bolts ASTM Standard F describes three tests that must be performed on expandable bolts. They are as follows. Expansion Test The expansion test entails fully expanding the product at the manufacturer s recommended pressure. Ferrule Test The tensile strength of the ferrule-to-bolt weld must be found to meet or exceed the bolt profile breaking load. This is verified by gripping the ferrule on one end of the bolt and the other, plugged end of the

22 Chapter 2. Testing of Rock Bolts 8 bolt itself (Figure 2.5), and applying tension. The specific test methodology must be supplied to the customer. Figure 2.5: Two alternative configurations for the ferrule test (ASTM F432, 213) Expandable Bolt Mechanical Property Test A fully inflated bolt (it is permissible to use the bolt subject to the expansion test) is tensioned in order to find the yield and ultimate tensile loads. These must reach the specifications laid out in ASTM Standard A Laboratory Rock Anchor Capacity Pull Test The laboratory rock anchor capacity pull test is one of two tests described in ASTM Standard D It may be applied to elements with mechanical, resin or other similar anchor systems (ASTM D741, 28). A bolt is installed in a steel tube, using standard installation procedures so far as they can be reasonably replicated, i.e. using the torque, spin time, etc. specified by the supplier. Two potentiometers are used to measure both the bolt end displacement and the plate displacement (Figure 2.6). The bolt is then tensioned until failure using a hydraulic ram, while load and displacement of the bolt head are recorded. The working and ultimate load of the bolt are found from the resulting load/displacement graph, and the energy absorbed by the bolt is calculated (ASTM D741, 28). Figure 2.6: Apparatus for the Laboratory Rock Anchor Capacity Pull Test (ASTM D741, 28)

23 Chapter 2. Testing of Rock Bolts Laboratory Drop Test The drop test is the second test described in ASTM Standard D It is to be applied to similar bolts as those described in the laboratory pull test, as the installation in the steel tube is nearly identical although for this test an impact plate and load cells are incorporated into the head assembly. The tube is installed in a drop frame (Figure 2.7), where an electromagnet raises a known mass to a predetermined height before its power supply is cut, dropping the weight. The energy input into the system is measured by calculating the velocity of the mass as it lands on the impact plate, displacing the bolt head through bolt-dependent displacement mechanisms. It is permissible to perform the experiment multiple times on the same bolt to investigate the effect of cumulative energy input (ASTM D741, 28). Figure 2.7: Apparatus for the Laboratory Drop Test (ASTM D741, 28) 2.3 In Situ Pull Test ASTM Standard D describes the only ASTM-standardized in situ test for rock bolts; the pull test. Additionally, the International Society for Rock Mechanics provides suggested methods for pull testing (ISRM, 1981). ASTM D : Standard Test Method for Rock Bolt Anchor Pull Test ASTM D dictates that when developing a rock bolt pull test program, the program should reflect all possible installations of bolt, such as rock unit, orientation relative to any anisotropy present in the rock mass, anchor configurations, etc. 1 to 12 tests are recommended for each set of variables. The apparatus described in the standard (Figure 2.8) includes a loading system with sufficient capacity to fail the bolt and induce at least 5 mm of displacement. Additionally, it should be compatible with any rock surface condition and not deviate more than 5 from the bolt axis. Either a load cell on the bolt or a pressure gauge on the ram is to be used to measure applied load, and displacement is to

24 Chapter 2. Testing of Rock Bolts 1 be measured to an accuracy of at least.25 mm and resolution of.13 mm. The displacement transducer must be supported from a point no closer than.9 m from the loading system if attached to the same rock face, or alternatively from the floor. A borehole diameter measuring gauge with a resolution of at least.25 mm is to be used to measure the drilled hole diameter at the location of the anchor, and a thermometer used to record temperature at a resolution of.5 C. The bolt, anchor and drilling equipment are to be used per typical operational procedure, although bolts are not to be tensioned. Figure 2.8: Apparatus for a rock bolt anchor pull test (ASTM D4435, 213) At least half of the bolts tested are to have three loading/unloading cycles performed to 1 /4, 1 /2 and 3/4 of the estimated failure load. For these cycles, 1 equal load increments are to be used, and load is to be applied smoothly and rapidly. These bolts are then pulled until failure in either the same increments as the previous cycle, or increments of 2.2 kn, whichever is less. Non-cycled bolts are tested until failure in either 2 equal load increments, or increments of 2.2 kn, whichever is less. Displacement is recorded to the nearest.13 mm after each loading increment. Bolts are to be pulled 12.7 mm beyond the failure displacement, defined either by the peak load or 12.7 mm of recorded displacement. After failure occurs, load is to be recorded every 1 mm of displacement, as opposed to recording at load intervals (ASTM D4435, 213). As per ASTM D , the stress on the bolt is to be calculated, as well as the corrected bolt head displacement (total displacement discounting elastic deformation of the bolt). Working and ultimate capacities (Figure 2.9) are to be interpreted from the resulting load displacement graph. Working capacity is defined as the load on the anchor system at which significantly increasing displacement begins, and ultimate capacity as the maximum load sustained by the anchor system (ASTM D4435, 213).

25 Chapter 2. Testing of Rock Bolts 11 Figure 2.9: Conceptual load versus bolt head deflection curve for a rock bolt pull test (ASTM D4435, 213) Suggested Methods for Determining the Strength of a Rockbolt Anchor (Pull Test) Published by the International Society for Rock Mechanics (ISRM), the scope of this suggested method is to measure the short-term strength of a rockbolt anchor installed under field conditions (ISRM, 1981). It is suggested that at least 5 tests are performed for a given set of rock and installation conditions. The standard rock bolt assembly is to be used for the pull test, along with a hydraulic jack and handpump with a travel of at least 5 mm and a method of measuring load with an accuracy of 2% of the maximum load reached in the test. Displacement should be measured by a device (a dial gauge is suggested) accurate to.5 mm, to be mounted on firm rock (ISRM, 1981). Figure 2.1 shows a schematic of the suggested apparatus. The bolt should be installed in conditions representative of those commonly encountered on site with standard installation, and if possible the bolt should not be tensioned. Dimensions of the drill hole, bolt, and anchor should be measured, and the type and condition of the rock should be noted. It is suggested that the pre-load (the load at which displacement is recorded) should not exceed 5 kn. Further load should be applied in increments of either 5 kn or 5 mm, whichever is less, at a rate of 5-1 kn/min. Load and displacement readings should be allowed to stabilize before recording, and the time taken for this stabilization noted. The bolts should be pulled until 4 mm of displacement, yield, or failure, whichever occurs first. Total displacement should calculated, along with the anchor strength (defined as the maximum load before yield or failure of the anchor). If the tendon yields or fails before the anchor, the failure/yield load is recorded as the minimum anchor strength. Elastic elongation should be calculated and contrasted with bolt behaviour, and the effect of cement cure time examined (ISRM, 1981).

26 Chapter 2. Testing of Rock Bolts 12 Figure 2.1: Apparatus for a pull test (ISRM, 1981) 2.4 Discussion of Testing Methods Most of the laboratory tests outlined are largely dependent on the mechanical properties and dimensions of the rock bolts tested. While these tests may be highly controlled and are likely quite repeatable (depending on materials and manufacturing processes used), they ignore vital aspects of rock mass reinforcement. As such, they only provide a measure of rock bolt performance under controlled conditions that may not represent those found in situ. A rock bolt system is composed not only of the bolt-nutplate assembly, but also the surrounding rock mass and, for some types of bolt, a mechanical or chemical anchor. Each of these, and their interactions with one another, may be affected by how the bolt is installed in operational scenarios. As such, to properly evaluate support element performance, more realistic analogues of bolt installation conditions are necessary. A laboratory pull test has certain advantages over a field test; more control may be exerted over certain parameters, and thus it may act as a better test to compare the relative performance of some bolts or grouts for that specific apparatus. Due to the controlled environment, more advanced instrumentation is significantly easier to implement than in the field. However, the laboratory pull test as described by ASTM D741-8 has limited application to field results, depending on bolt type. Bolts that are in direct contact with the rock mass (friction and expandable bolts) may outperform their in situ counterparts; as both the bolt and the installation tube have quite uniform diameters, the load distribution along the length of the bolt may not be realistically replicated. The strength of the anchor/rock or grout/rock interface may also not be accurately represented for other types of bolt. This could and has been addressed to some extent by going beyond the scope of the standard and installing bolts in blocks of concrete or rock in the lab (Li et al, 214). However, even this test is missing two crucial aspects of support element systems; the specific rock mass the bolts are to reinforce, and installation of the bolts as performed on a mine site. Although the in situ pull test incorporates these two missing aspects of testing into an evaluation of performance, this is not to say that it is a perfect descriptor of the efficacy of a reinforcement element. There are problems with this test on a conceptual, as well as practical, level. Bolts are only loaded axially at their heads, while it is unrealistic to assume that bolts underground will only be subject to

27 Chapter 2. Testing of Rock Bolts 13 similarly oriented and positioned forces. Shear and rotational forces may occur at one or multiple points along the bolt. The working environment is a difficult one, limiting instrumentation due to delicacy or set up time. With this lack of complete instrumentation, there also comes uncertainty with regard to how the bolt physically fails or yields, or the displacement mechanisms recorded ASTM D concedes that interpretation of the [load-deflection] curve often requires some engineering judgement. 2.5 Summary Several rock bolt types are used in underground mines. Each of these bolts may be subject to variations of several tests described by ASTM, designed to quantify various aspects of behaviour and performance. The in situ pull test best incorporates certain aspects of operational bolt performance, namely the bolt installation process and the rock mass in which the bolt will be installed. Despite the shortcomings of the pull test, as a site-specific evaluation of the performance of a rock bolt, it is a widely applicable and relatively inexpensive test influenced by a greater number of relevant factors than any of the laboratory tests discussed. This thesis focused on the behaviour of rock bolts subjected to pull tests. Chapter 3 provides a description of the mines that provided the the results of the pull tests, and the specific bolts tested. This information was used to construct a database of pull tests performed at these mine sites within a specific time frame.

28 Chapter 3 Composition of the Database This chapter provides a brief description of the regional geology of the Sudbury Basin, and of the mines participating in the development of the pull test database. The latter part of this chapter provides a description of the specific bolts tested in the database, and compares the database assembled with previous work. 3.1 Pull Test Setting Pull test data was collected from six copper-nickel mines around the Sudbury Basin, all owned and operated by Vale: Coleman, Copper Cliff, Creighton, Garson, Stobie and Totten (shown in Figure 3.1). The operations and their geological setting are described herein. Milnet N Z km AZ Nickel Offset North Range Shaft Whistle Sudbury Igneaous Complex Trillabelle Collins Strathcona Coleman Longvac Fecunis Fecunis Lake Levack Levack West Boundary Hardy Windy Lw Granophyre Quartzxrich gabbro Norite Sublayer L w Sultana Chicago Worthington Totten L w Proterozoic Victoria Fraser McCreedy Chelmsford L w Clarabelle Creighton LEGEND Chemsford Formation Onwatin Formation Onaping Formation CreightonI Murray granites L w Murray Capreol L w Little Stobie Garson Stobie Frood SUDBURY Ramsey Lake Copper Cliff North Copper Cliff South Quartzite Norduna East Falconbridge Falconbridge GreywackeI volcanic rocks Archean Granite gneiss and plutons South Range Shear Zone Fault Olivine diabase dykes NixCuxPGE deposits Figure 3.1: Map of Sudbury area showing locations of relevant mines (after Eckstrand & Hulbert, 27) 14

29 Chapter 3. Composition of the Database Regional Geology The Sudbury Basin is located at the contact between the South Province (South Range), composed of Proterozoic Huronian supracrustal lithologies including sedimentary and volcanic rocks, and the Superior Province (North Range), a collection of greenstone and metasedimentary belts, felsic plutons and gneissic terrains (Card et al., 1974). In 1964, Dietz proposed an interpretation of the Sudbury Structure as a meteor impact. He theorized that about 1.7 Ga, a meteorite 4 km in diameter struck Earth at 15 km/s, excavating a crater 3 miles (48.3 km) wide and 2 miles (3.2 km) deep. Magma welled up into the crater, resulting in the Sudbury Igneous Complex (SIC), which was then overlain by sediments. Several subsequent orogenies, particularly the Grenville Orogeny (about 1 Ga), deformed the basin into its present oval shape (Dietz, 1964). This model has been refined over the years, and it is now widely accepted that the basin is the result of an extraterrestrial impact at 1.85 Ga (Long, 29), resulting in a 2 km diameter crater after the collapse of the initial crater, and the SIC is largely a product of impact melt as opposed to a mantle-derived magma (Naldrett, 29). Ore deposits in the Sudbury Basin can be broadly divided into four categories. SIC-footwall contact deposits occur at the contact between the base of the SIC and the underlying rocks in the footwall, while footwall vein deposits are up to 7 m into the footwall from the SIC with magmatic or hydrothermal origin. Offset dike deposits are quartz diorite dikes with mineralized cores, and sheared deposits are SIC-footwall contact deposits typically in the east of the South Range that have been affected by ductile shear (Golightly, 29). Vale s Sudbury mines, described in Sections through 3.1.7, exploit all four types of mineral deposits between them. Table 3.1 shows the rock units present on the six mines and the UCS values used on site. In some cases the rock type is present on site, but no UCS values were readily available. Table 3.1: Intact rock UCS values in MPa by mine site Coleman Copper Cliff Creighton Garson Stobie Totten Ore Granite Gneiss 24 Mafic Gneiss 28 Granite Breccia 21 Sudbury Breccia 3 Olivene Diabase Dyke Quartz Diorite Metasediments Granite Norite Trap Dyke Amphibolite 2-24 Greenstone 29 Metabasalt 248 Conglomerate 223 Metagabbro 19.4

30 Chapter 3. Composition of the Database Coleman Mine Located in the North Range, 3 km north-west of Sudbury, Coleman Mine entered production in 197 and is now one of Vale s most productive mines. Four ore bodies (Main Ore Body, West Ore Body, 153 Ore Body and 17 Ore Body) are currently being exploited (Vale, 211), extracting a combined 1,515, metric tons of ore in 213 (Vale, 213). Several faults and an olivine-diabase dyke contribute to the challenge of managing seismicity on site. Up until 21, various configurations of cut-and-fill were the main mining method. Since then, bulk mining (mainly slot and slash) has become more prevalent for pillar recovery (Vale, 211). Figure 3.2 shows a longitudinal section of the mine. The 17 orebody extends below 57 feet (174 m), and as a result of high stresses and certain geologic structures such as the Lunchroom Fault, seismicity is a concern at Coleman. 145 pull tests were collected from Coleman, performed between September 21 and March OB MOB 17 OB Figure 3.2: Longitudinal section of Coleman Mine (Morissette et al, 214) Copper Cliff Mine Copper Cliff Mine is located on the Copper Cliff Offset, a quartz diorite dike in the South Range extending from the SIC to a system of faults to the south. Clarabelle pit entered production in 196, with a shaft coming online at Copper Cliff North in 1968, and Copper Cliff South in 1969 (Cochrane, 1984). The mine consists of numerous steeply dipping, pipe-like ore bodies (Morissette et al, 214) mined with methods including blasthole, vertical retreat, slot and slash and uppers retreat (Vale, 21). A cross fault, olivine-diabase dyke and trap dikes all act as sources for seismic activity across several levels of the mine (Morissette et al, 214). 164 pull tests were obtained from Copper Cliff Mine between October, 28 and March, 215. Figure 3.3 shows a longitudinal section of Copper Cliff Mine.

31 Chapter 3. Composition of the Database 17 Figure 3.3: Longitudinal section of Copper Cliff Mine (Chinnasane et al, 214) Creighton Mine Creighton Mine is one of the oldest operating mines in Ontario, having entered production in 191 (Vale, 214). It is located on the outer rim of the South Range. Currently, the majority of production comes from the 74 level (2255 m) and below, with most of the of the shallower resource already mined out. Significant seismic activity occurs on site as a result of its great depth and structures, including shear zones and faults. Several orebodies are mined in tandem, all steeply dipping. Mineralization occurs at the contact between the norite hanging wall and the granite-gabbro footwall, and is exploited using the slot-and-slash mining method (Morissette et al, 214). 294 pull tests were collected from Creighton Mine, performed between January 23 and January 215. Figure 3.4 shows a longitudinal section of Creighton Mine. SE 7LShaft 3LShaft 5LShaft 11LShaft 9LShaft Surface NW 1LL 35Lm 3 Shaft Ramp Footwall Orebodies Orebodies Hangingwall Orebodies GraniteIGabbro Complex 38LL 1158Lm Sudbury Igneous Complex Shear Zone 6 Shaft 58LL 1768Lm 8 Shaft 461Orebody 7LL 768LL 82LL 2134Lm 2341Lm 25Lm 2m 2 N 8 N 1LL 348Lm Figure 3.4: Longitudinal section of Creighton Mine (Snelling et al, 213)

32 Chapter 3. Composition of the Database Garson Mine Garson Mine consists of two operations; the surface ramp, and the shaft-accessed main mine which extends down 52 feet. Four main orebodies are present, coincident with four shear zones (Mukwakwami et al, 213). Slot and slash is used to mine the two ore bodies at depth (1 Shear and 4 Shear), both intersected by an olivene diabase dyke which defines east and west segments for each. The two orebodies have strike lengths of around 6 m, and dip at about 7 (Abdellah et al, 214). The shear zones, dykes and faults combined with the depth of mining result in seismic activity in the lower parts of the mine. 119 tests were obtained from Garson, performed between June 211 and March Stobie Mine The Frood-Stobie deposit is located on the Frood-Stobie concentric offset dike hosted in Sudbury Breccia just north of the city of Sudbury. Although the Frood-Stobe dike is concentric while the dikes hosting Totten and Copper Cliff mines are radial, they are comparable insofar as they are steeply dipping and composed of quartz diorite (Roussel et a., 23). Vale suspended operations at the Frood section of the mine in 212 (Carmichael, 212), although Stobie remains a significant producer at present. 15 pull tests were collected from Stobie Mine, performed between December 211 and December 214. Figure 3.5 shows a long section of Stobie Mine. Figure 3.5: Longitudinal section of Stobie Mine Totten Mine Totten Mine is one of several mines situated on the Worthington Offset, and is the most recent Vale operation to come online in Sudbury. After 7 years of development, Vale produced 64, tons of ore in 213 and will ramp up to full production in 216 (Vale, 215). The sulphide ore is found in an offset

33 Chapter 3. Composition of the Database 19 quartz diorite dike dipping at 8, the geological model of which is often compared to that of the Copper Cliff Offset (Lightfoot & Farrow, 22). 113 tests from between April 21 and December 214 were obtained. Figure 3.6: Longitudinal section of Totten Mine (After Sudbury Platinum Corporation, 215) 3.2 Reinforcement Elements in the Pull Test Database This section describes the bolts that constituted the pull test database. The bolts are briefly described, and manufacturer-provided specifications for bolt performance are outlined Friction Rock Stabilizers The FRS was developed by Scott (1977) as a reinforcement element that provides resistance to pull along its length by using friction generated between the element and the surrounding rock. They were originally branded as Split Sets by Ingersoll-Rand, but with the expiration of the patent they have become available from other suppliers under various names. This list of names includes Friction Set, Friction-Lok, Friction Stabilizer, Friction Bolt, as well as Split Set among others. As such a wide range of brand names exist for similar bolts, this class of bolt is referred to collectively as friction rock stabilizers (FRS) for the purposes of this thesis to avoid confusion. The FRS consists of a tube of steel with a gap, or split, along its length such that the bolt may compress radially when installed in a hole with a diameter smaller than that of the bolt. Resulting outward pressure on the surrounding rock mass generates frictional resistance to pull out. These rock bolts have a tapered toe-end to facilitate insertion into a drill hole, and generally have either a crimped head or, more recently, a ring welded on to the head to hold a plate at the end of the bolt. Over the course of the period examined, Vale had two primary suppliers of FRSs for testing. They will be referred to herein as Supplier A and Supplier B, supplying FRS A and FRS B respectively. Limited testing was also performed on a third supplier s (Supplier C) FRS bolts.

34 Chapter 3. Composition of the Database 2 FRS A The FRS A (Figure 3.7) is manufactured in 4 nominal diameters: 33, 35, 39 and 46 mm (FA33, FA35, FA39 and FA46 respectively), although the 33 mm variant was neither used nor tested at Vale s Sudbury operations during the time frame for which pull test data was collected. After the tube is electrolytically galvanized, a ring is welded to the head of the bolt, and the weld and ring are painted over with zincbased paint to prevent corrosion. Bolts manufactured in this sequence are referred to as pre-galvanized (i.e. the bolt is galvanized before assembly; Lynn Mainville-Beach, June 214). Length-normalized load capacities are calculated with an anchorage length equal to the bolt length minus 6 (15 cm). This is the specified anchorage length used for such calculations in Supplier A s pull test reports, and accounts for the taper length and a portion of the bolt that is not fully inserted into the rock mass so that the pull testing apparatus may be attached. Figure 3.7: FRS A schematic (Courtesy of Supplier A) FRS B The FRS B (Figure 3.8) is manufactured in the same 4 nominal diameters as FRS A bolts (FB33, FB35, FB39 and FB46), although as with the FRS A, the 33 mm variant was not used or tested by Vale s Sudbury operations. The bolts were manufactured with a crimped head until 212, and are now manufactured with a welded ring head. Bolts are post-galvanized; the galvanization is performed after the ring head is welded on to the bolt (Lamothe, June 214). Length-normalized load capacities are once again calculated using a length 6 (15 cm) less than the total bolt length. Figure 3.8: FRS B schematic (Courtesy of Supplier B) Table 3.2 compares what is referred to as the minimum breaking capacity by Supplier A and the minimum ultimate tensile strength by Supplier B of the FRS A and FRS B, as well as the initial capacity claimed by the suppliers, and the bit sizes suggested for installation. The FRS Bs have slightly larger minimum ultimate tensile strengths than their FRS A equivalents, although the recommended initial load ranges are identical. There is also variation in the bit size recommendations; Supplier B recommends a larger minimum diameter for the 33 mm bolt, smaller minimum and maximum diameters for the 35 mm bolt, and a larger maximum diameter for the 46 mm bolt.

35 Chapter 3. Composition of the Database 21 Table 3.2: Comparison of FRS supplier specifications (Courtesy of Suppliers A and B) Minimum Ultimate Tensile Strength Initial Load Capacity Nominal Bit Size FA33 71 kn 27 to 54 kn 3 to 33 mm FB33 89 kn 27 to 54 kn 31 to 33 mm FA35 71 kn 27 to 54 kn 32 to 35 mm FB35 89 kn 27 to 54 kn 31.8 to 33.3 mm FA39 89 kn 27 to 54 kn 35 to 38 mm FB39 12 kn 27 to 54 kn 35 to 38 mm FA kn 54 to 89 kn 41 to 44 mm FB kn 54 to 89 kn 41 to 45 mm Rebar Rock Bolts Rebar rock bolts are partially threaded steel bolts that are grouted in a setting medium upon installation (in the case of Vale s Sudbury operations, resin is used). Rebar supplied by Suppliers A and B, as well as a third supplier, Supplier C, were tested. Each bolt is installed in the corresponding supplier s brand of resin. The rebar are usually tensioned by tightening the nut on the bar after the fast-setting resin at the base of the hole has set, and before the slower setting resin along the rest of the column has set. The large majority of the bolts tested were 2 mm in diameter, with a limited number of tests on 22 mm bolts. Both Suppliers A and C manufacture their rebar using Grade 6 steel (minimum yield strength of 42 MPa; ASTM A615, 215), while Supplier B uses grade 4W steel (minimum yield strength of 4 MPa; CSA G3.18, 29). Figure 3.9 shows schematics of the three rebar, and Table 3.3 compares the strengths of 2 mm bolts. Figure 3.9: Schematics of rebar manufactured by Supplier C (top; courtesy of Supplier C), Supplier A (middle; courtesy of Supplier A) and Supplier B (bottom; courtesy of Supplier B) Table 3.3: Steel properties for a 2 mm threaded rebar (Courtesy of Suppliers A, B and C) Gr. 6 Gr. 4 W Minimum Yield Strength 89 kn 86 kn Minimum Ultimate Tensile Strength 134 kn 116 kn As shown in Table 3.3, the Grade 6 rebar (Suppliers A and C) has a marginally higher yield strength as well as higher tensile strength than Supplier B s Grade 4 W rebar. It should also be noted that each bolt type has a different surface pattern, which may influence the interaction between the bolt and grout during and after mixing. All three of the suppliers rebar are manufactured with 1 UNC threads.

36 Chapter 3. Composition of the Database Modified Cone Bolt The Modified Cone Bolt (Figure 3.1) is a yielding reinforcement element, designed to absorb kinetic energy in the event of a rock burst. It consists of a smooth bar with a conical head that ploughs through the resin column as displacement of the plate at the surface of the rock mass occurs. The current configuration of Conebolt supplied by Mansour, the MCB33, derives its name from the recommended drill bit size for bolt installation (33 mm). Limited testing was also performed on the discontinued MCB38. The Modified Cone Bolt has a paddle on the end of the conical head to improve resin mixing as the bolt is inserted into the drill hole. For debonding and protection against corrosion, the smooth bar of an MCB33 has a plastic coating, or sleeve, while greasing the smooth bar of an MCB38 was previously performed to limit bonding between the tendon and the resin (Cai et al, 21). The MCB33 is manufactured using a modified C155 carbon steel. Table 3.4 shows the minimum and typical yield and ultimate loads at the thread of the bolt. Figure 3.1: MCB33 (Courtesy of Mansour) Table 3.4: MCB33 mechanical properties (Courtesy of Mansour) Minimum Typical Yield Strength 98.5 kn kn Ultimate Tensile Strength kn kn D-Bolt The D-Bolt (Figure 3.11) is a yielding reinforcement element supplied by Normet. It consists of segments of smooth bar punctuated by anchor sections that are composed of perpendicularly oriented flattened paddles. The D-Bolt is designed such that load is distributed evenly along the smooth steel section, allowing for relatively large deformations and thus capacity to absorb energy. Due to the strength of the steel and relatively stiff bolt response, the D-Bolt may also hypothetically be used as a static support element (Li, 211). The D-Bolt is manufactured with a smooth bar diameter of either 2 mm or 22 mm, with pull tests on both diameters of bolt present in the database. Figure 3.11: D-Bolt schematic (Normet, 214)

37 Chapter 3. Composition of the Database 23 Mansour manufacture a similar, 2.5 mm diameter bolt: the VersaBolt (or DS-Bolt). Figure 5.19 shows one possible configuration of the Versabolt. Table 3.5 shows the yield and ultimate tensile loads of the D-Bolt and VersaBolt. Figure 3.12: VersaBolt schematic (Courtesy of Mansour) Table 3.5: VersaBolt and D-Bolt mechanical properties (Courtesy of Mansour, Normet) VersaBolt D-Bolt (2 mm) D-Bolt (22 mm) Typical yield strength 138 kn 15 kn 19 kn Typical ultimate tensile strength 192 kn 21 kn 25 kn Expandable Bolts Expandable bolts are reinforcement elements that are inflated once placed in a drill hole. They consist of a folded tube sealed at one end, with a pump adapter on the other. Water is pumped into the bolt to a specified pressure, expanding the bolt and contouring it to the walls of the drill hole. The original brand of expandable bolt was the Swellex, supplied by Atlas Copco. Upon the expiration of the patent, similar alternatives from different suppliers came onto the market. The Swellex line was the most prevalent expandable bolt in the database, while significantly fewer tests were performed on DSI s Omega bolt and Jennmar s Python bolt. Figure 3.13 shows a Python bolt, as well as cross-sections before and after inflation. 1 2 Figure 3.13: Schematic of a Python bolt, and cross sections before (1) and after (2) inflation (Courtesy of Jennmar) Atlas Copco manufacture three different types of Swellex; the Premium (Pm) line for standard applications, the Manganese (Mn) line for large deformation circumstances, and the Spartan (Sp) line for low convergence, low energy scenarios (Atlas Copco, 212). DSI supply a Standard and a Plus line, which are equivalent to the Pm and Mn lines respectively, and Jennmar only have one line. Table

38 Chapter 3. Composition of the Database compares different bolt variants. Swellex Spartan bolts are omitted as they were not tested at Vale s operations, and there were no DSI or Jennmar equivalent at the time of writing. Note that breaking load and elongation values are expressed as minimums for Swellex and Omega bolts, but as typical values for Python bolts. Table 3.6: Summary of Swellex, Omega and Python bolt mechanical properties (Atlas Copco, 212; Courtesy of Jennmar, DSI) Type Variant Breaking Load (kn) Elongation Material Thickness (mm) Swellex Pm % 2 Mn % 2 Omega 12 Tonnes Standard 12 1% 2 12 Tonnes Plus 115 2% 2 Python Standard 12 25% 2 Swellex Pm % 2 Mn % 2 Omega 16 Tonnes Plus 15 2% 2 Python Midi 16 25% 2 Swellex Pm % 3 Mn % 3 Omega 24 Tonnes Standard 24 1% 3 24 Tonnes Plus 22 2% 3 Python Super 24 25% 3 In addition to the different bolt variants supplied by Atlas Copco, the bolts may also be manufactured with a plastic coating, denoted by adding a prefix of Pc to the bolt name (e.g. PcPm12). Two types of plastic coating are used; a polyvinyl (PVC) coating is used exclusively on the Pm12 bolt, while a polyethylene (PE) powdered coating is used on other variants (Leung, 214). 3.3 Database In total, pull test data for at least 26 bolt configurations from 7 suppliers across 5 classes of bolt was collected Database Description The pull test database was constructed as a tool that may be used to investigate overall trends and behaviours of various rock bolts by analysing a large number of pull tests, as opposed to examining results within isolated testing campaigns. A wide variety of rock bolts were pull tested over the period of data collection, allowing the direct comparison of performance when installed and tested in broadly similar conditions. The large amount of empirical data collected made it possible to compare conceptual results with behaviour interpreted from in situ pull testing. This also allowed for the determination of input parameters for design, recognizing variation present in field data. The database was constructed using pull test reports issued by the personnel conducting the tests (usually the supplier of a certain bolt will perform the pull tests on that bolt) to the mines. While the content and quality of the reports varied between and within suppliers, they generally contained data on the installation of the bolt, an indication of ground conditions, and information about the bolt being

39 Chapter 3. Composition of the Database 25 tested. The results of testing take the form of a load displacement curve, a yield load (i.e. working capacity) and/or a maximum recorded load. The pull test reports were evaluated on an individual basis to ensure consistency in the data input into the database. For example, the working capacity was evaluated directly from the load displacement curve of a test; the yield value recorded by the author of the pull test report was not accepted without verification. Table 3.7 summarizes the number of pull tests performed on each configuration. Table 3.7: Number of pull tests by bolt type Class Name Manufacturer Configuration Number FA35 92 FRS A Supplier A FA39 3 FA46 52 FB35 1 FRS FRS B Supplier B FB39 11 FB FRS C Supplier C FC35 6 FC46 6 MD Bolt Sandvik N/A 1 TOTAL 555 Rebar A Supplier A 2 mm 52 Rebar B Supplier B 2 mm 54 Grouted Bolts Rebar C Supplier C 2 mm 18 (Static) Fibreglass Rebar FiReP N/A 1 Unspecified rebar 7 TOTAL 141 Pm12 57 Swellex Atlas Copco Pm24 36 Mn12 17 Expandable Bolts Mn24 2 Omega DSI 12t 4 24t 1 Python Jennmar 16t 5 TOTAL 14 Modified Cone Bolt MCB33 78 MCB38 8 D-Bolt Normet 2 mm 15 Yielding Bolts 22 mm 2 DS-Bolt Mansour 2.5 mm 4 Yield-Lok Jennmar N/A 1 TOTAL 135 Other Eyebolts in FA39 Supplier A N/A 11 Mechanical Bolt in FA39 N/A 3 TOTAL 14 TOTAL 985 Some of the bolt types shown in Table 3.7 could have been further subdivided, however inconsistent recording of precise bolt configuration made it difficult to distinguish between these subdivisions. For example, many FRSs were noted to be galvanized. However, in a large number of cases, it was not noted whether bolts were galvanized, and there was a small sample size of bolt explicitly note as nongalvanized. Given the size of the non-galvanized sample, it proved to be impossible to separate the

40 Chapter 3. Composition of the Database 26 influence of galvanization from other factors. A similar situation exists with Swellex bolts being uncoated, versus having various plastic coatings. In these cases, bolts are grouped and these distinctions are not incorporated into the analysis Comparison to Other Pull Test Databases Traditionally, pull tests are performed for specific purposes, such as corroborating the supplier s claims of load capacity for a bolt. There are exceptions, and similar pull test databases have been previously assembled for large-scale analysis. Tomory et al. (1998) built a database of over 9 pull tests on two diameters of Split Set (SS33 and SS39), and analysed their performance in terms of load capacity and how it related to various factors associated with their installation. Soni (2) collected 39 pull tests performed on Swellex bolts, and also attempted to correlate performance with installation parameters. The database assembled for the purposes of this thesis is different for several reasons. The first is that multiple types of bolt are investigated; both Tomory et al. (1998) and Soni (2) focus on one type of bolt (Split Sets and Swellex respectively). This database includes not only different types of bolt, but also similar bolts from different suppliers. While Tomory et al. (1998) compare two diameters of Split Set, the database assembled for this thesis has a significant amount of pull test data for three diameters of both FRS A and FRS B bolts. As all pull tests were performed for one company (Vale) in the Sudbury Basin (i.e. in relatively close proximity to one another), there is some degree of consistency in the products delivered to the mine sites, as well as installation, testing and reporting procedures. All tests were conducted in only 6 mines in similar geological settings (compared to over 5 mines in the analysis of Tomory et al; 1998, and mine sites across North America and Europe in the case of Soni; 2), lending further consistency to the testing conditions. Finally, a significant difference between the database assembled for this thesis and those of Tomory et al. (1998) and Soni (2) is the inclusion of displacement data. While the previous analyses were performed on the load capacities of the investigated rock bolts, this thesis seeks to also address bolt head displacement behaviour with load in order to develop guidelines of expected behaviours and input parameters for more advanced design methodologies Specific and General Limitations of the Pull Test Database Despite the large amount of data collected, there were inconsistencies and shortcomings that should be noted. Although collecting pull test data from six mines provides the undeniable advantage of large data volume, each mine has different equipment, equipment operators, ground control personnel, and ground conditions. All of these have the potential to add variability to the process of installing a bolt, and thus possibly its performance when subject to a pull test. On the other hand, with all six mines located in the same region, only lithologies present in and around the Sudbury Basin are represented in the database. In addition, as each supplier generally performs their own pull testing, different apparatuses and practices were used to conduct the pull testing, introducing another source of variability and possible bias. An omission of many pull test reports was a the use of a standardized method of rock mass characterization. While even one to two word descriptors such as good or heavily fractured were used, such terms are part of a non standard and very subjective terminology. While rock mass classification schemes such as RMR or Q often include qualitative measures, they are less subjective than such broad

41 Chapter 3. Composition of the Database 27 descriptors of quality. As such it is seen as a significant shortcoming that no standardized methodology of assessing, or even language describing, rock mass quality was available from the pull test reports or the mine sites themselves. As a result of Vale s Sudbury Operations motivations for performing pull tests (quality control, product interaction and investigation of new reinforcement elements), almost all pull tests were performed on bolts that had been installed hours or minutes before testing. As certain time related effects may affect performance, such as corrosion and exposure to vibrations resulting from seismicity and blasting, this is a significant area of study that is not addressed. Bolts that had been installed before the test date were usually randomly selected, with an unknown or unnoted installation date and no assessment of corrosion. Due to safety concerns, bolts were not loaded until failure during testing. While working capacity is a valuable design metric, an investigation into the full behaviour of the bolt until it reaches its ultimate capacity and/or failure could not be undertaken for most types of bolt in the database. It must be acknowledged that the database represents pull tests over a period of 4 years, or longer for some mines. In this time, the design of the reinforcement elements as well as pull testing procedures may evolve. For example, Supplier B transitioned from supplying FRS B with a crimped head to one with a steel ring, and the possibility of reducing the diameter of certain FRS A configurations was under investigation during this time frame. As such, the analysis performed is representative of the specimen population as a whole during this window, and not necessarily of current or future bolt performance. 3.4 Summary The six mines for which pull test data was compiled are all situated in or near the Sudbury Igneous Complex. A wide range of rock units are present as a result of the complex geology of the area, a challenge compounded by inconsistent geotechnical data logging between the mines. While the bolt types discussed have different operational principles behind them, the provided specifications between suppliers of the same bolt configuration generally agree with one another. Although suppliers do not provide an explicit value of strength to be used for the design of an excavation, they do often provide minimum and typical material strengths in terms of both yield and ultimate load. Databases of pull tests have been previously assembled, however the data collect for this thesis originates from 6 mines in relative proximity to one another. More types of rock bolt are included than in other databases, and bolt behaviour as well as load capacity is examined. Chapter 4 details the pull testing process in the campaigns that constitute the database and compares the methods used in practice with those suggested by ISRM or prescribed by ASTM.

42 Chapter 4 Review of Implemented Pull Test Methods Chapter 2 provided a description of the ASTM Standards for in situ pull testing of rock bolts, as well as the ISRM suggested methods for pull testing. This chapter provides a commentary on pull tests as implemented in the reported database. It is based on all pull test reports in the database, and witnessing a number of pull test campaigns performed by different suppliers at different mine sites. This chapter will also discuss the significance of any deviation from the ASTM standards and ISRM suggested methods. 4.1 Implementation of Pull Tests in Practice Standard procedure at Vale s Sudbury operations, as in most Ontario mines, is to have at least one representative from both the rock bolt supplier and the mine s ground control staff present at the test. As the pull tests that constitute the database were generally performed by the supplier of the bolt being tested, there are variations in the apparatuses and procedures used to perform the pull tests. The pull test reports generally have a generic procedure outlined therein, in which significant deviations may be recorded. The author was present to witness two campaigns of pull testing; MCB33s tested by Mansour at Garson Mine, and 22 mm D-Bolts tested by Normet at Creighton Mine. Figure 4.1 shows the apparatus used by Normet set up to test a 22 mm D-Bolt. A bolt extension, also known as a pulling rod, comprised of a thick threaded steel bar is attached to the bolt being pulled with a threaded adaptor. A loading frame is mounted and made flush with the rock, if necessary by inserting plates between the rock and the loading frame. This may also orient the apparatus to ensure near-axial application of load. A hydraulic ram is then mounted on the assembly, followed by a wing nut to secure the system and transmit load from the ram to the pulling rod and rock bolt. A Vernier calliper with an electronic display is affixed to the apparatus, which measures the relative displacement between the bolt extension and the loading frame (i.e. ram travel). 28

43 Chapter 4. Review of Implemented Pull Test Methods 29 Figure 4.1: Normet s pull test apparatus mounted on a 22 mm D-Bolt Load is applied by pressurizing the hydraulic ram with a hand pump, and measured from a calibrated pressure gauge. Measurements are made by reading displacement off of the Vernier calliper at predetermined load intervals (for the 22 mm D-Bolt, these intervals were generally one reading every two tons of load up to 1 tons, and one per ton above that, although this may vary with the type of bolt being tested). Testing is generally performed until the bolt reaches its working capacity, although tests may be stopped before this is reached. In contrast, an FRS is tested until it slips (i.e. its ultimate capacity), although displacement is not measured for the majority of pull tests on this type of bolt. While suppliers employ different apparatuses, this procedure was the one most widely used to contribute to the database, with manual application of load and data recording. In general, Supplier A, and on occasion other suppliers, used a more digitized apparatus which logs data electronically. A displacement sensor measures displacement between the pulling rod and the lower part of the hydraulic ram (essentially the same displacement measurement as that shown in Figure 4.1), and load is measured using a load cell. Recordings are made at a certain frequency, as opposed to intervals of load or displacement. Load may be applied with either a manual or an electric pump. 4.2 Deviations from ASTM Standard D and ISRM Suggested Methods for Rockbolt Testing ASTM D (Standard Test Method for Rock Bolt Anchor Pull Test) dictates with great detail the apparatus and procedure used for an in situ pull test. Although some aspects of ISRM s suggested

44 Chapter 4. Review of Implemented Pull Test Methods 3 methods are not as rigorous, it is still a thorough methodology. Differences between the pull test as practised at the mine sites discussed and the standard and suggested methods are described herein Deviations in Apparatus The loading systems and transducers of the various suppliers generally meet the standard and suggested methods. The loading systems are as described in ASTM D , and have travel ranges exceeding the specified 5mm, and the pressure gauges/electronic transducers have the specified resolution of 445 N (although it is not clear whether this includes the effects of friction and the like ; ASTM D4435, 213). However, the apparatus for the measurement of displacement is often quite different to what is specified. While a calliper mounted on the loading system is used in practice, the standard specifies a dial gauge (or another displacement transducer that may still comply with the standard) that is either supported independent of the rock face, or.9 m from the reaction frame if supported by the face (ASTM D4435, 213). ISRM also suggests a dial gauge, and specifies that it must be supported on a stable surface (ISRM, 1981). This means that fundamentally different displacements are being measured while ASTM and ISRM pull tests measure the displacement between the bolt head and a stable datum (i.e. an unaffected part of the excavation surface), the pull test as practised measures displacement that occurs between the pulling rod and the loading frame, or ram travel. This may result in the incorporation of displacements incurred on the surface of the excavation beneath the loading frame in addition to those attributed to the movement or deformation of the rock bolt. ASTM additionally mandates that a thermometer be used to measure temperature, a parameter that may be relevant for resin or cement grouts. No recording of temperature was made in any pull test report collected for the database Deviations in Procedure Both ASTM and ISRM require bolt installations representative of typical operational procedures. ASTM suggests 1 to 12 pull tests should be performed for every combination of factors (such as installation equipment, orientation relative to anisotropy, etc.), and ISRM suggests at least five. The size of the pull test campaigns varies in the database; some include three tests, some over 2. Generally, campaigns are performed on five bolts for a set of parameters, such as installation in ore versus waste. Both ASTM and ISRM also demand measurements of borehole dimensions (ASTM specifies diameter, ISRM diameter and length). Borehole diameter is occasionally recorded for the tests collected, although the methodology used to conduct the measurement is not always clear (i.e. the number and location of measurements in the drill hole). Similarly, the measurement of bolt diameter (in practice only performed for FRSs) may be a single value, or an average of several measurements along the bolt s length. ISRM also suggest drill hole straightness, cleanness, dryness and orientation be recorded, none of which are noted in the pull test reports. As is the case with the sections relevant to the apparatus set up, a portion of both ASTM and ISRM s procedural guidelines are concerned with ensuring that bolt installation is performed using standard operational practices. The principal motivation for regular testing at most mines is quality control, so typical installation of the bolts is performed. However, this may mean that certain aspects of the ASTM and ISRM procedures are not observed, such as cleaning the drill hole before the bolt is installed.

45 Chapter 4. Review of Implemented Pull Test Methods 31 Only ISRM s suggested methods refers to the implementation of a pre-load. A pre-load is the limited loading of the bolt before recording displacement data in order to tighten the apparatus mounted on the bolt. A 5 kn maximum is suggested, although the typical pre-load used in practice was 2 tons (17.8 kn) and could be as high as 4 tons (35.6 kn). While ASTM does not make reference to pre-load, it does describe quite a detailed and extensive loading procedure. For at least half of the bolts, three loading/unloading cycles in 1 equal increments are specified, loading the bolt up to 1 /4, 1 /2 and 3 /4 of the estimated failure load. This guideline did not appear to have been followed in any pull test campaign, and bolts were fully loaded in the first and only loading cycle. Unloading and reloading only occurred when significant movement of the apparatus resulted in a loss of load on the bolt. ASTM specifies that non-cycled bolts are to be loaded to failure in 2 equal increments or increments of 2.2 kn, whichever is less (ISRM suggests increments no greater than 5 kn), but in general increments of 1 ton (8.9 kn) were used when increments were recorded manually. There are also guidelines regarding failure of the bolt; ASTM specifies the bolt must be pulled to failure (defined by ASTM as the peak load, or a deflection of 12.7 mm), and then pulled an additional 12.7 mm, with load recorded every 1 mm. ISRM specifies testing to be performed until either 4 mm of displacement, yield or failure of the anchor. In general, pull tests were performed until the working capacity (i.e. the point at which a significant increase in the displacement observed per load increment; ASTM, 213c) of the element was surpassed. Peak load (ultimate capacity) or failure load was not generally found for bolts besides the FRS, as there are safety concerns associated with energy release when failing a bolt. There are several calculations mandated by ASTM which are either not performed, or are not included in the pull test report. These include stress, elastic deformation and corrected bolt head displacement. It should be recognised that these calculations are not straightforward for all bolt types; for example, elastic deformation in a grouted bolt is difficult to calculate unless one makes the assumption that all displacement before the bolt yields may be attributed to elastic deformation Practical Considerations in Pull Testing The ISRM suggested methods for pull testing rock bolts have not been updated since 1974, and in fact are only suggested methods, not requiring compliance. ASTM in itself has no part in requesting or enforcing compliance to its standards. Unless it is a contract requirement, the use of ASTM standards in not required. Pull testing is a time consuming process that poses logistical challenges and requires the co-ordination of multiple parties. It can interfere with operations, involves the transportation of a significant amount of equipment to and from the site of the tests, and can only be performed when at least one representative of the supplier, their test equipment, one member of the ground control staff, bolt installation equipment and its operator are available. Combining these challenges with a demanding working environment, pull testing is a slow process that may take weeks to organize. While minor improvements, such as the recording of temperature, could perhaps be made on the tests as performed relatively easily, ASTM s standard is difficult to fully implement on a routine basis. The testing campaigns collected were typically performed on 5 bolts, or on occasion 1. In all campaigns, one loading cycle with increments of at least 1 ton (if loaded manually) were performed on each bolt. Had the ASTM standard been followed, at least four times as many measurements at 5 lbf (2.2 kn)increments would have been made. Additionally, half of the bolts would have been loaded three times in addition to the full test. This would result in perhaps two to three bolts being tested in the same time frame, an unacceptably low number for a quality

46 Chapter 4. Review of Implemented Pull Test Methods 32 control investigation. ISRM s suggested methods are more forgiving, especially in terms of procedure, however are somewhat dated. They pre-date the introduction of FRSs, expandable bolts and yielding bolts to the industry, and as a result are focused on cement-grouted rebar and mechanical bolts, which must be seen as a significant shortcoming of the procedures described. 4.3 Pull Test Data A pull test measures two variables; load and displacement. From these measurements, certain properties of the bolt-anchor-rock mass system may be inferred or calculated, and act as quantifiers of bolt performance. The two types of metric this thesis primarily focuses on are working capacity (or ultimate capacity for an FRS) and measures of stiffness Working Capacity and the Measurement of Load While the ultimate capacity of the element is arguably the best reflection of reinforcement strength, pull tests in which the bolt is pulled until total failure are rarely conducted, as there are significant safety hazards associated with the amount of elastic energy released from a failing bolt. Working capacity is a safer and more easily achievable alternative for most bolts, representing a value at which relatively large, permanent displacements occur with little additional application of load. The use of working capacity as a design value over ultimate capacity is in part related to the objective of a ground support system. For such a system to be successful, the excavation must be stable, safe, and operational. Working capacity signifies the beginning of increasing displacement per unit load, which may be at least in part attributable to plastic deformation of some types of rock bolt. This increase in displacement could result in closure of the excavation, and/or unravelling. Both of these effects may result in conditions that interfere with regular operations and may require rehabilitation. In this case the support system may not be considered successful; even though individual elements have not necessarily failed, displacements resulting from loads exceeding the working capacities of the rock bolts in the system result in a non-operational excavation condition. How a reinforcement element may react to dynamic loading resulting from seismicity may also depend on whether its working capacity has been exceeded. Figure 4.2 shows the results of laboratory static (left) and dynamic (right) loading tests of a 22 mm D-Bolt. In both cases, the bolt failed. 3 3 Impact load Load (kn) Load (kn) Plate load Plate End Displacement (mm) Displacement (mm) Figure 4.2: Static laboratory pull test on a 22 mm D-Bolt 2.1 m in length (left; Doucet & Voyzelle, 212) and dynamic impact test on a 22 mm D-Bolt 1.5 m in length (right; Li & Doucet, 212)

47 Chapter 4. Review of Implemented Pull Test Methods 33 The results of static and dynamic tests cannot be directly compared, as the difference in loading rate results in the reinforcement element assuming different properties. For example, in Figure 4.2, impact testing results in a higher yield load and larger ultimate displacement than a static laboratory pull test on a longer 22 mm D-Bolt. However, by indirectly comparing the tests it can be demonstrated why working capacity may be used as a design value for static scenarios. The energy capacity of a rock bolt is equal to the area under the load displacement curve (Li et al, 214). The D-Bolt absorbs energy by distributing strain across a length between two anchors, allowing for larger deformations at equivalent loads than reinforcement elements continuously bonded to the rock mass, such as rebar, where strain is concentrated on a relatively short section of the element. If a D-Bolt bears a load less than its working capacity, most of the D-Bolt s energy capacity is conserved as little deformation has been incurred. However, if the working capacity is surpassed, the energy capacity of the bolt diminishes much more significantly as the magnitude of plastic deformation per unit load is greater than that of elastic deformation. As a result, if a D-Bolt (or other reinforcement elements that physically yield without slipping) is subject to static loads greater than its working capacity, it will have significantly less capacity to absorb energy under dynamic loading. In most pull test reports, the working capacity is referred to as yield for all bolt types. Yield strength is defined as the engineering stress at which, by convention, it is considered that plastic elongation of the material has commenced (ASTM E6, 29). Figure 4.3 shows two methods of determining yield strength according to ASTM Standard E6-9b. Figure 4.3: Methods of determining yield strength: halt of the pointer method (left) and offset method (right; ASTM E6, 29) In Figure 4.3, the yield strength (S y ; UYS and LYS are upper and lower yield strength respectively) are determined at either a perfectly plastic response to load (i.e. an increase in displacement without a corresponding increase in load), or at a predefined offset in strain from linear elastic behaviour. The resolution of recorded load values for the data collected is variable and in some cases quite poor. If load is recorded manually, increments of.5, 1 or even 2 tons (4.45, 8.9 or 17.8 kn) are used in different tests. Digital recording is frequency-based, resulting in variable load resolution which is a function of loading rate and measurement frequency. Neither of the methods of determining yield strength presented in Figure 4.3 are practical when applied to the data collected. Perfectly plastic behaviour is seldom observed since displacement is recorded in increments of load and defining a strain offset to determine yield when bolt behaviours may be non-linear and highly variable is not feasible. The working capacity was determined for the purposes of this thesis as the greatest load measured

48 Load Chapter 4. Review of Implemented Pull Test Methods 34 before a softening of the bolt response attributed to plastic deformation of the reinforcement element material or movement of the entire element is observed, as demonstrated in Figure 4.3. Working Capacity Displacement Figure 4.4: Determination of working capacity from a pull test The working capacity may be indicative of a lower bound of yield strength for some bolts, however some bolt types may reach their working or ultimate capacity before the bolt material itself actually yields. For example, an FRS will generally slip rather than yield or fail if pull testing is performed soon after installation, and its ultimate capacity will be defined by the maximum load sustained by the bolt without slipping. As working and ultimate capacity are generally applicable measures of bolt load capacity, they are the primary load metrics discussed in this thesis Recording Displacement During a Pull Test Elastic deformation of the bolt contributes to the bolt head displacement observed during a pull test. However, other sources of displacement may be included in the measurement, especially if using a displacement measuring system mounted directly on the rock bolt loading apparatus. These displacements may include compression or fracturing of the rock mass and surface support system, movement of the entire bolt relative to the rock mass, or shifting of the testing apparatus, among others. This is demonstrated in Figure 4.5. d b represents the displacement of the bolt head resulting from bolt deformation (as in Figure 4.5), movement of the bolt relative to the toe of the hole, or a combination of the two as load is applied. d s is the displacement of the excavation surface in direct contact with the pull test apparatus relative to the excavation surface unaffected by the testing, defined by the extent of the rock mass or any surface support, as elements such as mesh or liners may deform as load is applied. In the case of Figure 4.5, d s is a result of fracture compression and generation, and movement of the resulting blocks. d a is the apparent displacement, equal to the summation of d b and d s. This is the value that is measured by a pull test apparatus that measures total displacement induced by the apparatus (as in Figure 4.1), as opposed to the reaction of the bolt. The use of a dial gauge mounted independently of the face in which the bolt is installed per ASTM D and ISRM s suggested methods minimizes the effect of the rock mass response by measuring bolt head displacement relative to a stationary reference point, such as the floor or a stable portion of the wall (i.e. displacement d b ).

49 Chapter 4. Review of Implemented Pull Test Methods 35 d b = displacement attributable to bolt deformation or movement d s = displacement induced on the surface of the excavation d a = d b + d s = apparent displacement d a d s d b Figure 4.5: Measurement of displacement for a pull test on a generic reinforcement element The measurement of d a may not accurately reflect the performance of the bolt alone if the objective of testing is solely the investigation of the bolt response to load, however it could still be representative of certain loading scenarios. Figure 4.6 shows a representation of a generic rock reinforcement system (Thompson et al, 212). Figure 4.6: Generic reinforcement system (Thompson et al, 212) A pull test using an independently mounted dial gauge measures two of the interactions shown in Figure 4.6: element internal fixture, and internal fixture rock. By recording the displacement of the loading system as was practised in all cases used to build the pull test database, a third interaction is simulated. The pull test apparatus acts as an external fixture, and its interaction with the rock is recorded as well as the element-internal fixture and internal fixture-rock interactions. This results in a loading scenario analogous to the one presented in Figure 4.7. Figure 4.7 presents a simplified scenario in which a point-anchored bolt is loaded axially by a wedge. The excavation surface around the face plate of the bolt compresses as it would during a pull test, displacing by an amount d s relative to the rest of the wedge s surface. The bolt deforms, resulting in a displacement d b at the bolt head. The summation of these two displacements, d a, is the displacement undergone by the surface of the wedge relative to the surface of the stable excavation. In the case of a point anchored reinforcement element, the d a of a pull test may reflect the d a of such an axial loading scenario, depending on factors such as plate size relative to the footprint of the pull test apparatus on

50 Chapter 4. Review of Implemented Pull Test Methods 36 the surface of the excavation, use of surface support, wedge size, etc. For other types of reinforcement elements, the agreement depends on where along its length a bolt is loaded, how load attenuates to the head of the bolt (influencing d s ), and the length of bolt that will deform (influencing d b ). d a = d b + d s d b d s d a Figure 4.7: Measurement of displacement for axial loading of a point anchored rock bolt The rock mass is an integral part of the system, but its complexity makes its role relatively difficult to monitor during a pull test. Shearing of a bolt may occur across one or a series of joints, especially if the bolt is not tested immediately after installation, which may affect its behaviour. Loss of resin into fractures in the rock mass may impact the volume of effective resin and its mixing, thereby potentially affecting the performance of any resin-grouted bolt. The drillability of the rock influences the diameter of a drill hole relative to the size of a drill bit and could impact the performance of frictionally coupled bolts. The strength and degree of fracturing of the rock mass and surface support may affect the measured displacements, as shown in Figure 4.5. It follows that the behaviour of a rock bolt during a pull test is influenced by the condition and properties of the rock mass in which it is installed Rock Bolt Stiffness When designing a ground support system, very little deformation of an excavation may be desirable, in which case stiffer bolts are preferred. In other instances (such as squeezing ground conditions), a more ductile ground support system may be implemented to allow deformation of the rock mass while preventing reinforcement elements from surpassing their ultimate tensile strain capacity (Potvin & Hadjigeorgiou, 28). A conceptual example is provided by Brady & Brown (26), as the convergence confinement method of support system design (shown in Figure 4.8). Inward radial displacement occurs at a point in a tunnel as a result of stress in the rock mass and a diminishing support effect provided by the face as it advances away from the point in question. In order to mitigate radial displacement, a ground support system is installed, which provides a certain support pressure. An equilibrium is reached when the ground reaction curve and the support reaction curve intersect. As more displacement occurs, less support pressure is generally required to establish an equilibrium. Figure 4.8 shows various support systems installed after different degrees of radial displacement. It can be seen that a stiff support system installed promptly when little radial displacement has occurred (System 3) must provide a large

51 Chapter 4. Review of Implemented Pull Test Methods 37 support pressure relative to either a more ductile system (System 4), or systems installed once more radial displacement has occurred (i.e. the face has advanced further; Systems 1, 2 and 5). Figure 4.8: Design of support systems using the ground reaction curve (Brady & Brown, 26) Stiffness is measured in terms of load per unit displacement, kn/mm. The calculation of stiffness is a way of investigating how stress attenuates down the bolt. Continuously coupled bolts will not have an even load distribution along their lengths, and their stiffness may describe how well they are bonded or coupled to the rock mass. It would appear that stiffness is reported in several forms in technical literature. The two principal measures discussed in this thesis are illustrated in Figure 4.9. Secant stiffness is the average relationship between displacement and load from initial loading up until the working capacity (i.e. a secant stiffness at the working capacity). Stiffness calculated along a linear portion of the load displacement relationship is the tangent stiffness (Bieniawski et al, 1978). The former will incorporate displacement effects not necessarily directly related to the behaviour of the rock bolt such as fracturing of the rock mass, compression of pre-existing fractures and movement of the loading rig among others. The tangent stiffness was used in an attempt to minimize the influence of these sources of displacement, and may potentially provide a better description of bolt performance.

52 Load (kn) Chapter 4. Review of Implemented Pull Test Methods SecantcStiffness TangentcStiffness Displacement (mm) Figure 4.9: Calculation of secant and tangent stiffness from a pull test This thesis also uses the metric unloading stiffness, which is calculated by examining the recovery of displacement (recorded as negative) when the bolt is unloaded, representing elastic recovery of the bolt. This should further reduce the effect of elements of the testing system extraneous to the rock bolt being tested. In some cases, the pull test report would explicitly state that unloading was examined, qualifying the data. For the tests in which unloading was recorded but was not an explicit objective of testing, two clearly visible recordings of load and displacement in the unloading phase must be visible (see Figure 4.1); as logging was frequency based, loading may have continued beyond the maximum displacement recorded, affecting the slope of the unloading line. Similarly, when a load of kn is reached, the pull test apparatus and rock mass are no longer held together in compression, and large displacements not associated with elastic recovery of the bolt may occur as the system loosens. Figure 4.1 shows a testing campaign in which only Test 6 provides acceptable data as measurements of load and displacement are recorded twice between the beginning and the end of unloading (circled in red). Figure 4.1: Pull test campaign results for partially encapsulated Rebar A from November 16 th, 212 at Coleman Mine (Mainville et al, 212)

53 Chapter 4. Review of Implemented Pull Test Methods 39 In addition to the metrics discussed, load and stiffness metrics applicable only to the Modified Cone Bolt are introduced in Chapter 5. As the MCB33 is designed to plough through the resin column (unlike any other bolt presented in this thesis), it has a unique behaviour for which the metrics of performance discussed are not satisfactory Limitations of the Metrics Measured in a Rock Bolt Pull Test It is acknowledged that the presented measures of performance are not free of bias. It may be difficult to establish whether a bolt is in fact yielding, or if displacement from another source is occurring near the expected working capacity. In cases that are particularly difficult to interpret, the working capacity is simply not included in the analysis. As a result, the data that is included in the analysis is of reasonable quality. The introduction of human judgement results in a degree of subjectivity of the working capacity value, and as a result the secant stiffness. Similarly, the linearity of portions of the load displacement response used to calculate the tangent stiffness was not quantified, but determined from the test. The linear section generally include at least 4 measurements or 3 segments (as in Figure 4.9) for manually recorded tests, and if two such sections were present the longer one would be recorded. It was more difficult to adhere to these data limitations on the instrumented tests as many of the stiffness calculations were based on a graphic of the load displacement behaviour, but linearity was only evaluated for segments greater than 2.5 tons. As a result of these limitations not all data was found to be compliant and used for further analysis. 4.4 Summary Pull tests as performed at Vale s Sudbury operations deviated from the standard laid out by ASTM and the methods suggested by ISRM. Perhaps the most significant difference in implemented apparatus and procedure is the method of measuring displacement. While both ASTM and ISRM suggest independently mounted displacement measurement systems, all suppliers use an instrument mounted on to the loading system itself. This results in measurement of the response of the rock mass surface immediately adjacent to the bolt tested, and does not solely represent how the bolt and its anchoring system perform. This thesis uses metrics of load capacity and stiffness to quantify bolt performance. Data was carefully interpreted in order to attribute observed displacements to certain effects or mechanisms. Despite the identified limitations, there is great value gained from pull tests conducted by experienced personnel that employ engineering judgement in collecting and interpreting the pull test data. In Chapter 5, the behaviour of rock bolts subject to a pull test is interpreted, and statistics on the performance metrics described in this chapter are presented for a series of bolts.

54 Chapter 5 Summary Statistics and Interpretation of Pull Test Data This chapter provides a description of the statistics and analyses used to characterise the behaviour of rock bolts when subjected to a pull test. It also explores the variation in performance of rock bolts as predicted by field data and a series of conceptual assumptions. 5.1 Summary Statistics and Statistical Techniques The construction of a large database of rock bolt pull tests in the Sudbury Basin has allowed for a statistical analysis on all test data collected for a given time frame. Contrarily, the entities responsible for testing usually treat individual campaigns independently of one another, and use pull tests as a tool to demonstrate compliance with supplier specifications. Various methods of characterizing a dataset and investigating correlations exist. In this section, the statistical measures used to summarize the performance of the rock bolts are described, as are the statistical tools used for analysis Summary Statistics The mean of a dataset is the arithmetical average, expressed as x for a data set {x 1, x 2,..., x i }, and is calculated as x = 1 n n x i i=1 where n is the number of samples. The sample standard deviation (s) is also routinely quantified to describe dispersion, calculated as n i=1 s x = (x i x) 2 n 1 Both of these metrics are used to calculate the coefficient of variation for the sample (ĉ v ) as a way of describing the size of the sample standard deviation relative to the mean of the data (Baecher & Christian, 23). This allows for the direct comparison of dispersion between variables with different magnitudes of mean. 4

55 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 41 ĉ v = s x x The shape of a distribution may be described by skewness and kurtosis. The measure of skewness elected for use is the adjusted Fisher-Pearson standardized moment coefficient (G 1 ). G 1 = n (n 1)(n 2) n ( x i x s Skewness is a measure of the tendency of the mean of a dataset to be on one side of the median as the result of the distribution exhibiting a tail. A positively (or right) skewed distribution has a mean greater than the median (and a tail pointing right), and will have a positive G 1. A normal distribution is described by G 1 = (Doane & Seward, 211). The second statistic used to describe distribution shape is a measure of excess kurtosis, referred to as kurtosis herein, calculated as ( ( ) ) n(n + 1) n 4 xi x 3(n 1)2 Kurtosis = (n 1)(n 2)(n 3) s (n 2)(n 3) i=1 Kurtosis is a measure of peakedness of the distribution, or more accurately a measure of weighting of the shoulders of the distribution relative to the centre. In the format shown, a normal distribution has a kurtosis value of ; kurtosis greater than indicates a peakedness of the distribution, and kurtosis below indicates a flatness or evenness of the distribution (Balanda & MacGillivray, 1988). Percentiles are calculated using the method recommended by NIST. For a set of N measurements sorted in increasing rank {Y 1, Y 2,..., Y N }, the value of the pth percentile (Y p ) is calculated by setting i=1 p(n + 1) = k + d where k is an integer value and d the remaining decimal. There are three ways Y p is subsequently calculated: 1. If < k < N, Y p = Y k + d(y k+1 Y k ) 2. If k =, Y p = Y 1 3. If k >= N, Y p = Y N ) 3 Note that the 5 th percentile is the median of the data (NIST-SEMATECH, 23) Statistical Techniques Various methods of analysis are used to investigate relationships in the data. The first of these is least-squares linear regression. For a bivariate analysis, least-squares linear regression is expressed as y i = β + β 1 x i + ɛ i where β 1 = S n xy i=1 = (x i x)(y i ȳ) S n xx i=1 (x i x) 2 and β = ȳ β 1 x

56 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 42 and ɛ i is a random error term. The output for a bivariate analysis is a line for which the mean distance between the regression line and all data points is minimized. The coefficient of determination (R 2 ) may be calculated for a linear regression, as R 2 = RSS n T SS = i=1 (ŷ i ȳ i ) 2 n i=1 (y i ȳ) 2 where ŷ i is the fitted value for the term β + β 1 x i. The coefficient of determination assesses the level of correlation between y and x; R 2 = 1 describes a perfectly linear relationship, and R 2 = indicates no linear relationship is present (Fox, 28). Two methods are used to determine if a statistically significant difference in means of two or more sets of data exists, given the null hypothesis that the means of two populations are equal (µ 1 = µ 2 ). The first method, used to compare only two datasets, is the studentized t-test. Various configurations of this test exist, all of which assume distribution normality. The two used for the purposes of this thesis are the t-test for data sets Y 1 and Y 2 with equal variance, and the t-test for populations with assumed unequal variance. For the former, the t-statistic is calculated as t = Ȳ 1 Ȳ2 1 s Y1Y 2 n n 2 where s Y1Y 2 = (n 1 1)s 2 Y 1 + (n 2 1)s 2 Y 2 n 1 + n 2 2 Accounting for the degrees of freedom, calculated as v = n 1 + n 2 2 the t-statistic is compared to the t-distribution. A level of significance is chosen (5% is used in this thesis). If the t-statistic calculated is greater than the t-value that corresponds to the degrees of freedom and significance level chosen (t crit ), the null hypothesis is rejected, suggesting a significant difference between population means. In addition, a p-value is estimated. If t > t crit, the p-value will be less than the significance level chosen (NIST-SEMATECH, 23). If samples are assumed to have unequal variance, the same procedure is followed but the calculation of t and v are as follows: t = Ȳ 1 Ȳ2 s 2 1 /n 1 + s 2 2 /n 2 v = (s 2 1/n 1 + s 2 2/n 2 ) 2 (s 2 1 /n 1) 2 /(n 1 1) + (s 2 2 /n 2) 2 /(n 2 1) Single factor, one-way Analysis of Variance (ANOVA) is used to determine whether a significant difference in mean exists between three or more sets of data (although it may be used for two datasets as well). Table 5.1 shows the general format of an ANOVA table, comparing J groups, each containing I entries.

57 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 43 Table 5.1: General One-Way ANOVA table (NIST-SEMATECH,23) Source Sum of Squares DoF Mean Square F Factor SS F = J (ȳ i. ȳ.. ) 2 I 1 MSF = SS F /(I 1) MSF/MSE Residual SS E = (y ij ȳ i. ) 2 I(J 1) MSE = SS E /(I(J 1)) Total SST = (y ij ȳ.. ) 2 IJ 1 where and ȳ i. = 1 J ȳ.. = 1 IJ J y ij j=1 I J y ij i=1 j=1 Similar to the t-test, the F value calculated is compared to the F -distribution, accounting for degrees of freedom (DoF) and significance level. If F > F crit, the null hypothesis is rejected and a difference between means is suggested at a predetermined level of significance (NIST-SEMATECH, 23). Although many more sophisticated methods of data analysis exist, their application was limited due to the nature of the database. The database was assembled from independent pull test campaign reports that inconsistently record a variety of parameters. As a result, large gaps in data exist; for example, of 545 pull tests performed on various configurations of FRS, rock mass quality was recorded for only 126 of those tests. Such a data set is incompatible with some methods of analysis, such as the assembly of a generalized linear model. Other statistical analyses are suitable for an incomplete dataset, for example principal component analysis (PCA). However, due to the often limited frequency and widely varying combinations of parameters recorded, PCA was found to be an ineffective method of analysis for this database. As such, analysis is limited to the methods that have been discussed in this section. 5.2 Friction Rock Stabilizers There are six configurations of FRS discussed in this thesis; three diameters of FRS (35, 39 and 46 mm) from two suppliers (A and B) are compared. A discussion on theoretical and observed behaviour follows Theoretical Behaviour of a Frictional Rock Stabilizer Li & Stillborg (1999) present analytical models for various types of rock bolt, including frictionally coupled bolts (such as the FRS; Figure 5.1).

58 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 44 Figure 5.1: Shear stress distribution along a frictionally coupled bolt subject to axial load (Li & Stillborg, 1999) The interface between the rock mass and the rock bolt is assumed to have constant shear strength s. Once the strength of the interface is overcome, the bolt is said to have decoupled at that point and cannot resist further application of load. A decoupling front develops (x 2 ), beyond which the shear stress on the interface decreases following function τ b (x). Once the decoupling front has reached the end of the element, the entire bolt is free to slip (Li & Stillborg, 1999). Figure 5.2 shows how load attenuates down the bolt. Although the figure presented is modelled off of a Swellex bolt, the concept is similar for an FRS. The shear strength of the interface between a Swellex bolt and the rock mass is greater than that of an FRS due to the mechanical interlock of the bolt and rock, otherwise they are presented as the same. Figure 5.2: Shear stress and axial load along a Swellex rock bolt (Li & Stillborg, 1999) Load decreases linearly in the decoupled length of bolt, and then continues to decrease at a rate proportional to the shear stress (i.e. the load in the bolt is proportional to the integral of the shear stress on the interface). As the decoupling front proceeds along the length of the bolt, a progressively longer portion of the bolt is subject to higher loads. This results in an apparent softening of the bolt behaviour, as total deformation is proportional to the integral of function P (x), representing load in the bolt.

59 Loadm(kN) Chapter 5. Summary Statistics and Interpretation of Pull Test Data Observed Behaviour of Friction Rock Stabilizers Figure 5.3 shows the load-displacement behaviour of an FA39 tested at Garson Mine. Six discrete loading phases are observed. 9 8 A B C D 7 6 E F Displacementm(mm) Figure 5.3: Pull test performed on an FA39 In Phase A load is built on the bolt, although significant displacement occurs. Most displacement may be attributed to the loading rig adjusting position as it becomes fully flushed with the surface of the rock mass, as well as the result of the initial rock mass and surface support response to pressure from the pull test apparatus. To eliminate this phase, a pre-load may be applied to rock bolts during a pull test before displacement is recorded. Phase B shows a response with decreasing stiffness as the decoupling front progresses and a larger length of bolt is subject to higher axial load, until limited slip of the bolt appears to occur in Phase C (i.e. the decoupling front has progressed along the entire length of the bolt). Once slip stops, the decoupling front regresses which results in a stiffer reaction observed in Phase D before slip occurs again entering Phase E. Load is then repeatedly built before slip occurs again and load is released. The test is stopped once pull testing personnel are satisfied that higher loads will not be reached, and Phase F shows unloading of the bolt as it re-couples with the rock mass and some elastic deformation is reversed (although there will still be elastic energy stored in the bolt due to frictional resistance acting against elastic recovery). Displacement data from the test in Figure 5.3 was collected by a data logger; load intervals are dictated by a frequency at which data is collected, and are inconsistent due to a variable loading rate. This results in a load-displacement graph that may appear different from one for which data is manually collected. In a manually-recorded pull test, testing personnel will generally wait for the load and displacement readings to stabilize before recording them and will seldom record two displacements for a single load, which makes perfectly plastic behaviour difficult to observe. Manual recordings of load and displacement also generally do not incorporate the pre-loading (Phase A) and unloading (Phase F) portions of a pull test for any bolt type.

60 Chapter 5. Summary Statistics and Interpretation of Pull Test Data Characterization of Performance Metrics for Friction Rock Stabilizers Ultimate Capacity Most types of rock bolt in the database were only tested until their working capacity is observed. However, FRS pull tests are performed to determine ultimate capacity. As the failure mechanism of a recently installed FRS is slip, there are fewer safety concerns associated with loading to ultimate capacity than exist for a bolt for which ultimate capacity is dictated by bolt failure. Additionally, most FRS pull test reports do not fully record the load displacement behaviours of the tested bolts and only note a maximum load. As such, ultimate capacity is investigated as opposed to working capacity for this bolt type. Figure 5.4 shows the distributions of the ultimate capacities for FRS A and B bolts. Bolts are separated by supplier and nominal diameter, and loads are recorded per unit length of anchorage (kn/m). The length of bolt providing anchorage is assumed to be 6 (.152 m) less than the total length of the bolt, as the tapered end and the section of bolt to which the pull test apparatus is attached are not in contact with the rock mass. Table 5.2 shows summary statistics for the maximum loads for the six bolt variants. The average measured diameter of the bolts is also noted. Diameter measurements were not performed using a consistent method between campaigns (although most reports mention the use of Vernier callipers); some diameters are an average of 3 or 5 measurements along the length of the bolt, some are a single measurement at the midpoint, and some are not explained. Table 5.2: Statistics regarding the ultimate capacities of FRSs Variant n x s ĉ v Skewness Kurtosis Average Diameter FA kn/m 1.3 kn/m mm FA kn/m 7.2 kn/m mm FA kn/m 11.1 kn/m mm FB kn/m 13.7 kn/m mm FB kn/m 13.3 kn/m mm FB kn/m 1.8 kn/m mm ALL kn/m 11.7 kn/m N/A There appears to be no statistically significant difference between the ultimate capacities of bolts with different nominal diameters, or between bolts from either supplier. ANOVA was performed on the data sets composed of FRS A bolts, FRS B bolts, and across all FRS configurations. Failure to reject the null hypothesis occurred for both the FRS A and B (p =.745 and.45 respectively), as well as for the ANOVA of all configurations (p =.695). This indicates that there is not a statistically significant difference in the ultimate capacity between configurations within or across supplier. While a larger diameter FRS will have greater surface area in contact with the rock mass on which to generate friction, it would appear that the larger diameter bolts do not generate the equivalent stress normal to the bolt-rock mass interface as the smaller bolts. This is demonstrated as resistance to pull is essentially the same between the different diameters of FRS. In fact, the 35 mm nominal diameter FRSs had the highest average ultimate capacities (although the difference is marginal and not statistically significant). This implies that if immediate resistance to axial loading is the main objective of bolt installation, there is apparently no advantage gained in the selection of a larger diameter FRS. Amalgamating all FRS configurations into one dataset results in a distribution of ultimate capacities that is very close to

61 Frequency Cumulative5Frequency Frequency Cumulative3Frequency Frequency Cumulative5Frequency Frequency Cumulative5Frequency Frequency Cumulative5Frequency Frequency Cumulative5Frequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data Frequency Cumulative 1% 8% 6% 4% 2% % Ultimate5Capacity5(kN/m) 16 1% 14 8% % 8 6 4% 4 2% 2 % Ultimate5Capacity5(kN/m) 12 (a) FA35 1% 18 (b) FB35 1% 1 8% % % 4% % 4% 2 2% 4 2 2% % Ultimate5Capacity5(kN/m) % Ultimate5Capacity5(kN/m) 12 (c) FA39 1% 25 (d) FB39 1% 1 8% 2 8% 8 6% 15 6% 6 4 4% 1 4% 2 2% 5 2% % Ultimate5Capacity5(kN/m) (e) FA46 % Ultimate3Capacity3(kN/m) (f) FB46 Figure 5.4: Ultimate capacity per unit length distributions for FRSs with nominal diameters of 35, 39 and 46 mm normal, with very low values of both skewness and kurtosis. Although the FA39 has the lowest ĉ v, this is likely the result of the low number of tests performed on it. One would expect similar standard deviations for this number of tests of a normally distributed random variable, however this is complicated by the fact that testing is conducted in campaigns. 3 pull tests performed on the FA39 are present in the database, but these are conducted in only 6 campaigns, i.e. 6 sets of conditions. The other diameters of FRS had a greater number of campaigns performed,

62 Loadm(kN) Chapter 5. Summary Statistics and Interpretation of Pull Test Data 48 and thus were potentially exposed to a wider range of conditions, resulting in larger standard deviations. This may also be the cause of the relatively high skew, as it is possible that the FA39 may have been tested more often in conditions that would result in marginally lower ultimate capacities. Kurtosis values are generally consistent. The exception is the FA46, with a kurtosis of.88, indicating a peakedness of the distribution. In Figure 5.4e, two outliers are observed: one in the 5-1 kn/m bin, and one in the 7-75 kn/m bin. This is the only distribution in Figure 5.4 that has single outliers this obvious. As a result, the shoulders have a lower weight than the centre of the distribution, elevating the value of kurtosis. With a greater sample size, this may be expected to reduce as the distribution fills in. Tomory et al. (1998) found that the 33 mm nominal diameter Split Set (SS33) and the 39 mm Split Set (SS39) also performed very similarly. The average ultimate capacity of their dataset of over 9 pull tests was 1.9 tons/ft, or 31.9 kn/m, with ĉ v =.42. Although the mean ultimate capacity is significantly lower than those observed in Figure 5.2, much of their database was composed of Split Sets installed using jacklegs, while the emergence of bolters since 1998 and their use at Vale s operations in Sudbury may explain the higher capacities observed in more recent times. It should also be noted that Tomory et al. collected data from over 5 mine sites. This means it is likely the pull tests contained therein were performed across a wider range of installation and ground conditions, and possibly testing methods/equipment. As a result, the higher coefficient of variation is expected. Overall, it does not appear as though one FRS configuration significantly outperforms any other in terms of either ultimate capacity or consistency of performance. Although there are irregularities in values of the coefficient of variance, skewness and kurtosis, these would likely be addressed by expanding the database. Stiffness Only campaigns of FA35 and FA39 bolts recorded the bolts load displacement behaviour in its entirety. As such, an analysis of stiffness is limited to these two bolt configurations. Figure 5.5 shows how two measures of stiffness are calculated for an FRS FirstmDropmStiffness SecantmStiffness Displacementm(mm) Figure 5.5: Stiffness metrics for an FRS

63 Frequency CumulativekFrequency Frequency CumulativekFrequency Frequency CumulativeSFrequency Frequency CumulativeSFrequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data 49 Measuring the stiffness of an FRS is made difficult by the constantly changing coupling length and the occurrence of slip before maximum load is achieved. The first drop stiffness outlined in Figure 5.5 represents the stiffness of the bolt at the displacement at which the first drop in load is observed. Whenever the bolt slips, a drop in load is expected. This drop is not consistently captured in the data recording due to the logging frequency. Slip distance appears to often be relatively short (less than 1 mm) before the bolt reaches a position with a higher frictional state at loads below the ultimate capacity. Where the first drop in load is observed is where it takes the bolt a greater period of time than the data logging frequency to reach this position and rebuild an equivalent load, and can thus be assumed to be a greater slip distance than any previously observed slip. The secant stiffness is representative of the displacement at which the maximum load is achieved. Both measures have their drawbacks, however they are satisfactory as broad descriptors of behaviour; the resolution of the data recorded is not sufficient to build a more robust model of bolt response. Figure 5.6 shows the distribution of stiffness for the FA35 and FA39. Table 5.3 summarizes the data. 6 1/ 7 1/ Frequency Cumulative 8/ 6/ 4/ 2/ / 6/ 4/ 2/ / FirstSDropSStiffnessS(kN/mm) / FirstSDropSStiffnessS(kN/mm) (a) FA35 first drop stiffness (b) FA39 first drop stiffness 6 1% 1 1% 5 8% 9 8 8% 4 6% 7 6 6% % 4 3 4% 1 2% 2 1 2% % SecantkStiffnessk(kN/mm) % SecantkStiffnessk(kN/mm) (c) FA35 secant stiffness (d) FA39 secant stiffness Figure 5.6: Stiffness distributions for the FA35 and FA39 The difficulties of assessing FRS performance emerge; standard deviations calculated for the distributions are very large relative to the means of the data (ĉ v > 1 in the case of first drop stiffness of the FA39). The heavy skew of the data calls into question the use of statistics such as standard deviation and kurtosis in these circumstances, as the distribution does not appear to be normal. The FA35 seems to have a higher secant stiffness than the FA39, however given the similarity of the first drop stiffnesses

64 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 5 Table 5.3: FRS stiffness summary statistics Stiffness Bolt n x s ĉ v Skewness Kurtosis First drop FA kn/mm 26. kn/mm FA kn/mm 26. kn/mm Secant FA kn/mm 18.1 kn/mm FA kn/mm 4.7 kn/mm between the bolts and the high coefficient of variations, it is unclear whether this is indicative of a consistent difference in responses. It is reiterated that these metrics are used as broad observational descriptors of behaviour. They are heavily influenced by several parameters unrelated to the bolt or how it interacts with the rock mass, such as data logging frequency and how much displacement the personnel conducting the test allow before stopping the pull test. As such, no further analysis was performed on the stiffness of FRSs although the information gained will be used to describe the expected behaviour of the bolts in a pull test. 5.3 Rebar Rock Bolts Theoretical Behaviour of a Rebar Rock Bolt As with frictional bolts such as the FRS, rebar offers resistance to pull along its length. Unlike the FRS, it is bonded to the rock mass with a continuous column of grout, the behaviour of which must be considered in a rebar model. Figure 5.7 shows a model of shear stress along a rebar rock bolt subject to a pull test. Figure 5.7: Model of the shear stress profile in a grouted rock bolt (Li & Stillborg, 1999) In this model, a length of bolt (x ) is fully decoupled from the rock mass. Between x and x 1, the bolt is partially decoupled with the shear stress along the interface between the bolt and the grout (τ b ) equal to the residual strength of the interface (s r ). Beyond point x 1, the grout is less damaged and has higher strength, with the peak strength (s p ) occurring at point x 2. Beyond this point, the full strength of the grout is not mobilized, and shear stress attenuates down the remainder of the bolt following function τ b (x). Figure 5.8 shows a similar shear stress profile, but shows function P (x), representing tensile load in the bolt (Li & Stillborg, 1999).

65 Load (kn) Chapter 5. Summary Statistics and Interpretation of Pull Test Data 51 Figure 5.8: Model of tensile load and shear stress profile for a rebar (Li & Stillborg, 1999) As in the case of the FRS model, load (and thus stress) in the bolt decreases at a rate proportional to the shear stress. As x 2 (the location of s p, the maximum strength) progresses down the bolt, an increasing length of the element is subject to higher stresses, and a decrease in measured stiffness is expected. Should x 2 reach the end of the bolt, the rebar would be pulled out through the grout. However, all pull tests in the database displayed behaviour suggesting P (x) exceeds the working capacity of the bolt at x =, i.e. at the bolt s threads Observed Behaviour of Rebar Rock Bolts Figure 5.9 shows the load displacement relationship for a pull test performed on a 2 mm Rebar B. These tests have a relatively straightforward interpretation; elastic deformation of the debonded portion of the rebar and rock mass compression result in the displacements observed in Phase A, before the rebar reaches its working capacity, and yields into Phase B A B Displacement (mm) Figure 5.9: Pull test performed on a 2 mm Rebar B The stiffness of a 1.8 m 4W steel bar with a diameter of 2 mm and an elastic modulus of 2

66 Load (kn) Chapter 5. Summary Statistics and Interpretation of Pull Test Data 52 GPa (as is used to manufacture Rebar B; Lamothe, November 214) is 34.9 kn/mm. The stiffness for the 1.8 m rebar shown in Figure 5.9 with an approximately linear response between 26.7 and kn (R 2 =.998) is 34.2 kn/mm. If displacement is interpreted as equal to the deformation of the rebar, this would suggest that stress is distributed uniformly along its length. However, this may in fact demonstrate the degree to which sources of displacement beyond elastic deformation of the bolt interfere with the measurement displacement during a pull tests. As very small displacements are being measured, displacements on the scale of one millimetre attributed to the rock mass or surface support will strongly affect the stiffness calculation. This may also explain why no progressive softening of the bolt response is observed. Figure 5.1 shows the same relationship for a 2 mm Rebar A, and perhaps a more typical pull test, where adjustments and rock fracturing are observable on the load displacement graph Displacement (mm) Figure 5.1: Pull test performed on a 2 mm Rebar A A tangent stiffness of 59.3 kn/mm is calculated for the linear portion of the graph, circled in red (R 2 =.999). The minimum elastic modulus of Rebar A, like Rebar B, is 2 GPa (Mainville-Beach, October 214), and thus also has a stiffness of 34.9 kn/mm if subject to constant axial stress along its length. This stiffness calculation agrees more closely with the analytical model of fully grouted rock bolts (Li & Stillborg, 1999; Martín et al, 21), suggesting that stress and strain is not evenly distributed along the length of the bolt. In any case, it is apparent by contrasting these two examples that the properties of the rebar material are not the only factors influencing the stiffness of the bolt system. It is noted that if the load were to be evenly distributed along a bolt s length, a longer bolt would appear less stiff. Although different lengths of rebar are present in the database, length normalized stiffness metrics are not used as the degree of deformation the bolt undergoes is much better described by the location of maximum shear stress on the bolt (x 2 in Figure 5.7) than by the overall length of bolt.

67 Frequency Cumulative5Frequency Frequency Cumulative5Frequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data Characterization of Performance Metrics for Rebar Rock Bolts Working Capacity Figure 5.11 shows the distribution of working capacities for rebar supplied by Suppliers A and B. The rock bolts were typically installed using the proprietary resins of each manufacturer. 12 1% 14 1% Frequency Cumulative 8% 6% 4% 2% % 6% 4% 2% 1 % Working5Capacity5(kN) 1 % Working5Capacity5(kN) (a) 2 mm Rebar A (b) 2 mm Rebar B Figure 5.11: Working capacity distributions for rebar Working capacity appears to be distributed normally, although data resolution poses a problem for pull tests conducted by Supplier B. Although the Rebar B working capacities are not as well distributed as those of Rebar A, this may be the result of differing data recording methods. Tests performed by Supplier B typically are performed in loading increments, with manual recording of displacement, while pull tests conducted by Supplier A generally log data digitally. As loading increments in a manual test are usually one ton, working capacity may only be calculated to the nearest ton (8.9 kn). As such, no working capacities may be recorded for Rebar B between 14 and 15 tons (124.6 kn and kn), thus the lack of entries in the 125 kn to 13 kn bin and the resulting appearance of the distribution. On the other hand, Supplier A s load recordings are distributed according to loading rate and measurement frequency. As such, the variance of the distribution observed is dependent not only on the variability in rebar material properties, but also the variability in loads recorded by the apparatus. This introduces a type of error that is also distributed normally as a random variable. Table 5.4 provides the summary statistics for the working capacities of the rebar. Table 5.4: Summary statistics for the working capacity of rebar Supplier n x s ĉ v Median Skewness Kurtosis A kn 7.5 kn kn B kn 7.4 kn kn The average working capacities of the two rebar are very similar. The small discrepancy between the two averages may be a result of the difference in the method of data recording rather than performance of the elements themselves. A 1 kn difference in the medians of the two datasets also suggests that little difference would exist between the two, if the same method of data recording were used for both rebar. It may also be concluded that the negative skew and high kurtosis observed for the Rebar B distribution is the result of the rounding of working capacity down to the nearest ton.

68 Frequency CumulativesFrequency Frequency Cumulative(Frequency Frequency CumulativeNFrequency Frequency CumulativeNFrequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data 54 Stiffness Tangent and secant stiffness were calculated for the results of the pull tests on rebar. A large portion of the pull tests performed were partial encapsulation tests. In these tests, a limited length of active resin (usually one 12 or 18 cartridge of fast setting resin) is used in tandem with two or three cartridges of dummy or inert resin used to simulate typical mixing conditions for the active cartridge. These tests are performed to verify the competence of a limited length of resin. While this test configuration should not affect the working capacity of the rebar, it may influence the load displacement behaviour of the bolt. As such, a distinction is made between the fully and partially encapsulated tests in Figures 5.12 and 5.13, which show the secant stiffness and tangent stiffness respectively of both rebar. Table 5.5 shows summary statistics for the stiffness of the two brands of rebar. 12 1p 7 1% PartiallyNEncapsulated FullyNEncapsulated Cumulative 8p 6p 4p 2p % 6% 4% 2% SecantNStiffnessN(kN/mm) p SecantNStiffnessN(kN/mm) % (a) Rebar A secant stiffness (b) Rebar B secant stiffness Figure 5.12: Secant stiffness for 2 mm rebar 9 1) PartiallysEncapsulated FullysEncapsulated Cumulative 8) 6) ) 2) TangentsStiffnessspkN/mmd ) Tangent(Stiffness((kN/mm) (a) Rebar A tangent stiffness (b) Rebar B tangent stiffness Figure 5.13: Tangent stiffness for 2 mm Rebar

69 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 55 Table 5.5: Summary statistics for the stiffness of rebar Stiffness Supplier n x s ĉ v Skewness Kurtosis Secant A kn/mm 13.4 kn/mm B kn/mm kn/mm Tangent A kn/mm kn/mm B kn/mm 1.23 kn/mm % of the rebar pull tests in the data base were performed on 1.8 m, 2 mm diameter rebar with an elastic modulus of 2 GPa, and thus a stiffness of 34.9 kn/mm when non-grouted. This is a very conservative lower bound of stiffness in which recorded displacement can only be attributed to the deformation of the rebar, as it assumes an even distribution of stress along the bolt. Figures 5.12a and 5.12b show that a large number of secant stiffness calculations fall below this value. This is likely due to the incorporation of sources of displacement besides deformation in the measurements. It is interesting to see that the tangent stiffness also frequently falls short of 34.9 kn/mm, suggesting that these alternate displacement mechanisms may in some cases have linear load displacement responses at the load resolutions in question, making it difficult to isolate the bolt/resin response from that of the rock mass. High values of skewness and kurtosis are calculated for Rebar A in particular. Although these are calculated for a dataset that incorporates both partially and fully encapsulated bolts, it can be observed in Figures 5.12a and 5.13a that high stiffness outliers exist which are responsible for these values. Table 5.6 separates stiffness statistics by encapsulation length. Table 5.6: Comparison of stiffness between partially and fully encapsulated rebar Variable Supplier Encapsulation n x s ĉ v A Full kn/mm 17.1 kn/mm.59 Secant stiffness Partial kn/mm 9. kn/mm.39 B Full kn/mm 14.2 kn/mm.52 Partial kn/mm 4.2 kn/mm.28 A Full kn/mm 31.2 kn/mm.62 Tangent stiffness Partial kn/mm 9.3 kn/mm.31 B Full kn/mm 1.2 kn/mm.33 Partial kn/mm 4.7 kn/mm.27 Interestingly, the secant stiffness of fully encapsulated rebar from both suppliers are quite similar (averages of 29.2 kn/mm and 27.4 kn/mm respectively), with similar dispersions (coefficients of variation equal to.59 and.52). However, Rebar A have greater values of tangent stiffness (an average of 5.6 kn/mm compared to 3.9 kn/mm). Comparing the distributions presented in Figures 5.12 and 5.13, the fully encapsulated Rebar B appear to have a much more defined distribution shape for both tangent and secant stiffness, suggesting the sample size of Rebar A is too small to adequately define these statistics, despite being the same size as the Rebar B dataset. This may be the result of the difference in data recording methods. While the displacements measured by Supplier A are taken instantaneously at loads defined by loading rate and data logging frequency, it is common practice in manual data recording of pull tests to wait for displacement measurements at a certain load to stabilize before reading them, potentially resulting in what appears to be lower stiffness. The reason the secant stiffness measurements are similar for the two brands of rebar despite the disparate distributions of tangent stiffness may lie in the pre-loads used. Relatively large displacements may be observed at low loads while the rock mass

70 SecantmStiffnessmukN/mmp Tangent/Stiffness/(kN/mm) Chapter 5. Summary Statistics and Interpretation of Pull Test Data 56 and pull test apparatus tighten up. 8% of the pull tests on fully encapsulated Rebar A were performed without a pre-load, compared to only 11% of pull tests performed on Rebar B. This can explain both the ragged distribution shape of the stiffness metrics distributions (as the amount of displacement at low loads is highly variable), and the high tangent stiffness relative to the secant stiffness calculated for Rebar A. In addition to tangent and secant stiffness calculations, a total of 1 unloading stiffness measurement were calculated from the available data, including both partially and fully encapsulated Rebar A. Figure 5.14 compares the stiffness of the unloading phase with the corresponding tangent and secant stiffness calculated for that rock bolt. 7 Fullmencapsulation Partialmencapsulation Fullmencapsulationmregression Partialmencapsulationmregression All ym=m.199xmrm13.34 R²m=m y/=/.5356x/+/1.543 R²/=/.7528 y/=/.511x/+/ R²/=/ y/=/.3272x/+/7.254 R²/=/ UnloadingmStiffnessmukN/mmp Unloading/Stiffness/(kN/mm) (a) Secant stiffness (b) Tangent stiffness Figure 5.14: Comparison of Rebar A unloading stiffness to secant (a) and tangent (b) stiffness As shown in Figure 5.14, a positive correlation exists between unloading stiffness and both secant and tangent stiffness. However, the relationship has a much higher coefficient of determination for the unloading stiffness tangent stiffness relation. Note that this relationship appears to show that tangent stiffness overestimates the amount of displacement attributable to the elastic deformation of the bolt by a factor of 2 for partially encapsulated pull tests, and a factor of 3 for the fully encapsulated tests. This demonstrates that tangent stiffness may be used as an indicator of the quality of the bond between rebar and rock mass, although not necessarily a direct quantifier of elastic deformation of a rebar rock bolt subject to a pull test. 5.4 Modified Cone Bolts Theoretical Behaviour of a Modified Cone Bolt The Modified Cone Bolt is a yielding support element designed primarily to absorb energy resulting from dynamic loading scenarios imposed by seismicity. As such, analysis of its behaviour in static or quasi-static conditions is generally considered secondary. Laboratory testing of the cone bolt has shown that the displacement of the head of a cone bolt is strain rate dependant; in dynamic scenarios plough of the bolt contributes significantly to displacement, while in static loading scenarios displacement is mostly a result of deformation of the bolt itself (Simser et al, 26). This is illustrated in Figure 5.15.

71 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 57 (tonnes) mm cone plow 6 mm cone plow (mm) Figure 5.15: Laboratory pull test of an MCB33 (Simser et al, 26) Around 5 mm of displacement occurred before strain hardening is induced. Only 17 mm of this was attributed to cone plough. Of the remaining 17 mm of displacement recorded, just 6 mm is attributed to plough. This suggests that there is no constant load in a quasi-static loading scenario at which plough occurs consistently before the bolt fails, although plough does occur as load increases. A conceptual depiction of the behaviour is shown in Figure Steel stretching Load Eventual failure of the bolt Limited plow of the cone (~ 2-4 mm) Deformation Figure 5.16: Conceptual load displacement behaviour of a cone bolt subject to quasi-static loading (Simser et al, 26) Observed Behaviour of Modified Cone Bolts The pull tests on cone bolts in the assembled database appear to generally react to pull testing in one of two ways. Figure 5.17 shows a typical load displacement relationship for the first of these behaviours, representative of 83% of the MCB33 pull test data. This relationship is split into four phases. A pre-load of 3 tons (27 kn) is applied before displacement is recorded. Phase A is a stiff initial response, which softens into Phase B. Phase C shows displacement that is close to perfectly plastic, before the system stiffens again in Phase D. The stiff initial response of the bolt may be primarily attributed to elastic deformation of the bolt tendon before plough occurs at the A B transition. Subsequently, displacement is attributed to both plough and elastic deformation of the bolt as load increases, until the tendon yields into Phase C. Displacement observed in Phase C is likely primarily a result of plastic deformation; as the bolt tendon is a 2.4 m smooth bar, load should be evenly distributed along its length, particularly as plough cannot occur before it is entirely decoupled from the resin. This means that the 3-4 mm of plastic deformation observed in Phase C corresponds to a strain of about 1.5% before strain hardening seems to occur at the C D transition. A strain of this

72 Load (kn) Load (kn) Chapter 5. Summary Statistics and Interpretation of Pull Test Data 58 magnitude occurring between yield and strain hardening may be observed for a carbon steel, depending on composition and manufacturing process (ASM International, 22). It should be noted that testing on the bolt shown in Figure 5.17 was stopped before failure occurred A B C D Displacement (mm) Figure 5.17: Pull test performed on an MCB33 with plough The linear plough elastic deformation phase of the bolt was not observed in all pull test on the MCB33. 17% of the tests exhibited a behaviour that suggested that the bolts responded to load with relatively little plough, although they may exhibit a progressive softening of the bolt/grout system. Figure 5.18 illustrates an example of such behaviour Displacement (mm) Figure 5.18: Pull test performed on an MCB33 without a linear plough response The dashed red line in Figure 5.18 represents the linear elastic behaviour of a bar of equivalent length, diameter and elastic modulus (2 GPa) as the tested bolt. Figure 5.19 focuses on the initial response recorded for the same bolt, which seems to closely follow a second order polynomial trend between the start of the test and the yield of the tendon (R 2 =.9992). This gradual deviation from the elastic response of the tendon suggests that as load increases, there is movement of the cone and the resin is

73 Load (kn) Chapter 5. Summary Statistics and Interpretation of Pull Test Data 59 becoming more damaged at an accelerating rate, but is competent enough to resist a linear plouging response until the bolt tendon yields R 2 t=t Displacement (mm) Figure 5.19: Close-up of the bolt response shown in Figure 5.18 Examining both of these divergent displacement behaviours highlights the difficulties associated with assessing the performance of a bolt. Cone bolts are designed to plough through resin under dynamic loads, so from one perspective it may be considered proof of concept and a successful pull test if the bolts plough under static conditions (as in Figure 5.17). However, if the cone bolt is loaded in static conditions it is possible that as little deformation as possible is desirable (as in Figure 5.18), and significant displacements should ideally only occur during dynamic events if a bolt is to be said to perform well in both loading conditions (although there is no widespread methodology used to test the dynamic capabilities of a rock bolt in situ) Proposed Interpretation of Modified Cone Bolt Behaviour Figure 5.2 provides a slightly modified interpretation of the conceptual behaviour for the cone bolt, based on a review of the undertaken pull tests. It must be noted that this model only applies in quasi-static axial loading scenarios. At low loads, displacement is primarily due to elastic deformation of the bolt tendon. At a certain load threshold, the cone begins to plough through the resin. A new load response is established, for which displacement is attributed to both elastic deformation of the bolt (as load continues to increase), and plough of the cone through the resin. Note that at the resolution data was collected for most MCB33 pull tests, this behaviour appeared linear. At higher resolutions repeated load build and drop may be observed if the bolt ploughs incrementally. The tendon subsequently yields and deforms in a perfectly plastic manner while little to no plough occurs. The tendon then strain hardens, load on the bolt builds, and as a result plough resumes. In the case presented, there does not appear to be a single threshold of load at which plough is sustained. Dynamic loading of the MCB33 results in very different load displacement behaviour (St-Pierre et al, 29; Doucet & Voyzelle, 212) as the properties of both the steel of the cone bolt tendon and the encapsulating resin vary with loading rate. This behaviour of the cone bolt may be broken down into 5 parameters (Figure 5.21). While length

74 Load (kn) Load Chapter 5. Summary Statistics and Interpretation of Pull Test Data 6 Yield of bolt tendon Strain hardening To failure of bolt Plastic deformation of tendon, limited plough Elastic deformation and plough Plough begins Elastic deformation of bolt tendon Displacement Figure 5.2: Amended conceptual load displacement behaviour of a cone bolt of the bolt will be a controlling factor on any stiffness metric dependent on the deformation of the bolt tendon, all MCB33s in the database were the same length. As a result, stiffness is not normalized to length, and is expressed in kn/mm SteelkYieldkStrength PloughkStiffness SecantkStiffness PloughkPoint InitialkStiffness Displacement (mm) Figure 5.21: Performance metrics measured from a cone bolt pull test The initial response to loading is measured by the initial stiffness. The plough point is the load at which behaviour deviates from the initial stiffness, and plough begins. The plough stiffness measures the gradient of the combined plough/elastic deformation response, until the steel yield strength. The secant stiffness is the average stiffness between the first measurement of load and displacement and yield of the bolt. The term working capacity is not used to describe the capacity of a cone bolt as its determination in this context is somewhat ambiguous. Unlike the other grouted bolts discussed, the response to load consistently deviates from linearity before the tendon yields. As such, the working capacity could be defined as the plough point, but this neglects the fact that substantially more load may be borne by the element at larger displacements. This also results in a potential complication if the bolt ploughs from the beginning of a pull test, as the plough response becomes the linear response from which deviation

75 Frequency Cumulative Frequency Frequency Cumulative)Frequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data 61 defines working capacity. Values of cone bolt capacity used in the design of a support system should depend on the design methodology, chiefly whether or not displacement is taken into account. As such, the term working capacity will not be used in the context of cone bolts, and either the plough point or the yield strength will be specified Characterization of Performance Metrics for Modified Cone Bolts Load Metrics Figure 5.22 shows the distributions of plough point and yield strength. Table 5.7 shows the summary statistics calculated for these two metrics % 8% 6% % 8% 6% 6 4% 6 4% 4 2 2% 4 2 2% % Plough Point (kn) % Yield)Strength)(kN) (a) Plough Point (b) Yield Strength Figure 5.22: Load metric distributions for the MCB33 Table 5.7: Summary statistics of MCB33 load metrics Variable n x s ĉ v Skewness Kurtosis Plough Point kn 22.7 kn Yield Strength kn 16.3 kn The observed distributions can be explained with a careful review of the pull tests compiled. The initial peak in Figure 5.22a and the resulting heavy skew of the MCB33 yield strength distribution is attributed to one testing campaign from 26 which contributed 5 of the 7 values in the 8-9 kn bin for the yield strength, and as a result is excluded from the calculation of statistics. The yield strength is the variable with the lowest coefficient of variation, which is to be expected as it should be heavily dependant on the properties of the steel as opposed to the resin. The plough point appears to be normally distributed about an average of 53.2 kn, less than half of the yield strength of the bolt. There are a large number of bolts that plough between 2 and 3 kn, which may be representative of bolts where ploughing began during pre loading. Stiffness Metrics Distributions for initial stiffness, plough stiffness and secant stiffness are shown in Figure Summary statistics are presented in Table 5.8.

76 Frequency Cumulative Frequency Frequency Cumulative(Frequency Frequency CumulativemFrequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data Initial(Stiffness((kN/mm) (a) Initial Stiffness % 8% 6% 4% 2% % 7 1a 6 8a 5 4 6a 3 4a 2 2a a PloughmStiffnessm(kN/mm) (b) Plough Stiffness 1% 8% 6% 4% 2% Secant Stiffness (kn/mm) (c) Secant Stiffness % Figure 5.23: Stiffness metric distributions for the MCB33 Table 5.8: Summary statistics of MCB33 stiffness metrics Variable n x s ĉ v Skewness Kurtosis Initial Stiffness kn/mm 4.1 kn/mm Plough Stiffness kn/mm.9 kn/mm Secant Stiffness kn/mm 3.51 kn/mm The stiffness of a 2 GPa, 2.4 m bar with a diameter of 17.2 mm is 19.3 kn/mm. As the initial reaction of the MCB33 to load is though to be a result of elastic deformation of the tendon, one would expect the initial stiffness to be approximately equal to 19.3 kn/mm. However, very few bolts exhibit a response this stiff. As seen in Section 5.3.3, a linear measure of stiffness may overestimate displacement attributable to bolt deformation by a factor of 2 to 3, which could explain why so few bolts exhibit a stiffness of 19.3 kn/mm. The secant stiffness of the MCB33 has a strong right skew and high coefficient of variation; as it is a function of the four other parameters, non-normality is not unexpected. The plough stiffness is quite evenly distributed between 1 and 3 kn/mm as indicated by the negative kurtosis, showing that the cone resin interaction appears to be highly variable. This could suggest that the in situ dynamic performance of the cone bolt may be subject to a similarly variable energy capacity. It must be noted that the pull test is strictly analogous to static or quasi-static loading of a bolt. The dynamic performance of a reinforcement element may be assessed using impact testing (as reported in

77 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 63 Hadjigeorgiou & Potvin, 211), or using passive monitoring in situ (as in Morissette et al, 214). The MCB33 has been demonstrated to perform significantly differently when subject to impact loading versus quasi-static (Doucet & Voyzell, 212), as such the results of one loading rate may not indicate a similar effect for another. 5.5 D-Bolts Theoretical Behaviour of a D-Bolt D-Bolt performance is difficult to evaluate from a conventional pull test. The principle behind the design of the bolt is to evenly distribute load across the length of the smooth bar between two anchors, potentially resulting in significant differences in load between two adjacent smooth sections. Published laboratory static testing was performed across a simulated joint, where load is applied at the midpoint between two anchors (Li, 212). Figure 5.24 shows the apparatus and test set-up used for this test, and Figure 5.25 shows the results. Figure 5.24: Apparatus for a simulated joint laboratory test on a D-Bolt (Li, 212) Figure 5.25: Results of simulated joint laboratory tests on 2 mm D-Bolts (Li, 212) The test shown in Figure 5.25 is performed on two bolts grouted in cement: OP1 and OP2. OP2 has a shrink sleeve on the test section while OP1 does not, otherwise the bolts are identical (Li, 212. The D-Bolts used by Vale do not have sleeves, and are grouted in resin). Bolt load is very similar for both tests, and shows typical steel stress-strain behaviour. For test OP1, the plate experiences an increase in load around when the test section or tendon starts to yield, while OP2 s plate is loaded slightly

78 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 64 before. With such a small sample size, it is difficult to determine whether this difference in behaviour may be attributed to the difference in bolt surfaces. In any case, it is apparent that load does appear to propagate past the anchors into adjacent smooth sections, insinuating that there is some limited movement of the anchors through the grout. The D-Bolts in the database are 2.4 m in length with three.12 m long anchor sections, at.34 m, 1.14 m and 2.14 m along the length of the bolt. Although an in situ pull test is different from the procedure shown in Figure 5.24 as load is applied to the head of the bolt, possible behaviour may be hypothesised. Load could be limited to the first,.34 m segment of the bolt between the head and the first anchor assuming good resin encapsulation, resulting in a very stiff response (the test sections for the bolts shown in Figure 5.25 were.9 m in length; Li, 212). If load does propagate past the first anchor, a drop in stiffness would be observed as the adjacent smooth section will also be strained. Table 5.9 shows the stiffness that would be expected of both the 2 and 22 mm D-Bolts used at Vale s Sudbury operations, depending on the length of bolt exposed to load. This is defined by the anchor at which no movement occurs, and thus prevents load from further propagating down the bolt. The steel used in the manufacture of D-Bolts has an elastic modulus of 2-21 GPa (Charette, 214). Table 5.9: D-Bolt stiffness by anchor stability Stable Anchor Length 2 mm D-Bolt 22 mm D-Bolt 1 st.34 m 184 kn/mm 227 kn/mm 2 nd 1.14 m 55 kn/mm 68 kn/mm 3 rd 2.14 m 28 kn/mm 34 kn/mm The calculations shown in Table 5.9 should be used only as a rough guideline. An assumption is made that the anchor sections are of equivalent diameter to the smooth sections. It also assumes that if, for example, the second anchor is stable, the first bears no load and the same magnitude of stress is present in both smooth sections on either side of it. If the first anchor were to be load-bearing, the stiffness of the element as a whole would be dependent on the magnitude of that borne load Observed Behaviour of D-Bolts Figure 5.26 shows two typical load displacement relationships for a 22 mm D-Bolt pull test. The dashed lines represent the expected stiffness for a 22 mm D-Bolt assuming different lengths defined by anchor placement are subject to the applied load (see Table 5.9). The progressive stiffening of the bolt in Figure 5.26 tested at Creighton in the elastic phase is not an isolated case in the data collected. As with other bolts, compression of the rock mass may account for displacement in some circumstances. The fact that this effect is observed consistently in D-Bolt pull tests may indicate that D-Bolts are usually installed in ground susceptible to rockbursts, which can be heavily fractured (and thus more compressible) as a result of high stresses. It appears, examining only the stiffness, as though the bolt is not acting as would be expected if load was isolated solely between the testing jack and the first anchor, but as if load were evenly distributed along its length. This could indicate movement of one or more anchors through the resin as they seat in place, or if the anchors are not properly encapsulated (possibly due to the aforementioned fractured ground).

79 LoadC(kN) Chapter 5. Summary Statistics and Interpretation of Pull Test Data Creighton,C214 CopperCCliff,C29 BeforeCfirstCanchorC(.34Cm) BeforeCsecondCanchorC(1.14Cm) BeforeCthirdCanchorC(2.24Cm) DisplacementC(mm) Figure 5.26: Pull tests performed on 22 mm D-Bolts An insight into how load is distributed along the bolt is provided by the displacement attributed to plastic deformation of the bolt before strain hardening occurs. Laboratory testing performed on D-Bolts by Li (212; Figure 5.25) exhibited slightly more than 1 mm of perfectly plastic deformation between yield and strain hardening on a smooth section of.8 m, corresponding to a strain of around 1.3%. The bolt pulled at Creighton in 214 shown in Figure 5.26 exhibits about 2.5 mm of near perfectly plastic deformation before strain hardening. Assuming similar steel was used for the manufacture of this bolt as was used for the bolts tested by Li (212), 2.5 mm of deformation representing a strain of 1.3% corresponds to a deformation length of about 19 mm. The distance between the end of the threaded section of the D-Bolt head and the first anchor is 215 mm. This calculation is complicated by the fact that the 22 mm D-Bolt uses an M24 x 3. thread (Normet, 214). This thread has a major (i.e. maximum) diameter of 24 mm and the thread crests are 3 mm apart. The minor (i.e. minimum) diameter for an M24 x 3. thread is 2.7 mm for flat form threads (ASME, 25). Assuming the threading process does not significantly alter the yield strength of the bolt material, the lowest diameter part of the bolt (the 2.7 mm minor thread diameter) will yield first, about 2 kn before the bar as cross-sectional area is proportional to yield load. The thread is by definition of varying diameter, so yield will be a more protracted response (as observed in Figure 5.26) in comparison to the yield of the smooth bar section alone in Figure As the major (maximum) diameter is 24 mm, the 22 mm smooth section will yield before the entire threaded section does. As a result, it can be concluded that the near perfectly plastic response observed in Figure 5.26 is attributed to the first smooth section between the thread and the first anchor, implying that although the bolt acts in a manner soft enough to suggest little to no anchorage provided by the first two anchors, the first section bears the most load. This does not agree with the findings of an analysis solely of secant and tangent stiffness, thus demonstrating the value of measuring displacement beyond the working capacity and analysing bolt behaviour. However, the measured stiffness of the bolts should not be ignored. D-Bolts are often pulled until just after, or even before, their working capacity. Figure 5.27 shows the results of three 2 mm D-Bolts installed in the same rock mass at the same location during one testing campaign. These bolts were installed at angles between 15 and 3 off of perpendicular from the wall face. It is not clear how this

80 Loadl(kN) Chapter 5. Summary Statistics and Interpretation of Pull Test Data 66 affects the results, although the appearance of a working capacity of about 12 kn was attributed to movement of the testing rig (the minimum yield load of a 2 mm D-Bolt is 14 kn) Boltl1 Boltl2 Boltl3 Beforelfirstlanchorl(.34lm) Beforelsecondlanchorl(1.14lm) Beforelthirdlanchorl(2.24lm) Displacementl(mm) Figure 5.27: Pull tests performed on 2 mm D-Bolts As in Figure 5.26, the expected stiffness of varying length of smooth bar in tension is displayed, in this case the diameter of which is 2 mm. Bolt 1 has the least stiff response, on average performing very similar to what would be expected if only the third anchor was stable. Bolt 2 is the most stiff, its behaviour indicating that strain in the bolt is largely concentrated before the first anchor. After an initially softer response, Bolt 3 develops a stiffness that almost exactly matches the expected deformation of a 1.14 m bar. The initial response could be explained by rock mass compression, or perhaps by the second anchor fully seating after limited displacement and preventing further development of load on the third anchor. With these three pull tests, assuming the installation and ground conditions are nearidentical between them, it is clear that a variety of behaviours can be expected from the D-Bolt in a pull test. Limited unloading data for D-Bolts was also present in the database. The results of a single campaign on 22 mm D-Bolts are shown in Figure Ground conditions at the location of the pull tests (which were performed in the back) are described in the pull test report as broken, and the large displacements observed during loading are likely a result of this. As was the case with rebar, the unloading phase is used as a way to measure elastic deformation of the bolt. Unlike the rebar pull tests, the unloading phase was explicitly targeted by the personnel conducting the pull test, and data was recorded manually. As such, the stiffness may be calculated with the maximum and minimum load values recorded. Table 5.1 shows the unloading stiffness calculated for the three pull tests. Table 5.1: Unloading stiffness calculated for pull tests performed on 22 mm D-Bolts Test Stiffness kn/mm 2 54 kn/mm 3 36 kn/mm

81 Load (kn) Chapter 5. Summary Statistics and Interpretation of Pull Test Data Test 1 Test 2 Test Displacement (mm) Figure 5.28: Pull tests performed on 22mm D-Bolts Each of the three tests exhibits different degrees of elastic recovery during the unloading phase, implying variable load distributions between the three bolts. Comparing the stiffness calculated for these bolts with the different values of stiffness expected of a 22 mm D-Bolt (Table 5.9), certain similarities are observable. Test 1 exhibits very little displacement during unloading, resulting in a very high stiffness value. It must be acknowledged that for such high values of stiffness, the measured displacements will be correspondingly very small, and any error in measurement will be magnified. Test 3 unloads in a manner that would be similar to that of a bolt with an even load distribution along its length until the third anchor, and Test 2 as if the bolt were subject to a reduced load between the second and third anchors. The fact that different bolt behaviours are characterized over limited testing agrees with the conclusions drawn from Figure 5.27; the stiffness of the D-Bolt is dependent on the mobility of its anchors, which is apparently subject to change between bolts even if installed in very similar conditions Characterization of Performance Metrics for D-Bolts Working Capacity An in-depth analysis of working capacity is omitted due to an insufficient amount of data, with a total of 6 reliably observed working capacity values (to the nearest ton) between the two D-Bolt diameters. As too few measurements of working capacity exist in the database to reliably characterise dispersion, all values of working capacity for the D-Bolt present in the database are shown in Table Table 5.11: Working capacities obtained from all D-Bolt pull tests Diameter Working Capacity 142 kn 2 mm 16 kn 142 kn 196 kn 22 mm 196 kn 196 kn

82 More Frequency CumulativesFrequency Frequency CumulativekFrequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data 68 Stiffness Figure 5.29 shows the distribution of stiffness for the two diameters of D-Bolt. Refer to Table 5.9 for the anticipated stiffness of each bolt for different distributions of load along its length Histogram 12.% 1.% 8.% 6.% 22smm 2smmFrequency 4.% Cumulatives) % 2.%.% 1) 8) 6) 4) % 8% 6% 4% 2 1 Bin 2) 2 1 2% TangentsStiffnesss(kN/mm) ) SecantkStiffnessk(kN/mm) % (a) Tangent Stiffness (b) Secant Stiffness Figure 5.29: Distributions of stiffness for 2 and 22 mm D-Bolts These figures show that in most of the pull tests performed, either all bolts acted in a manner that would suggest that the load distribution extended beyond the first anchor (and the second anchor for the majority of bolts), or rock mass compression played a large role in the tests. Two 2 mm bolts have tangent stiffnesses that strongly suggest the second anchor is firmly held in the resin, and although there is one test with higher stiffness, none of the tests approach the stiffness that would be observed if there were no movement in the first anchor. Table 5.12 shows summary statistics for the D-Bolts. Table 5.12: D-Bolt summary statistics Stiffness Bolt Diameter n x s ĉ v Skewness Kurtosis Tangent 2 mm kn/mm 22.3 kn/mm mm kn/mm 8.3 kn/mm Secant 2 mm kn/mm 18.1 kn/mm mm kn/mm 8.8 kn/mm The undertaken analysis is to be interpreted carefully. A very small volume of data is available, and in reality there are three possible clusters of bolt stiffness, one for each anchor. Normal distributions are not to be expected with such variable behaviour. Insufficient data was obtained to examine each of these clusters on an individual basis. The data appears to portray the 2 mm D-Bolts as generally acting stiffer than the 22 mm equivalent. This is likely misleading; two thirds of the 2 mm D-Bolt pull tests were performed at Copper Cliff Mine, on 355 and 371 Levels (1,8 and 113 m), while over half of the 22 mm D-Bolt pull tests were performed at Creighton Mine at twice that depth, between 768 and 794 Levels (234 and 242 m) in rock masses noted to be fractured. While the quality of the rock was not noted for the shallower Copper Cliff tests, the lower far field stresses may correlate to a less intensely damaged rock mass in which the bolts were installed.

83 Load (kn) Chapter 5. Summary Statistics and Interpretation of Pull Test Data Expandable Bolts Theoretical Behaviour of an Expandable Bolt In theory, the concepts that dictate the behaviour of an expandable bolt should be similar to those of an FRS as both are friction bolts. Figure 5.2 shows the load and shear stress distributions for a Swellex bolt subject to a pull test, and was used to explain the behaviour of an FRS in Section 5.2. The main difference between the behaviour of an FRS and an expandable bolt is the higher shear strength of the bolt rock mass interface. This is provided by the mechanical interlocking of the bolt and the rock mass once the bolt is inflated (Li & Stillborg, 1999). As a result, an expandable bolt is more likely to reach its yield or ultimate tensile strength before it slips entirely out of the hole, although this depends on the length of bolt embedded in the rock mass Observed Behaviour of Expandable Bolts Although several brands of expandable are available from suppliers, the analysis was performed on four types of Swellex: the Pm12, Mn12, Pm24 and Mn24. There are four typical behaviours for a pull test on an inflatable bolt observed in the database. Figure 5.3 shows examples of three, with ideal elastic stiffness shown for bolts of equivalent length, cross sectional area and elastic modulus (2 GPa). Note that although the Mn bolt is a yielding reinforcement element, both the Pm and the Mn bolts have very similar elastic moduli; the additional deformation capacity of the Mn line is a result of its plastic behaviour (Bureau, 25). 12 WorkingWCapacity Pm12W(2.4Wm) PcPm12W(1.8Wm) PcMn12W(2.4Wm) 1.8WmWStiffness 2.4WmWStiffness Displacement (mm) Figure 5.3: Pull tests performed on Pm12 and Mn12 expandable bolts Some expandable bolts exhibit an extremely stiff response before gradually softening as load increases, implying that a limited section of the bolt is initially deforming and the decoupling front mobilizes as load increases (Li & Stillborg, 1999). Conversely, some bolts have a very soft initial response before stiffening. This may be due to the response of the rock mass, but may potentially also indicate bolt movement. The third behaviour is simply a linear response of uniform stiffness, and the fourth behaviour is highly variable, with no discernible pattern in response to load. A number of partial embedment tests are included in the database. In these tests, the bolt is inserted

84 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 7 into a steel tube with a diameter less than that of the installation hole, and a limited length left exposed (in the cases present in the database, this is usually 1, or.348 m). When the bolt is inflated, only the exposed part of the bolt comes in contact with the rock mass and provides anchorage. Table 5.13 shows bolt behaviour by bolt type, omitting partial embedment tests. Acknowledging very limited data, it is interesting to note that the Pm12 bolts appear to soften or displace in a uniform manner more often than the other bolt types, although it must be noted that all pull tests on the Pm12 were performed in sandfill or paste. Table 5.13: Swellex behaviour breakdown Variant n Softening Uniform Stiffening Variable Pm % 32% 5% 21% Mn % 25% 25% 37.5% Pm % 27% 12% 47% Mn % 22% 11% 44 % If the objective of a pull test on an expandable bolt was to determine the strength of the bolt rock interface (as it is for an FRS), then a length normalized value of capacity would be adopted as capacity would be dependent on the surface area (and thus length) of the interface. However, Swellex bolts in the database generally yield before slipping extensively. The working capacity of the bolt is thus dependent on the cross-sectional area of the steel, and independent of the bolt s length. As such, load is expressed in absolute terms (i.e. kn). In these cases it is only possible to calculate a minimum strength of the bolt rock interface. Bolts subject to a partial embedment pull test usually fail by slip, in which case the shear strength of the interface can be calculated in terms of load per length. While a limited number of fully embedded pull tests on both Pm12 and Pm24 did fail by slipping, all tests were performed in sandfill. Table 5.14 summarizes all slipped tests. Insufficient data is available to distinguish between bolt configurations. Table 5.14: Coupling strength of partially embedded and slipped Swellex pull tests Embedment Installation medium n Minimum Maximum Average Partial Rock or ore kn/m 263 kn/m 228 kn/m Full Sandfill or paste kn/m 66 kn/m 39 kn/m It is worth noting that 8 of the 12 slipped fully embedded Swellex bolts came from only 3 testing campaigns, suggesting that slip in fill generally occurs when there is an underlying issue with either bolt installation or the fill itself. It is acknowledged that for the fully embedded tests, there does exist an upper bound on the value of couple strength calculable, as it is not possible to determine for tests that do not slip. As such this should not be seen as representative of all bolts, while the partial embedment tests may be a better approximation. With an average coupling strength of 228 kn/m, the mechanical interlocking of an inflatable bolt with the rockmass seems to greatly contribute to the strength of this interface; as seen in Section 5.2, an FRS tends to reach its ultimate capacity (i.e. fully decouple) at 4 kn/m. Contrasting the number of tests where Swellex bolts slipped in the current database with the number of slipped tests in a pull test database previously assembled by Soni (2) shows a large difference in the proportion of slipped tests relative to what Soni terms as destructive or non-destructive tests. 12 out of 111 tests (excluding partial embedment tests and tests performed with malfunctioning equipment) in

85 Frequency CumulativepFrequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data 71 the database assembled for this thesis are known to have slipped. Soni s database included 173 slipped tests out of 34 entries in the data set (Soni, 2). This is a result of a fundamental difference in the definition and interpretation of a slipped test. Tests in which 15 mm of displacement or more was observed were classified as slipped by Soni (2). For the purposes of this thesis, a slipped test occurs if the pull tester is unable (and acknowledges their inability) to build further load on the bolt as it continues to displace at loads that do not suggest plastic deformation of the bolt. Soni is not necessarily incorrect; it is possible that slip is the displacement mechanism that contributes to the low stiffness calculated for the bolts in the database, but it would be slip comparable in observed effect to cone plough, only occurring as load increases rather than the sustained slip at constant load observed in the case of an FRS Characterization of Performance Metrics for Expandable Bolts Working Capacity Figure 5.31 shows the distributions of working capacity for the Pm12 and Mn12, with further information in Table Working capacity is measured to the nearest 1 ton (8.9 kn), as this is generally the resolution used for pull tests performed by Atlas Copco. In general, pull tests on the Pm24 and Mn24 bolts were verification tests, likely due to the large loads involved, and very rarely were the bolts loaded until they yielded. Load resolution for these tests was 2 tons (17.8 kn), further complicating the interpretation of the data. As such, an analysis on the working capacity of the Pm24 and Mn24 bolts was not performed. 13 1P P Mn12 6P Pm12 7 Cumulative 6 5 4P 4 3 2P 2 1 P WorkingpCapacityp(kN) Figure 5.31: Working capacities of Swellex Pm12 and Mn12 Table 5.15: Summary statistics of Swellex Pm12 and Mn12 working capacity Bolt n x s ĉ v Pm kn 5.5 kn.6 Mn kn 7.4 kn.8 Mn12 bolts appear to have a higher working capacity than the Pm12 bolts. This is attributed to one outlier and the low number of tests on the Mn12. As expected, the coefficients of variation are relatively

86 Frequency Cumulative3Frequency Frequency CumulativeNFrequency Chapter 5. Summary Statistics and Interpretation of Pull Test Data 72 low, as the working capacity is dependent on the material properties of the bolt. Stiffness Figure 5.32 shows the distributions of bolt secant stiffness for all variants, and Table 5.16 shows summary statistics of secant stiffness. 14 1% 7 1% Mn12 Pm12 Cumulative 8% 6% 4% Mn24 Pm24 Cumulative 8% 6% 4% 4 2 2% 2 1 2% Secant3Stiffness3(kN/mm) % SecantNStiffnessN(kN/mm) % (a) Pm12, Mn12 (b) Pm24, Mn24 Figure 5.32: Secant stiffness of Swellex variants Table 5.16: Swellex secant stiffness summary statistics Variant n x s ĉ v Skewness Kurtosis Cross-sectional area Pm kn/mm 2.71 kn/mm mm 2 Mn kn/mm 3.31 kn/mm mm 2 Pm kn/mm 7.22 kn/mm mm 2 Mn kn/mm 7.17 kn/mm mm 2 The average secant stiffness for the Pm24 and Mn24 bolts is significantly greater that of the Pm12 and Mn12 bolts. This corresponds to the larger cross-sectional area of the Pm/Mn24 configurations. However, the values of secant stiffness are well below what would be expected if the only displacement mechanism were to be axial deformation. Using a maximum bolt length of 2.44 m for the Pm12 and Mn12, a stiffness of 2.1 kn/mm would be observed for pure axial deformation spread evenly across the length of the bolt, and for the Pm24 and Mn24 with a maximum length of 3.6 m, a stiffness of 26.7 kn/mm would be observed. If a deformation model similar to that of an FRS is considered with a decoupling front marking the onset of steel deformation as postulated by Li & Stillborg (1999), higher stiffness should be observed considering the strength of the couple appears to be in the region of 2 kn/m (Table 5.14). Of the Pm12 and Mn12 bolts, none approached this stiffness, and only one Pm24 and one Mn24 bolt surpassed their respective minimum anticipated stiffness. Of the few values of tangent stiffness recorded (16 across all bolts), none reached their minimum anticipated values. This strongly suggests a displacement mechanism beyond axial elastic deformation is being mobilised. Within bolt sizes, the Pm and Mn variants seem to perform quite similarly to one another. It is also noted that the skew value of the Pm12 is significantly less than that of the other 3 bolts; the other distributions appear more log-normal in nature, although the sample size seems to be too small to be definitive. The Pm12 also has the lowest coefficient of variation. This may be the result of installation

87 Secant Stiffness (kn/mm) Secant Stiffness (kn/mm) Chapter 5. Summary Statistics and Interpretation of Pull Test Data 73 only in backfill, a relatively controlled substance when compared with the variety of rock masses that may be encountered in the Sudbury Basin. Figure 5.33 shows the secant stiffness for the different bolt types sorted by the medium in which they were installed, and Table 5.17 shows summary statistics. t-tests are performed assuming unequal variances (the properties of sandfill and of various lithologies are not assumed to be equally variable), and the p-value is calculated for a two-tailed test Sandfill Rock/Ore Pm12 Mn Sandfill Rock/Ore 35 3 Pm24 Mn (a) Pm12 and Mn12 (b) Pm24 and Mn24 Figure 5.33: Secant stiffness of Swellex sorted by installation medium Table 5.17: Secant stiffness summary statistics on Swellex sorted by installation medium Bolt Medium n x s ĉ v p Pm12 & Mn12 Sandfill kn/mm 2.5 kn/mm.34 Rock/ore kn/mm 3.3 kn/mm Pm24 & Mn24 Sandfill kn/mm 5.7 kn/mm.56 Rock/ore kn/mm 8.7 kn/mm As the Mn variants of Swellex are designed as yielding bolts, they are generally not installed in backfill, and very few of the Pm bolts tested were installed in rock or ore. The Pm12 bolts installed in sandfill appear to react with similar stiffness as the Mn12s installed in rock or ore; with a p-value of.524, the null hypothesis is not rejected. However, both the Pm24s and Mn24s installed in rock or ore seem to react significantly stiffer than the Pm24s installed in sandfill (from the t-test, p=.27). This discrepancy in observed effect of the installation medium may be due to a lack of data, be it volume or the testing of Pm12s exclusively in sandfill, and Mn12s only in rock and ore. Having observed that in the assembled database slipping failure of the entire length of a Swellex bolt only appears to occur in sandfill, it does appear that inflatable bolt performance in rock and in backfill is not equivalent. This is attributed to the difference in material properties between the two installation media. 5.7 Other Reinforcement Elements Not all reinforcement elements pull tested by Vale in Sudbury since 211 are to be discussed at length. Some rock bolts had insufficient data available to perform a more extensive analysis on performance. Their behaviour is discussed herein, but the discussion must be kept in the context of a small available data set.

88 Chapter 5. Summary Statistics and Interpretation of Pull Test Data Yield-Lok Two pull test campaigns were performed on Yield-Lok bolts by Vale s Sudbury operations between 211 and 214, both at Totten Mine. A total of 9 bolts were tested, one of which is discarded as the data is incongruous with the other tests. One campaign was a verification test, and bolts were not pulled to their working capacity (in the case of the Yield-Lok defined by the yield strength of the tendon), resulting in very limited data. A summary of the working capacity and secant stiffness of the bolts is shown in Table Table 5.18: 8 (2.44 m) Yield-Lok pull test result summary Campaign Working Capacity Secant Stiffness 148 kn 18.6 kn/mm Totten - August 4, kn 12.3 kn/mm 149 kn 17.6 kn/mm N/A 9.6 kn/mm N/A 17.6 kn/mm Totten - April 9, 213 N/A 36.6 kn/mm N/A 18.1 kn/mm N/A 16. kn/mm Although only three working capacities were recorded, they were all around 15 kn. The 213 testing campaign stopped testing at about 16 tons (142 kn) without observing yield. The expected stiffness of a 3/4, 8 (19 mm, 2.44 m) bar is 23.4 kn/mm. One test acts in a stiffer manner, although the bolt was unloaded midway through the test and then reloaded, which may have affected the result. The lower stiffness of the other tests suggests that there is some degree of movement of the upset head through the polymer in static conditions, although it is possible that this is the rock mass response being measured Fibreglass Rebar One testing campaign of 1 (25 mm), 66 (1.68 m) fibreglass rebar was undertaken at Creighton with different combinations of resin types, for a total of 8 tests. Although they were not tested until yield, laboratory tests were provided by the supplier (FiReP) on 5 samples 21 mm in diameter. The failure loads of these samples were between kn and kn, averaging kn, corresponding to an average ultimate tensile strength of 135 MPa. Unlike steel rebar, fibreglass rebar, or more specifically fibreglass reinforced plastic (FRP) rebar, does not exhibit large plastic displacements, but deforms elastically until sudden failure (Duthinh & Starnes, 21). As such, working capacity is essentially synonymous with ultimate capacity in the case of FRP reinforcement. Another aspect of FRP behaviour that is different from that of steel is a significantly lower elastic modulus. While the steels discussed in this thesis have elastic moduli of around 2 GPa, the laboratory testing of the FRP rebar resulted in an average elastic modulus of 57.7 GPa. For a 71 (1.8 m), 25 mm bar this would be equivalent to a stiffness of 15.7 kn/mm. 6 of the 8 pull tests performed exhibited consistent stiffness between 19.2 and 2.1 kn/mm (the two outliers were 18.3 kn/mm and 13.5 kn/mm). This shows that despite the variance in resins used, performance was generally consistent and load was being effectively transmitted to the resin. Having said this, most tests were only performed until around

89 Chapter 5. Summary Statistics and Interpretation of Pull Test Data tons (169 kn), with one test reaching the 25 ton (222 kn) capacity for the pump. This falls well short of the FRP rebar failure loads anticipated, so resin performance was not verified for higher loads DS Bolt One testing campaign was performed on 4 DS Bolts (now known as the VersaBolt; Lamothe, September 214). Comparable to the D-Bolt, it is a smooth bar punctuated by oval anchors along its length. 2.5 mm bolts were pulled at Totten on April 19 th, 213. Load and displacement data was recorded for 3 of these tests. Working capacity and tangent and secant stiffness are shown in Table It should be noted that all bolts exhibited a very gradual yielding behaviour in comparison to the D-Bolt. Table 5.19: DS Bolt campaign summary Bolt Working Capacity Tangent Stiffness Secant Stiffness kn 22. kn/mm 19.5 kn/mm kn 36.9 kn/mm 24.5 kn/mm kn 22.8 kn/mm 18.3 kn/mm The typical thread and bar yield strengths of the 2.5 mm DS Bolt are 125 kn and 138 kn respectively (courtesy of Mansour), indicating yield of the thread Other Expandable Bolts Although Swellex were by far the most tested expandable bolts, three other configurations were tested. One campaign of Jennmar s Midi (16 kn ultimate capacity) Python bolt was performed at Creighton in 21, although the only data recorded is the maximum load exerted on the bolt during the pull test. Additionally, one campaign of DSI s Omega Bolt (both 12 Tonne and 24 Tonne configurations) was also performed at Creighton in 213. The results of the pull tests are shown in Table 5.2. Table 5.2: Summary of expandable bolt campaigns not including Swellex Bolt Configuration n Embedment Working Capacity Secant Stiffness Python Midi 5 N/A N/A N/A Omega 24t 1 Full 178 kn 1.2 kn/mm Omega 12t 4 Partial N/A N/A Table 5.2 shows that the data collected for these bolts is of insufficient quality for the purposes of this thesis. As such, no further analysis was performed MD Bolt A campaign testing 14 MD Bolts, supplied by Sandvik, was conducted at Copper Cliff mine. Only peak load and displacement at that peak load were recorded. It was not apparent how peak load was defined in the report for this trial, potentially denoting slip, yield, failure or an arbitrary load at which the test was stopped. Additionally, without intermediate displacements recorded during loading, the behaviour of the bolt cannot be fully assessed from this dataset. As a result, the MD bolt was not investigated further.

90 Chapter 5. Summary Statistics and Interpretation of Pull Test Data Summary Performance metrics have different degrees of variability for the reinforcement elements discussed. Metrics that are dependant primarily on the material properties of the bolt itself (i.e. working capacity as defined by the yield strength of the steel) tend to have relatively low variance. Larger variance observed for other metrics (such as measures of stiffness, or ultimate capacities defined by frictional as opposed to mechanical properties) may be attributed to variable installation conditions as a result of a number of potentially influential factors, as will be discussed in Chapter 6. The ultimate capacity of an FRS appears to be independent of the nominal diameter of the bolt. FRS A and FRS B bolts perform similarly, with average ultimate capacities of about 4 kn/m. The stiffness of FA35 and FA39 bolts was found to be similar, although displacement is comprised of various degrees of limited slip and elastic deformation of the bolt. Consequently, stiffness as calculated in this thesis is a somewhat arbitrary descriptor of how an FRS behaves. The rebar rock bolts tested yielded in a relatively consistent manner, although stiffness is much more variable. In most cases, stiffness values are below the minimum that would be expected of an equivalent non-grouted steel bar. Investigating the unloading of the rebar shows that while the stiffness (in particular tangent stiffness) of the bolt may be proportional to the elastic deformation, it overestimates deformation by a factor of at least 2. Partial encapsulation tests show that even a limited bond length between the rebar and the rockmass has sufficient strength to surpass the yield strength of the rebar. For the cone bolts, five performance metrics were identified to describe behaviour. The bolt head is displaced with an initial stiffness while the cone remains anchored in the resin. The cone then begins to displace at the plough load, and a new linear relationship between load and displacement is developed. This was dubbed plough stiffness, and accounts for simultaneous bolt movement and deformation. This linear displacement behaviour occurs until the bolt tendon yields, at which point the test is usually stopped as large plastic deformation of the bolt tendon is induced. This sequence of events is not universal; a fraction of the pull tests show that the MCB33 may potentially yield before any significant plough occurs. Divergent behaviours appear to be characteristic of D-Bolt pull tests for the collected data set. Although relatively little data was available, it could be observed that stiffness of the bolt may be a function of the length between the point of load application and the first, second or third anchor on the bolt, or a value in between or even beyond any of the anchors. The anchors have the potential to prevent load transmission down the entire bolt length. However, movement of one or more anchors on the bolt results in stress transmission and a less stiff response. A large variety of expandable bolt test data was collected; bolts manufactured with different steels, coatings and capacities were tested at Vale s Sudbury operations. Limited data volume resulted in the analysis of four types of Swellex: the Pm12, Mn12, Pm24 and Mn24. These bolts also had several possible behaviours: the bolt may displace linearly with load, but may also stiffen, soften, or displace in a non-uniform manner. This resulted in variable stiffness, although the stiffness generally reflected the capacity of the bolt used; higher capacity bolts are made with thicker, larger tubes resulting in stiffer behaviour. As there exist four bolt configurations split between installation in backfill and rock, and such divergent behaviours are consistently observed, meaningful analysis is difficult. Bolts for which a small amount of data is available include the Yield-Lok, FRP rebar, the DS Bolt, one Python bolt configuration, two Omega Bolt configurations and the MD Bolt. Further discussion of these bolts in this thesis is limited.

91 Chapter 5. Summary Statistics and Interpretation of Pull Test Data 77 The interpretation of the pull tests was consistently hindered by displacements attributed to rock mass or surface support compression, which did not represent a reaction of the bolt or its anchoring mechanism. This led to stiffness calculations resulting in lower-than-anticipated values. This could be addressed by implementing an alternative method of measuring displacement, such as that described in ASTM D However, stiffness as measured in the creation of the database may still be used as a relative metric, albeit requiring slightly more interpretation. Table 5.21 summarizes the findings of this chapter. Note that in the case of the FRS, ultimate capacity as opposed to working capacity is shown, with capacity given per metre of anchorage length. In the case of the MCB33, the yield load is shown. Table 5.21: Summary of working capacities for all bolts pull tested Bolt type Bolt name n Working capacity s ĉ v FA kn/m 1.3 kn/m.26 FA kn/m 7.2 kn/m.19 FA kn/m 11.1 kn/m.28 FRS* FB kn/m 13.7 kn/m.34 FB kn/m 13.3 kn/m.35 FB kn/m 1.8 kn/m.29 All kn/m 11.7 kn/m.3 Grouted (Static) Rebar A (2 mm) kn 7.5 kn.6 Rebar B (2 mm) kn 7.4 kn.6 MCB33* kn 16.3 kn.14 D-Bolt (2 mm) kn N/A N/A Grouted (Yielding) D-Bolt (22 mm) kn N/A N/A Versabolt kn N/A N/A Yield-Lok kn N/A N/A Expandable Swellex Pm kn 5.5 kn.6 Swellex Mn kn N/A N/A Recognizing the limitations of the testing methods, the statistical analysis provides useful indicators of the working capacity of several bolt types in underground hard rock conditions. Chapter 5 focuses on how factors related to the installation of the bolt, the rock mass in which the bolt was installed in and the characteristics of the bolt itself affect the performance of reinforcement elements in underground hard rock mines.

92 Chapter 6 Factors Influencing Pull Test Performance The behaviour of rock bolts as recorded by a pull test is influenced by a number of factors, some of which may be specific to a particular type of bolt. These can be grouped into three categories: those pertaining to the bolt itself, those that are specific to the installation of the bolt, and factors associated with the rock mass in which the bolt is installed. As the behaviour and performance of different bolt types is dictated by different mechanisms, the influence of various factors are examined for each type of bolt individually. The nature, quantity, and quality of information recorded varied between rock bolt suppliers and was also dependent on the type of bolt being tested. Sufficient data was recorded only for an analysis on factors that affect the performance of FRS bolts, rebar rock bolts, and MCB33s. This chapter will describe these analyses. 6.1 Friction Rock Stabilizers The FRS was the most tested type of bolt in the database, and generally had the most thoroughly documented trials. This is likely due to the recognition that ultimate capacity is closely tied to parameters that dictate how tight the fit of the bolt in the hole is, and friction between the bolt and the rock mass. A total of 7 factors were closely examined: bolt length, method of installation, drive time, drill bit diameter, bolt diameter, geology and rock mass quality Influence of Length The ultimate capacity of an FRS is expressed in a load per unit length basis. As a continuously frictionally coupled bolt, the typical failure behaviour is slip, and the bolt s capacity is thought to be dependant on the cumulative shear strength of the bolt/rock mass interface over the entire length of the bolt. In general, a limited number of FRS lengths are found in the database. Only one length of the FA39 and FA46 were tested, and other bolts generally had a large majority of testing performed on a single length of bolt, indicating the maturity of ground support standards in use at the mines from which data was acquired. Analysis results are presented in Figure 6.1 and Table

93 Ultimate Capacity (kn) Chapter 6. Factors Influencing Pull Test Performance Anchorage Length (m) Figure 6.1: Relationship between ultimate capacity and length for the FA35 Table 6.1: Statistics on the relationship between ultimate capacity and length for the FA35 Anchorage length 1.37 m 1.52 m n 1 71 x kn kn s kn kn ĉ v t p.111 As only two treatments (lengths) are present in the FA35 data, a t-test was performed to compare their means. The critical t- and the p-values shown in Table 6.1 are for a one-tailed t-test assuming unequal variance (α =.5). As the p-value is.111, the null hypothesis is rejected, implying a significant difference between the two means. It is acknowledged that there was very little data for the bolt with 1.37 m of anchorage (1 data points from 2 testing campaigns), especially considering the wide range of installation conditions that may be encountered. This is addressed to some extent by assuming unequal variances when performing the t-test, but it should also be taken into account that there is only a.15 m difference in length. It seems unlikely that this relatively small additional length of the bolt/rock mass interface has a strength of 79 kn/m, while the entire interface length of the 1.37 m bolt has an average strength of 36 kn/m. The same statistical test (one-tailed, unequal variance, α =.5) is performed for the FB35. A third, intermediate length of bolt was pull tested, but with only four tests performed it was omitted from the analysis, and the t-test was performed on the 1.52 m and 1.83 m bolts (presented in Figure 6.2 and Table 6.2).

94 Ultimate Capacity (kn) Ultimate Capacity (kn) Chapter 6. Factors Influencing Pull Test Performance Anchorage Length (m) Figure 6.2: Relationship between ultimate capacity and length for the FB35 Table 6.2: Statistics on the relationship between ultimate capacity and length for the FB35 Length 1.52 m 1.83 m 1.68 m n x 63.4 kn 69.4 kn 4. kn s 22.6 kn 1.2 kn 17.4 kn ĉ v t p.49 The null hypothesis was rejected, albeit marginally, which once again suggests a length dependency of performance. Although the data set was larger for this test, it was still limited and the results of the t-test are difficult to interpret. In Figure 6.3 and Table 6.3, the equivalent graph and table for the FB39, the opposite trend seems to be present for the larger bolt diameter Anchorage Length (m) Figure 6.3: Relationship between ultimate capacity and length for the FB39

95 Chapter 6. Factors Influencing Pull Test Performance 81 Table 6.3: Statistics on the relationship between ultimate capacity and length for the FB39 Length 1.52 m 1.65 m 2.13 m n x 61.8 kn 56.9 kn 46.4 kn s 2.1 kn 1.4 kn 2.5 kn ĉ v Three lengths of FB39 were tested, so the t-test must be abandoned in favour of ANOVA. For the purposes of this analysis, single factor ANOVA using a fixed effect model was used. Single factor ANOVA on a fixed effect model is to be applied when one factor has been isolated. As the database is composed of different testing campaigns which have variable installation, rock mass and bolt parameters, it is very difficult to isolate a single factor when comparing two or more campaigns. As such, the results of the analysis presented in the Table 6.4 are to be interpreted cautiously, as the degree of confidence calculated is calculated under the assumption that a single factor is varied between the datasets. Table 6.4: Single factor ANOVA performed on the relationship between ultimate capacity and length for the FB39 Source of Variation SS df MS F p-value F crit Between Groups Within Groups Total A p-value of.32 was calculated, thus the null hypothesis was rejected. This analysis implies that the capacity of an FB39 in fact decreases with length, in direct contradiction to the findings of the analyses on the FA35 and FB35 datasets, the model proposed by Li & Stillborg, and a conventional understanding of friction. This shows the shortcomings of this type of analysis on data sets with high potential variability not necessarily captured by the low volume of data. As a result, the confidence in the findings of similar analyses on the length dependency of the FA35 and FB35 is affected. In order to address the problem of a lack of data volume, datasets composed of each bolt configuration were combined. Table 6.5 shows that all 1.52 m bolts were either 35 or 39 mm nominal diameter, and almost all 1.83 m bolts were 46 mm nominal diameter. As such, this analysis is only valid if the assumption is made that there is no difference in performance between different diameters and suppliers of FRS. Although from Section this does appear to be the case, it must be a considered a shortcoming of the analysis. The overall dataset is shown in Figure 6.4. Table 6.5: Breakdown of bolts contributing to major sets of anchorage length 1.52 m 1.83 m FA35 3.5% % FA % % FA46 % 31% FB % 12.7% FB % % FB46 % 56.3%

96 Ultimate/Capacity/pkNc Chapter 6. Factors Influencing Pull Test Performance 82 Distribution/p1/kN/m/binsc a 1a 2a 3a 4a 5a L=1.52m L=1.83m y/=/14.947x/h/ R²/=/ Anchorage/Length/pmc Figure 6.4: Relationship between ultimate capacity and anchorage length for all FRS bolts As seen in Figure 6.4, the coefficient of determination for the linear regression performed was very low. However, as the majority of the data is for bolt lengths of 1.52 m and 1.83 m, these two subsets were further analysed. The distributions for both are shown in Figure 6.4. The longer bolt does appear to generally have a larger load capacity as it not only has a larger mean, but the overall distribution occurs at higher loads. Table 6.6 directly compares the two data sets, including the results of a one-tailed t-test assuming equal variances (α =.5). Table 6.6: Comparison of 1.52 m and 1.83 m of anchorage length for all FRS bolts Anchorage Length n x s ĉ v t t crit p 1.52 m kn 18.7 kn m kn 18.6 kn The results of the t-test performed in Table 6.6 strongly suggest a difference between the mean ultimate capacities of the two FRS lengths. The low coefficient of determination calculated as part of the linear regression may be explained by the high variability in ultimate capacity. This limitation is overcome by the t-test through sheer volume of pull test results, allowing a thorough characterisation of distributions for two lengths of bolt. Tomory et al. (1998) do not discuss the relationship between capacity and FRS length. However, Tomory (1997) found no observable trends in an analysis that spanned a larger variety of bolt lengths than is presented here. This analysis was performed on mm Split Sets, over a much wider range of installation conditions (over 5 mines participated in the study), and it was hypothesized that a variety of factors obscured any potential relationship between bolt length and ultimate capacity (Tomory, 1997). Similar conclusions must be drawn for the unsuccessful linear regression presented in Figure 6.4. However, on both the bases of the analysis presented in Table 6.6, as well as existent theoretical justification (Li & Stillborg, 1999), a load per unit length basis of performance evaluation will be used for the remainder of the FRS analysis in this chapter.

97 UltimatedCapacityd(kN/m) Chapter 6. Factors Influencing Pull Test Performance Installation Method A parameter recorded on a relatively frequent basis in pull test reports was whether the FRSs were installed using a bolter or a jackleg. A large majority of pull test reports that note the method of installation were for testing campaigns in which a MacLean bolter was used to install the bolts (58.7% of pull tests were performed on FRSs installed with bolters, 7.7% on FRSs installed with jacklegs, and no method of installation was recorded for the remaining 33.6%). Pull tests performed on the FA35 are the exception. 31 FA35 bolts were explicitly noted as installed using jacklegs, compared to 44 FA35 bolts installed using a bolter. Figure 6.5 compares the two installation methods, and Table 6.7 compares statistics. A two-tailed t-test assuming unequal variances is performed (α =.5) Bolter-installed Copper Cliff 4/27/212 Copper Cliff 5/11/212 Totten 4/22/21 Stobie 5/28/214 1 Jackleg Bolter Figure 6.5: Comparison of ultimate capacities between jackleg and bolter installations of the FA35 Table 6.7: Ultimate capacity statistics for jackleg and bolter installations of the FA35 Jackleg Bolter n x 32.6 kn/m kn/m s 8.49 kn/m 7.47 kn/m ĉ v t 7.2 t crit 2.1 p There is a significant difference between the means of the two data sets, with bolter-installed FRSs exhibiting almost 5% more capacity than a jackleg-installed bolt. The jackleg bolts represent four testing campaigns in different conditions, and 9% of the bolter-installed bolts outperform 65% of the jackleg-installed bolts. The t-test strongly suggests that installation method is a critical factor in regards to FRS performance. In Figure 6.6 and Table 6.8, the dataset is expanded to compare installations of all FRS configurations.

98 Frequency Chapter 6. Factors Influencing Pull Test Performance 84 3% 25% 2% MacLean Jackleg 15% 1% 5% % Ultimate Capacity (kn/m) Figure 6.6: Comparison of ultimate capacities between jackleg and bolter installations of all FRS bolts Table 6.8: Ultimate capacity statistics for jackleg and bolter installation of all FRS bolts Jackleg Bolter n x 31.8 kn/m 39.8 kn/m s 7.83 kn/m kn/m ĉ v t 5.56 t crit 1.67 p. While relatively few bolts are added to the jackleg-installed dataset, many more bolter-installed bolts are incorporated into the analysis. The conclusions are similar, although Table 6.7 seems to overstate the difference in bolt performance. A bolter-installed FRS appears to have approximately 25% greater capacity than an FRS installed with a jackleg. This may be due to the stability that a bolter provides during drilling. A steadier drill would result in a smaller, more accurately drilled hole. This would in turn result in a tighter fitting FRS with higher radial stresses, and greater frictional resistance to pull. In addition to the productivity and safety benefits of mechanized bolting, it also appears to result in a higher quality FRS installations. In Section 5.2.3, it was found that the average FRS ultimate capacity was greater than that found by Tomory et al. (1998) for the Split Set (38.9 kn/m versus 31.9 kn/m). Tomory et al. (1998) state that the jackleg was commonly used for bolt installation, while for the database assembled for this thesis, bolter installation is much more common. In fact, comparing the average ultimate capacity of Split Sets pull tests collected by Tomory et al. (1998) and the average capacity of pull tests on jackleg-installed FRS in this thesis, very similar capacities are observed (31.8 kn/m in the current database versus 31.9 kn/m found by Tomory et al; 1998). As such, the discrepancy in average load values is attributed to

99 UltimateICapacityI(kN) UltimatenCapacityn(kN) UltimateICapacityI(kN) UltimateICapacityI(kN) UltimateICapacityI(kN) UltimatenCapacityn(kN) Chapter 6. Factors Influencing Pull Test Performance 85 the modern methods of bolt installation used at Vale s Sudbury operations Influence of Drive Time Drive time refers to the time it takes to fully insert an FRS into a borehole. The FRS is mounted on the head of the drill, and then hammered into the hole. Figure 6.7 shows the relationship between drive time and ultimate capacity for various FRS configurations. Data points regarded to be outliers are shown in red, and are excluded from the regression. As they are on the boundary of the data set, they are termed as having high leverage, (Fox, 28), and have the potential to greatly influence the regression. They are not seen as representative of the majority of FRS installations, and it is assumed that factors extraneous to a typical installation influenced the drive times of these bolts. Note that absolute values of ultimate capacity (kn) were used as it was assumed that a longer bolt will have a lengthier drive time yi=i1.4847xi+i38.47 R²I=I InstallationITimeI(s) yn=n.4436xn+n R²n=n InstallationnTimen(s) (a) FA35 (b) FB yi=i1.119xi+i R²I=I InstallationITimeI(s) yi=i2.4483xi+i25.55 R²I=I InstallationITimeI(s) (c) FA39 (d) FB yi=i1.5842xi+i4.991 R²I=I InstallationITimeI(s) yn=n1.272xn+n R²n=n InstallationnTimen(s) (e) FA46 (f) FB46 Figure 6.7: Relationship between drive time and ultimate capacity for all FRS bolts Most configurations of FRS appear to show a trend between installation time and ultimate capacity. The two exceptions are the FB35 and the FA39. In the case of the FA39, there are relatively few data points, and those present are concentrated in a narrow band of installation times. In the case of the

100 UltimatenCapacitynukNw Chapter 6. Factors Influencing Pull Test Performance 86 FB35, the low coefficient of determination appears to be due to different installation methods. Figure 6.8 distinguishes between the bolts known to have been installed with a MacLean bolter (black), and those installed with either a jackleg or with equipment not noted in their respective pull test report (red) yn=n2.2248xn+n44.69 R²n=n Bolter Jacklegnornunknown InstallationnTimenusw Figure 6.8: Relationship between drive time and ultimate capacity for FB35, distinguishing between installation methods Separating the MacLean-installed FB35 bolts results in a clearer relationship (R 2 =.493, compared to.3 for the full data set). This shows that although jackleg installed FRS bolts have lower load capacities, they require a similar amount of time for installation. Figure 6.9 and Table 6.9 show the results of analysing data only from pull tests known to have been installed with a MacLean Bolter. Table 6.9: Description of relationships between drive time and ultimate capacity for FRSs installed with a MacLean Bolter Variant n Average Ultimate Capacity Average Drive Time Regression Coefficient R 2 FA kn 19.5 s.86 kn/s.7 FA kn 11.7 s.98 kn/s.11 FA kn 16.7 s 1.53 kn/s.41 FB kn 11.9 s 2.25 kn/s.49 FB kn 15. s 3.5 kn/s.55 FB kn 18.1 s 1.8 kn/s.54

101 UltimateICapacityI(kN) Ultimate)Capacity)(kN) UltimateICapacityI(kN) UltimateICapacityI(kN) UltimateICapacityI(kN) UltimatenCapacityn(kN) Chapter 6. Factors Influencing Pull Test Performance yi=i.8646xi+i R²I=I InstallationITimeI(s) yn=n2.2248xn+n44.69 R²n=n InstallationnTimen(s) (a) FA35 yi=i1.119xi+i R²I=I InstallationITimeI(s) (b) FB35 yi=i3.52xi+i R²I=I InstallationITimeI(s) (c) FA39 (d) FB yi=i1.5291xi+i4.25 R²I=I InstallationITimeI(s) y)=)1.7974x)+)36.1 R²)=) Installation)Time)(s) (e) FA46 (f) FB46 Figure 6.9: Relationship between drive time and absolute ultimate capacity for all FRS configurations installed using a bolter From Table 6.9, it appears as though different bolt configurations have different degrees of sensitivity to the relationship between drive time and ultimate capacity; regression coefficients range from.86 kn/s for the FA35 to 3.5 kn/s for the FB39. This may be the result of an over-simplification of the relationship illustrated by performing a linear regression. Both the FB35 and FB39 have a large number of high leverage data points with installations less than 1 seconds. If the relationship were logarithmic or a power function as opposed to linear in nature (as is perhaps hinted for the FB35 and FA46 in Figure 6.9 and the fact that such a relationship should pass through the origin), data sets with concentrations of low values would have higher gradients than those without. Datasets of individual FRS configurations do not clearly capture this behaviour, and it is unclear whether these datasets may be combined. Figure 6.1 shows a power function fit to the data from all FRS configurations.

102 UltimatenCapacityn+kN) Chapter 6. Factors Influencing Pull Test Performance y = 1.481x R² =.3641 y = x.394 R² =.4217 FA35 FA39 FA46 FB35 FB39 FB InstallationnTimen+s) Figure 6.1: Relationship between installation time and ultimate capacity for all bolter-installed FRS fit with linear and power functions Figure 6.1 shows that a linear regression has a lower coefficient of determination than a power function fit to the same dataset, although neither correlation is very strong. The FB35 appears to develop high ultimate capacities with relatively low installation times. The FB39 has longer installation times for low ultimate capacity bolts, but the high regression coefficient results in relatively low installation times for high ultimate capacity bolts. The FB46 and FA46 display similar relationships between ultimate capacity and installation time. The FA35 develops high capacities at relatively low installation times, but has a low regression coefficient so only relatively long installation times exhibit higher capacities. Overall, it would appear that as trends between ultimate capacity and installation time are quite different between FRS diameters as well as suppliers, this relationship should be defined on the basis of an individual FRS configuration. A larger number of tests that cover a wider range of installation times for individual configurations are necessary in order to determine whether these trends may be characterised as linear or power functions. Tomory et al. (1998) fit linear functions to their dataset which, when extrapolated, intercepted the origin. This is not the case of the data presented in Figure 6.1, possibly due to the fact that Tomory et al. were primarily investigating bolts installed using a jackleg, while the data presented in Figure 6.1 is for bolts installed using a bolter. It is clear that there is a relationship between drive time and the ultimate capacity of an FRS. However, drive time is not an independent factor, but is just as dependent on installation conditions as the performance of the FRS itself. As such, its relationship with ultimate capacity is not particularly elucidating in terms of explaining how the performance of an FRS relates to its installation conditions. Having said this, tracking installation time could be a useful quality control tool. If the operator of the equipment installing the bolts consistently achieves drive times of less than 1 seconds, it may be prudent to implement a denser bolting pattern. Conversely, a 3 second drive time for any bolt suggests the bolt will perform well in a loading scenario analogous to a pull test Influence of Drill Bit Diameter The diameter of the drill bit is one of the factors that dictate the size of the hole drilled. Hypothetically, all other conditions remaining the same, one would expect to observe greater ultimate capacities with

103 UltimateDCapacityD(kN/m) UltimateDCapacityD(kN/m) UltimateDCapacityD(kN/m) UltimateDCapacityD(kN/m) UltimateDCapacityD(kN/m) UltimateDCapacityD(kN/m) Chapter 6. Factors Influencing Pull Test Performance 89 smaller drill bit diameters due to higher radial stresses induced in the bolt as a result of a tighter fit. Figure 6.11 shows the relationship between ultimate capacity and drill bit diameter for the different configurations of FRS R²D=D.93 4 R²D=D DrillDBitDDiameterD(mm) (a) FA DrillDBitDDiameterD(mm) (b) FB R²D=D R²D=D DrillDBitDDiameterD(mm) (c) FA DrillDBitDDiameterD(mm) (d) FB R²D=D R²D=D DrillDBitDDiameterD(mm) (e) FA DrillDBitDDiameterD(mm) (f) FB46 Figure 6.11: Relationship between drill bit diameter and ultimate capacity for all FRS configurations From Figure 6.11 there initially does not appear to be a substantive relationship between drill bit diameter and ultimate capacity, and the weak relationships that do exist for the smaller diameter bolts appear to be positive, defying the expectation of a negative relationship. The exception to this is the FB46, the data for which has the largest span in drill bit diameters tested. In order to further investigate this, two of the largest campaigns on pull tests in which the drill bit diameter was varied were examined, both on the FB46. These are shown in Figure 6.12.

104 UltimateDCapacityDgkN/m) UltimateDCapacityD(kN/m) UltimateDCapacityD(kN/m) Chapter 6. Factors Influencing Pull Test Performance CreightonD9/12/25 CreightonD4/28/28 R²D=D.2482 R²D=D DrillDBitDDiameterD(mm) Figure 6.12: Relationship between drill bit diameter and ultimate capacity for two testing campaigns performed on the FB46 Both of the campaigns presented in Figure 6.12 show negative relationships between drill bit diameter and ultimate capacity. This implies that while individual campaigns may reflect differences in bolt performance when a wide range of bit diameters are used, there are other factors that strongly influence the way bit diameter translates to hole size, and thus tightness of bolt fit and ultimate capacity. Results were filtered by installation method for further examination. There were no further findings for the FA39, FA46, FB39 or FB46 bolts. FA35 and FB35 pull tests (Figure 6.13) further clarify the relationship between ultimate capacity, drill bit diameter and installation method R²D=D.342 MacLeanDBolter Jackleg R²D=D.2354 MacLeanDBolter Unknown 2 R²D=D R²D=D DrillDBitDDiameterDgmm) 37. (a) FA DrillDBitDDiameterD(mm) 37. (b) FB35 Figure 6.13: Relationship between drill bit diameter and ultimate capacity for FA35 and FB35, separated by installation method Although no correlation between ultimate capacity and drill bit diameter is observed for bolts installed by a bolter in Figure 6.13a, bolts installed using a jackleg show a negative relationship between ultimate capacity and drill bit diameter (R 2 =.4411). In Figure 6.13b, no relationship is observed for bolts installed using an unknown method, but a weak relationship seen for bolts installed with a bolter (R 2 =.2354). The relatively low coefficient of determination calculated for either regression suggests that while ultimate capacity does appear to be influenced by the diameter of the drill bit, it is heavily influenced by other factors that obscure the relationship. These factors are likely associated with the translation of drill bit size to hole size, and could include the equipment operator s skill, or how the strength and quality of the rock mass influences the size of the hole drilled. It is thus difficult to describe

105 Bolt Diameter/Bit Diameter Chapter 6. Factors Influencing Pull Test Performance 91 a universally applicable model of FRS performance dependant solely on the diameter of the drill bit used Influence of Bolt Diameter There is insufficient variation in bolt diameter for specific FRS sizes for its effects to be reliably determined. However, an analysis is possible if bolt diameter and drill bit size are combined. A ratio between the diameters of the bolt and the bit used in the installation is calculated a higher ratio would indicate a tighter fit, should the drill bit accurately reflect the hole diameter. Figure 6.14 shows the nature of this relationship for the various bolts, and summary statistics provided in Table 6.1. The average ultimate capacities in Table 6.1 are calculated from the same data set used for the calculation of the bolt to drill bit diameter ratios FA35 FB35 FA39 FB39 FA46 FB46 Bolt Configuration Figure 6.14: Bolt diameter to drill bit diameter ratios for all FRS variants Table 6.1: Summary statistics for bolt diameter to drill bit diameter ratios for all FRS variants Variant n x s ĉ v Average Ultimate Capacity FA kn/m FB kn/m FA kn/m FB kn/m FA kn/m FB kn/m Figure 6.14 and Table 6.1 suggest that for the given data set, FRS A bolts are larger relative to the drill bit than the equivalent FRS B. However, it must be recognized that this is not representative of the entire database (only 228 of 545 FRS pull tests), and as seen in Section this does not necessarily translate directly into performance. Table 6.1 reinforces both of these points; although the FA35 and FB35 have quite different bolt to drill bit diameter ratios, the average ultimate capacities for the two data sets are similar. Conversely, the bolt to drill bit diameter ratios of the FA46 and FB46 are similar, but the FA46 dataset has a much higher average ultimate capacity. Only a comparison of the 39 mm nominal diameter bolts shows the expected effect. The FA39 dataset has a larger bolt diameter to drill bit diameter ratio, and a larger ultimate capacity than the FB39 dataset. Figure 6.15 shows the

106 UltimateoCapacityo(kN/m) UltimateDCapacityD(kN/m) UltimateDCapacityD(kN/m) UltimateDCapacityD(kN/m) UltimateoCapacityo(kN/m) UltimateDCapacityD(kN/m) Chapter 6. Factors Influencing Pull Test Performance 92 relationship between bolt to drill bit diameter ratio and ultimate capacity R²o=o R²D=D Bolto/oDrilloBitoDiameter (a) FA BoltD/DDrillDBitDDiameter (b) FB R²D=D.19 4 R²D=D BoltD/DDrillDBitDDiameter (c) FA BoltD/DDrillDBitDDiameter (d) FB R²o=o R²D=D Bolto/oDrilloBitoDiameter (e) FA BoltD/DDrillDBitDDiameter (f) FB46 Figure 6.15: Relationship between bolt diameter to drill bit diameter ratio and ultimate capacity Figure 6.15 appears to indicate that the relationship between the bolt diameter/drill bit diameter ratio and ultimate capacity is weak; coefficients of determination are generally low, and gradients are both positive and negative for different bolt configurations, when positive relationships would be expected. In an attempt to address this, as in Section 6.1.4, Figure 6.16 trims the data, sorting by test campaign in which multiple drill bit sizes are used (to increase the range of bolt diameter to drill bit diameter ratios covered). Linear regressions are shown for data sets with 1 or more data points.

107 UltimategCapacitygXkN/m) UltimatefCapacityf(kN/m) UltimatesCapacitys(kN/m) UltimatenCapacityn(kN/m) UltimateBCapacityBhkN/mO UltimatenCapacityn(kN/m) Chapter 6. Factors Influencing Pull Test Performance Totten,B4/22/21 Totten,B4/4/214 2 Creighton,B4/29/214BhOreO Creighton,B4/29/214BhNoriteO Totten,B9/23/21 R²B=B BoltB/BDrillBBitBDiameter 8 (a) FA Totten,n214/11/ Boltn/nDrillnBitnDiameter 1.18 (b) FB Coleman,n3/7/214 2 Garson,s9/3/213 Coleman,s4/2/214s(Ore) Coleman,s4/2/214s(Granite) Bolts/sDrillsBitsDiameter (c) FA39 Creighton,g6/1/214gXOre) Creighton,g6/1/214gXSubX) BoltgDiameterg:gDrillgBitgDiameter 1.18 (e) FA BoltnDiametern:nDrillnBitnDiameter (d) FB39 CopperfCliff,f214/1/31f(QuartzfDiorite) CopperfCliff,f214/1/31f(Ore) Totten,f214/12/ BoltfDiameterf:fDrillfBitfDiameter (f) FB46 Figure 6.16: Relationship between bolt diameter to drill bit diameter ratio and ultimate capacity for all FRS configurations by testing campaign Figure 6.16 shows that there is no clear relationship for the data sets with 5 or fewer entries, and only one data set (Totten, 4/22/21) shows a positive relationship between bolt to bit diameter ratio and ultimate capacity. This data set is the largest and also has the widest ratio range, so while it does make a case for the existence of a relationship when testing campaigns are evaluated individually, this is not validated with any other data in the database Geology Rock types in the pull test database broadly fall into five categories: ore, igneous rocks (norite, quartz diorite, granite, etc.), metamorphosed igneous rocks (greenstone, amphibolite and mafic gneiss), breccias (Sudbury and granite breccias), and metasedimentary rocks. Additionally, some pull tests were performed on bolts installed in sand fill. Figure 6.17 separates test results by these lithological categories. Drill bit diameter is indicated in this figure, with green indicating a smaller drill bit diameter, and red larger. Note that all bolts for which the data could be found are included, regardless of installation method.

108 UltimatekCapacityk(kN/m) UltimatekCapacityk(kN/m) UltimatekCapacityk(kN/m) UltimatekCapacityk(kN/m) UltimateUCapacityU(kN/m) UltimatekCapacityk(kN/m) Chapter 6. Factors Influencing Pull Test Performance LargeUdrillbitUdiameterU(relative)U SmallUdrillbitUdiameterU(relative) 4 2 Ore Igneous Metaigneous Breccia Metased Sandfill Ore Igneous Metaigneous Breccia Metased Sandfill 8 (a) FA35 8 (b) FB Ore Igneous Metaigneous Breccia Metased Sandfill Ore Igneous Metaigneous Breccia Metased Sandfill (c) FA39 (d) FB Ore Igneous Metaigneous Breccia Metased Sandfill Ore Igneous Metaigneous Breccia Metased Sandfill (e) FA46 (f) FB46 Figure 6.17: FRS ultimate capacity by lithology Despite ore in Sudbury being a generally weaker and softer material than the host rock (lower UCS and modulus of elasticity), it is difficult to see any category of lithology consistently over- or underperform relative to the others, potentially due to other factors or measures of rock mass characterization. Figure 6.18 shows ultimate capacity plotted against the UCS of the rock in which the bolt is installed. Performing linear regression on the UCS ultimate capacity data results in no obvious relationship for any bolt, except the FA35. The likely cause of this relationship, however, is the fact that the bolts installed with a jackleg prior to pull testing were all in the relatively strong host rock and not the weaker ore, resulting in a misleading relationship with rock strength. As part of Vale s quality control program, pull test campaigns covered more than one rock unit in some cases (usually ore and host rock). These serve to isolate the effect of the properties of the intact rock as much as possible. Usually, the same equipment, operator and drill bit (or set of drill bits) were used to install the bolts, and testing in one area reduces the variability of rock mass conditions between the two rock units. A direct comparison of the results from ore and waste rock is presented in Table 6.11.

109 UltimatePCapacityP(kN/m) Ultimate Capacity (kn/m) UltimatePCapacityP(kN/m) Ultimate Capacity (kn/m) UltimatePCapacityP(kN/m) Ultimate Capacity (kn/m) Chapter 6. Factors Influencing Pull Test Performance R²P=P R² = UCSP(MPa) 3 (a) FA UCS (MPa) (b) FB R²P=P R² = UCSP(MPa) 3 (c) FA UCS (MPa) (d) FB R²P=P R² = UCSP(MPa) UCS (MPa) (e) FA46 (f) FB46 Figure 6.18: Relationship between UCS and ultimate capacity Table 6.11: Comparison of average ultimate capacities for pull tests performed in ore and waste rock in the same campaign Campaign UCS ore Element n Ore n W aste x Ore x W aste FA kn/m 27.2 kn/m FA kn/m 34.6 kn/m Creighton, 5/16/ MPa FA kn/m 48.9 kn/m FB kn/m 3.2 kn/m FB kn/m 32.8 kn/m Garson, 11/24/ MPa FB kn/m 27.9 kn/m Stobie, 1/15/ MPa FB kn/m 31.2 kn/m Copper Cliff, 1/31/ MPa FB kn/m 28.3 kn/m Coleman, 4/2/ MPa FA kn/m 36.9 kn/m Creighton, 4/29/ MPa FA kn/m 43.9 kn/m Creighton, 6/1/ MPa FA kn/m 29.9 kn/m Copper Cliff, 3/3/ MPa FB kn/m kn/m FB kn/m 58.4 kn/m

110 Normalized Ultimate Capacity Chapter 6. Factors Influencing Pull Test Performance 96 In general, very little difference in ultimate capacity is observed between installations in ore versus waste, as 7 of the 13 campaigns listed show a difference of less than 2 kn/m between the average ultimate capacities installed in the two rock types. In order to further investigate, the data from each campaign is normalized with respect to the average ultimate capacity for the pull tests performed in waste, set to equal 1. This reduces variability across different testing campaigns, while preserving the contrast between the ultimate capacities measured in ore and waste rock within individual campaigns. Figure 6.19 and Table 6.12 show the results. Table 6.12 also shows the results of the reverse analysis, where the average result of the bolts installed in ore in a campaign are equal to 1. The t-tests performed assume unequal variances, and the t crit and t values are two-tailed (α =.5). Distribution % 1% 2% 3% 4% 5% 6% Ore Waste Figure 6.19: Distributions of FRS bolts installed in ore and waste rock normalized to the campaign average ultimate capacity for bolts installed in waste rock Table 6.12: Comparison of FRS bolts installed in ore and waste rock Medium n x s ĉ v t t crit p Ore Waste Ore Waste Table 6.12 shows that while pull test results for bolts installed in ore do appear to be more variable than those installed in the waste rock, there is on average no significant difference between their ultimate capacities immediately after installation. It should be noted that while the ore is generally weaker than the waste rock in Sudbury as was shown in Table 3.1, it is still quite strong compared to some rock types with UCS values between 91 and 17 MPa. The same investigation performed in rock weaker than the ore found in Sudbury may not yield the same results. Additionally, there are other factors that may influence how an FRS performs with time between the ore and the waste; as the ore is a sulphide, an FRS installed in the ore may corrode more quickly than one installed in other rock types. The nature of structures present in waste rock versus those in ore is another aspect of the rock masses to be considered, as well as difference in rock quality.

111 Ultimate Capacity (kn/m) Chapter 6. Factors Influencing Pull Test Performance Rock Mass Quality No standardized quantification of rock mass quality was routinely performed in areas where pull testing occurred. In some cases, however, pull test reports offered qualitative descriptors of rock mass quality, usually limited to one or two words, such as good or highly fractured. As such, two categories of rock quality are defined: Poor and Good. Poor encompasses descriptors such as fractured, broken, and poor quality, while Good is used for descriptors denoting competent ground. This is a vast simplification of rock mass characterization, but it must be recognized that the database includes reports from several manufacturers at six different mines, each with different personnel that may be contributing to the report using non-standardized language to describe the rock. Geotechnical data quality varies between mine sites; RMR is the most widely used classification scheme between them, but values of RMR are generally applied to a certain lithological unit on a mine-wide basis. Values are typically between 55 and 8, although may be as low as 45 and as high as 9 in some lithologies. Bieniawski (1989) distinguishes between Good and Fair rock at an RMR of 6; as such, it is assumed that testing performed in rock masses labelled as Poor have an RMR that falls below 6, and Good above 6. Only one FRS configuration, the FA46, had more than one testing campaign installed in both poor and good rock. The results of these campaigns are shown in Figure 6.2 and Table The t-test performed is two-tailed (α =.5) assuming unequal variances Poor Rock Mass Quality Good Figure 6.2: Ultimate capacities recorded for FA46 bolts installed in poor and good quality ground Table 6.13: Comparison of FA46 bolts installed in poor and good quality ground Ground n x s ĉ v t t crit p Good kn/m 12.3 kn/m.31 Poor kn/m 6.3 kn/m Performing the t-test results in the rejection of the null hypothesis, thus there appears to be a significant difference between the means of the data sets; FA46 bolts installed in good ground averaged 3% higher ultimate capacity than those installed in poor ground. Data, however, is limited, especially for those bolts installed in poor ground. To address this, the analysis is expanded to include all bolt

112 UltimateCCapacityCRkN/mM Chapter 6. Factors Influencing Pull Test Performance 98 configurations. Figure 6.21 shows the results. Bolts installed using a jackleg are omitted on the basis of the findings in Section 6.1.2, where it was shown that an FRS installed using a bolter has 25% greater capacity than one installed with a jackleg. Distribution R5 kn/m binsm, 1, 2, 3, 4, 5, 6, Garson, 9/2/213 Poor Good RockCMassCQuality: Poor Good Figure 6.21: Ultimate capacities recorded for FRSs installed in poor and good quality ground The highlighted pull tests performed in poor rock are from a campaign of FA39 bolts installed at Garson, 9/2/213, and are the obvious outliers in the data set. Further investigation showed that the report included a photograph of the bolts as installed, shown in Figure Although the bolts appear to be installed along a fracture, only Bolts 1 and 3 appear to be on the fracture on the surface of the excavation, and registered ultimate capacities of 49.8 kn/m and 29.9 kn/m respectively, two of the three lowest results of the campaign (Bolt 3 by a large margin). The surrounding rock mass appears to be intact, and the assessment of this particular testing area as having fractured rock, per the report, is questionable. As a result, Table 6.14 compares the dataset from good quality rock, and poor quality rock with and without this campaign included. Figure 6.22: FA39s pulled at Garson, 9/2/213

113 Chapter 6. Factors Influencing Pull Test Performance 99 Table 6.14: Comparison of all FRSs installed in poor and good quality ground Ground n x s ĉ v t t crit p Good kn/m 9.5 kn/m.24 Poor (including Garson, 9/2/213) kn/m 9.8 kn/m Poor (omitting Garson, 9/2/213) kn/m 5.4 kn/m If the Garson FA39 test is included, the p-value of.61 is calculated, resulting in a marginal t-test result; strictly speaking, the null hypothesis is not rejected if α =.5, and no difference is observed between the means at 5% significance. However if this campaign is omitted, very similar results are obtained as when only the FA46 is examined, indicating that the average ultimate capacity of an FRS installed in good quality rock is 3% larger than that of one installed in poor rock. A larger number of tests and less variance in the results of the bolts installed in good rock result in a very low p-value, suggesting a significant difference. As such, it is concluded a difference does exist in the performance of FRSs installed in poor versus good quality rock Summary of Investigation on FRS Pull Tests The pull tests performed on the FRSs were relatively well documented, with pull testing personnel recording various installation parameters more often than for any other type of rock bolt. This is likely due to the recognition that FRS load capacity is much more dependent on these factors than the load capacity of bolts that fail before being pulled out. The thoroughness of the reporting allowed for a detailed investigation into the parameters affecting the ultimate capacity of an FRS installed in situ. An analysis of the effect of length on ultimate capacity showed that while there does appear to be an influence, it is only apparent when comparing two large populations of data. This implies that other factors are more influential, as linear regression shows no relationship and analysis of individual bolt configurations yields contradictory results. There does appear to be a clear difference in the ultimate capacity of a bolt depending on whether its hole is drilled and it is installed using a jackleg or a bolter. Bolter installed bolts seem to have 25% higher capacities than bolts installed using a jackleg. Filtering by installation method also clarifies other relationships; for example, the coefficient of determination for the relationship between ultimate capacity and drive time is increased when only examining bolts installed by a bolter, showing a clear relationship. While drill hole diameter was seldom noted in the pull test report, drill bit diameter was more routinely recorded. Once again, installation method heavily influenced the analysis, and trends became more clear for some bolt configurations when a distinction was made between bolts installed with a bolter versus a jackleg, or examining individual campaigns that used multiple bits. Comparing the drill bit diameter to the measured bolt diameter resulted in inconclusive results, implying that drill bit diameter is not necessarily a precise descriptor of the size of the drillhole. Various lithologies are present at Vale s Sudbury Operations, although no trends were visible when subdividing these lithologies into several categories. For a more focused analysis, only campaigns where tests were performed in both ore and waste were compared, and showed no systematic difference in bolt capacity. However, when comparing bolts installed in different rock mass conditions, those installed in better quality rock appear to outperform those installed in worse quality rock. Similarities exist between the analysis presented and that of Tomory et al. (1998), as their objective

114 Chapter 6. Factors Influencing Pull Test Performance 1 was to investigate parameters that influence the performance of Split Sets, although there is relatively little overlap in findings. In Section 6.1.1, a significant difference in performance between FRS bolts of different lengths was found, which was not touched on by Tomory et al. (1998). While the emergence of the use of bolters allows for a comparison between installation methods, Tomory et al. (1998) did not perform an equivalent analysis, presumably due to lack of variation in installation equipment. While both this thesis and Tomory et al. (1998) examine drive time, this thesis found that it appears as though the relationship between bolt capacity and drive time is either non-linear, or inconsistent between different diameters and suppliers of FRS. Tomory et al. (1998) appears to have had greater success analysing the effect of drill bit diameter on performance, likely due to the larger dataset on fewer configurations. The findings presented in this thesis only show broad trends in the cases of some individual campaigns, and relatively little on a consistent basis. Similarly, results comparing different geologies (aside from ore versus waste rock) were inconclusive for this thesis, while the database assembled by Tomory et al. (1998) covered a wider range of geologies, and presented more conclusive results. However, no direct comparison between ore and waste installations of Split Sets was performed, for which there does not appear to be a difference in FRS ultimate capacity. Rock mass quality is another parameter investigated that seems to influence ultimate capacity which Tomory et al. (1998) did not discuss extensively. The factor that was thoroughly investigated by Tomory et al. (1998) but not covered in this thesis was capacity development with time. It was found that the capacity of the SS33 increased with time after installation. All pull tests in the database assembled for this thesis were conducted immediately after bolt installation, so this parameter could not be examined. 6.2 Rebar Rock Bolts As the working capacity of rebar is dependent on the properties of its steel as opposed to factors associated with its installation, the stiffness of the bolt s behaviour is primarily analysed. This may provide insight into how load attenuates down the length of the bar, and how effectively the grout transfers load from the bolt to the surrounding rock mass. A significant number of pull tests performed on rebar were partially encapsulated. In these tests one cartridge of resin (usually fast setting) was used to anchor the toe of the bolt, while two cartridges of inert resin are used to simulate mixing conditions without providing further anchorage. Although these tests should not differ from full encapsulation tests in terms of working capacity, a shorter length of rebar coupled to the rock mass may affect the stiffness of the bolt. In this section, the influence of length, encapsulation length, spin time, residence time in the ground and geology on the performance of rebar is examined Rebar Rock Bolt Length Depending on how load is distributed along a rock bolt, a longer bolt may appear less stiff in terms of load per displacement. However, if the length of bolt that is partially decoupled from the rock mass is the same for different lengths of bolt, a similar stiffness should be observed. Figure 6.23 investigates this effect for Rebar A and B.

115 Working Capacity (kn) Working Capacity (kn) Secant Stiffness (kn/mm) Secant Stiffness (kn/mm) TangentbStiffnessb(kN/mm) Tangent Stiffness (kn/mm) Chapter 6. Factors Influencing Pull Test Performance FullbEmbedment PartialbEmbedment BoltbLengthb(m) (a) Rebar A tangent stiffness Bolt Length (m) (c) Rebar A secant stiffness Bolt Length (m) (b) Rebar B tangent stiffness Bolt Length (m) (d) Rebar B secant stiffness Bolt Length (m) (e) Rebar A working capacity Bolt Length (m) 3. (f) Rebar B working capacity Figure 6.23: Rebar performance metrics by length It can be seen that for both suppliers rebar, high tangent and secant stiffness occur with the longer (2.4 m) rebar length. Rebar A data (unlike the Rebar B data) shows two data clusters for the greater length: one low stiffness, and one high stiffness. It is assumed that the longer Rebar B acts in a stiffer manner than its shorter counterpart is a result of the low number of tests collected, and is not an effect that would be systematically observed with further testing. The 2.4 m Rebar A had a lower working capacity than the 1.8 m rebar. This may be a result of a different steel and/or manufacturing plant or process used for the production of the longer bolt - these results come from 3 different testing campaigns at 2 mines, so it is not attributed to bolt or installation quality. 2.4 m Rebar B does have one low working capacity test, but the others are equivalent to results seen for the 1.8 m bolts Encapsulation Length As suggested in Section 5.3, the stiffness of rebar should be related to its fully coupled (i.e. grouted) length. Grout length may be quantified in terms of a ratio between the summed length of resin cartridges and the length of the rebar. Only 1.8 m rebar tests are discussed and are presented in Figure 6.24.

116 SecanthStiffnessh(kN/mm) SecanthStiffnessh(kN/mm) TangenthStiffnessh(kN/mm) TangenthStiffnessh(kN/mm) Chapter 6. Factors Influencing Pull Test Performance GrouthLengthh:hBolthLength (a) Rebar A tangent stiffness GrouthLengthh:hBolthLength (c) Rebar A secant stiffness GrouthLengthh:hBolthLength (b) Rebar B tangent stiffness GrouthLengthh:hBolthLength (d) Rebar B secant stiffness Figure 6.24: Relationships between stiffness and grout length : rebar length for Suppliers A and B Large variances and a lack of data between fully encapsulated bolts and bolts with less than one third of their lengths encapsulated makes an analysis of two data clusters more prudent. Table 6.15 compares stiffness calculated for full and partial encapsulation tests. Table 6.15: Comparison of partial and full encapsulation test statistics for Rebar A and B Supplier Stiffness Embedment n x s ĉ v Tangent Partial kn/mm 9.3 kn/mm.31 A Full kn/mm 13.7 kn/mm.34 Secant Partial kn/mm 9. kn/mm.31 Full kn/mm 13.7 kn/mm.33 Tangent Partial kn/mm 4.7 kn/mm.27 B Full kn/mm 5.2 kn/mm.2 Secant Partial kn/mm 4.2 kn/mm.28 Full kn/mm 9.6 kn/mm.44 There is a clear difference in tangent and secant stiffness between the partial and full encapsulation tests, presumably due to a smaller length of rebar coupled to the rock mass in the partial encapsulation tests. Interestingly, the partial encapsulation tests consistently achieve about 7% of the stiffness of the fully encapsulated tests, although they employ only 15% to 35% of the resin. To further investigate, the unloading stiffness of partially and fully encapsulated test are shown in Figure 6.25.

117 Unloading Stiffness (kn/mm) Chapter 6. Factors Influencing Pull Test Performance % 2% 4% 6% 8% 1% Grout coverage 1.8 m 2.4 m 1.8 m Maximum 2.4 m Maximum Figure 6.25: Unloading stiffness and grout length : rebar length for Rebar A Figure 6.25 shows the maximum stiffness that may be expected for 1.8 m and 2.4 m bolts, corresponding to grout coverage. This was determined by assuming uniform elastic deformation along the length of bolt not fully bonded to the rock mass, with no deformation occurring within the bonded section. At % grout coverage, the maximum stiffness is equivalent to the stiffness of the entire length of rebar, and at 1% grout coverage maximum stiffness is considered infinite. This also assumes that the length of resin in the hole is equivalent to the length of cartridge, as the cross sectional area of a 2 mm bolt (314 mm 2 ) plus a 28 mm resin cartridge (616 mm 2 ; combined total of 93 mm 2 ) is roughly equivalent to a 34 mm hole (98 mm 2 ) that would be drilled by a 32 to 33 mm drill bit. It is apparent that for the fully grouted bolts, significant load is distributed down the bolt and/or alternate sources of displacement greatly influence the results. It is also apparent that the partially encapsulated tests perform in a relatively stiff manner, even though these tests should also be subject to displacements extraneous to elastic deformation. This implies that resin may migrate up the bolt into the inert resin during spinning, resulting in bond length between the rock and the rebar significantly greater than the length of the cartridge Spin Time When any resin-grouted bolt is installed, it is spun during the installation process to mix and activate the resin. Under-spinning may result in a poor mix, and an incomplete chemical reaction leading to reduced resin strength. Over-spinning may induce fractures in the resin as it begins to set, which may also result in reduced performance. Figure 6.26 shows the relationship between spin time and various performance metrics for Rebar A. Limited data prevented a similar analysis of Rebar B.

118 TangentlStiffnessl(kN/mm)l Secant Stiffness (kn/mm) TangentEStiffnessE(kN/mm) Secant Stiffness (kn/mm) Chapter 6. Factors Influencing Pull Test Performance EmEFullEEncapsulation 1.8EmEPartialEEncapsulation SpinETimeE(s) (a) Tangent stiffness Spin Time (s) (b) Secant stiffness Figure 6.26: Relationship between stiffness and resin spin time for Rebar A As the majority of the data is concentrated between 8 and 1 seconds (per typical installation procedures), no clear trends are observed. There was no description of how the bolt was spun, i.e. how quickly the bolt was spun or inserted into the hole, or how long spinning continued after the bolt was fully inserted. A wider range of data would be expected to show a trend in stiffness, however this data set does demonstrate that there is some limited allowable deviation in spin time for which performance does not seem to be significantly altered Residence Time The large majority of the pull tests collected were performed on bolts that had been installed on the same day as they were installed. Occasionally, tests are performed on bolts that were installed before the date of testing. Figure 6.27 and Table 6.16 compare fully encapsulated Rebar A that were tested on the day of installation with those that were installed previously Daylofltest Previous Installation 1.8lm 2.4lm Day of test Previous Installation (a) Tangent stiffness (b) Secant stiffness Figure 6.27: Stiffness comparison of Rebar A installed on the day of testing versus previously Table 6.16: Statistics regarding residence time for Rebar A Stiffness Installation Length n x s ĉ v Day of installation 2.4 m kn/mm 49.9 kn/mm.65 Tangent 2.4 m & 1.8 m kn/mm 33.9 kn/mm.64 Previously installed 2.4 m kn/mm 25.2 kn/mm.56 Day of installation 2.4 m kn/mm 24. kn/mm.62 Secant 2.4 m & 1.8 m kn/mm 17.3 kn/mm.49 Previously installed 2.4 m kn/mm 11.2 kn/mm.59

119 Tangent Stiffness (kn/mm) Secant Stiffness (kn/mm) Tangent Stiffness (kn/mm) Secant Stiffness (kn/mm) Chapter 6. Factors Influencing Pull Test Performance 15 Recognizing that the dataset is limited, it appears that in some cases rebar stiffness may decay with time; bearing load and being subject to vibrations from blasting and seismicity may weaken the resin rebar bond, allowing a greater degree of stress propagation down the bolt. However, this should not be regarded as a universally applicable model. Only a very limited sample size was available, and the low stiffness of the previously installed rebar makes it tempting to assume that no significant shearing has affected the bolt s performance. Should shearing occur, higher stiffness could be observed due to locking of the rebar in place Geology Rebar bolts were pull tested in a variety of lithologies. Due to the relatively low number of tests, the only lithological distinction made is between ore and host rock. A comparison of stiffness is shown in Figure 6.28 and Table m Full 1.8 m Partial 2.4 m Full Ore Waste Ore Waste (a) Rebar A tangent stiffness (b) Rebar A secant stiffness Ore Waste Ore Waste (c) Rebar B tangent stiffness (d) Rebar B secant stiffness Figure 6.28: Stiffness comparison between lithologies for rebar It appears that Rebar A react in a more stiff manner in ore versus a stronger and stiffer host rock, although fully encapsulated 1.8 m bolts are conspicuously absent from the data set. Conversely, Rebar B react with a higher tangent stiffness in the host rock, although the difference in stiffness is less pronounced and is quite small in the case of the secant stiffness (based on a smaller sample size). This suggests that for Rebar A, stress is being transmitted further down the bolt if it is installed in host rock, resulting in more strain and lower stiffness. The results must be kept in perspective, with a limited number of tests performed and many potentially influential factors.

120 Chapter 6. Factors Influencing Pull Test Performance 16 Table 6.17: Comparison of stiffness across different lithologies Manufacturer Stiffness Configuration Rock Unit n x s ĉ v 1.8 m Full Ore Host kn/mm 13.7 kn/mm.34 Tangent 1.8 m Partial Ore kn/mm 8.5 kn/mm.25 Host kn/mm 6.5 kn/mm m Full Ore kn/mm 44.4 kn/mm.65 A Host kn/m 3.1 kn/mm m Full Ore Host kn/mm 13.7 kn/mm.4 Secant 1.8 m Partial Ore kn/mm 7.4 kn/mm.28 Host kn/mm 8.4 kn/mm m Full Ore kn/mm 18.5 kn/mm.51 Host kn/mm 4.9 kn/mm.41 Tangent 1.8 m Full Ore kn/mm 3.7 kn/mm.17 Host kn/mm 3.5 kn/mm m Partial Ore kn/mm 3.5 kn/mm.22 B Host kn/mm 5.3 kn/mm.28 Secant 1.8 m Full Ore kn/mm 11.9 kn/mm.53 Host kn/mm 8.1 kn/mm m Partial Ore kn/mm 3.7 kn/mm.28 Host kn/mm 4.4 kn/mm Summary of Investigation on Rebar Rock Bolt Pull Tests Stiffness was the primary metric of study for rebar rock bolts, as various parameters have the potential to influence how effectively the resin bonds the rebar to the surrounding rock mass. Insufficient data existed in the database to verify the relationship between rebar length and stiffness, and the observed effect (length and stiffness being positively correlated) is likely a result of other factors. Comparing the lengths of rebar encapsulated in resin also yielded surprising results; while fully encapsulated bolts behaved in a manner less stiff than expected (although this may be associated with the test apparatus and method as oppose to the bolt itself), the partially encapsulated bolts reacted to load with a greater than expected stiffness. This implies that the bond length between the rebar and the rock mass is significantly longer than the length of the cartridge containing the chemically active resin. As standard operational procedures were followed for the installation of rebar, only a narrow band of spin times were recorded (all times within 4 seconds of each other). With such a narrow data set, no relationship was found, although more variable spin times would presumably influence results to a greater degree. Pull tests were on some occasions performed on rebar that had been installed prior to the date of testing. These rebar appeared to behave in a less stiff manner, but it is emphasized that the data set used to draw these conclusions is very small. Generally, rock mass compression posed a major challenge to analysing bolt stiffness. Rebar is the stiffest reinforcement element discussed in this thesis in terms of secant stiffness, so small movements of the test rig on the scale of a millimetre may have a tangible influence on the results. As such, to conduct a more substantial analysis, a method of measuring displacement as described by ASTM D that measures only bolt head displacement and excludes rock mass compression is recommended.

121 PloughyStiffnessy(kN/mm) Secant Stiffness (kn/mm) Initial Stiffness (kn/mm) Plough Point (kn) Chapter 6. Factors Influencing Pull Test Performance Modified Cone Bolts Five metrics by which cone bolts may be evaluated from a pull test have been defined: plough point, initial stiffness, plough stiffness, secant stiffness and yield load. Of these five, the first four depend directly on how the bolt displaces (or does not displace) through the resin, while the yield load depends on the properties of the steel. Residence time, resin type and geology are examined as potentially influential factors. Additionally, correlations between the five metrics are analysed Residence Time A number of pull tests were performed on Modified Cone Bolts installed before the date of pull testing. Figure 6.29 and Table 6.18 compare bolts installed the day of testing with those that were installed previously. The t-tests performed are two tailed, assuming unequal variances (α =.5) Day of test Previous Installation Day of test Previous Installation (a) Initial stiffness (b) Plough point Dayyofytest Previous Installation Day of test Previous Installation (c) Plough stiffness (d) Secant stiffness Figure 6.29: Performance comparison of MCB33s installed prior to and on the day of testing Table 6.18: Performance comparison of MCB33s installed prior to and on the day of testing Metric Installation n x s ĉ v t t crit p Initial Stiffness Day-of kn/mm 3.2 kn/mm.32 Previous kn/mm 8.2 kn/mm Plough Point Day-of kn 21. kn.34 Previous kn 23.9 kn Plough Stiffness Day-of kn/mm.69 kn/mm.35 Previous kn/mm 1.56 kn/mm Secant Stiffness Day-of kn/mm 2.63 kn/mm.66 Previous kn/mm 2.6 kn/mm

122 Chapter 6. Factors Influencing Pull Test Performance 18 MCB33 behaviour appears to evolve with time. The bolts seem to plough at lower loads, although with higher plough stiffness as indicated by the two relevant t-tests indicating a significant difference in means (p =.32 and p =.28 respectively). Note that of the 14 plough point measurements for the bolts installed before the test date, 5 of them have noted plough points of 26.7 kn. This is an upper bound, as it represents the pre-load, and it is possible that plough may have already occurred at loads below this. It should also be noted that the average initial stiffness of the two installation periods is very similar, implying that the difference in plough stiffness is not due to factors such as the rock mass compressing. This difference in behaviour could be a result of exposure to seismicity or blasting that would have occurred between installation and testing of the bolts. Resin already damaged by vibrations in the rock mass would offer less resistance to the onset of a consistent plough response. Resin damage may also be the result of shearing of the bolt. When sheared, cone bolts convert a portion of the shearing load into axial (Gaudreau et al, 24), potentially initialising the ploughing process. Simser et al. (26) found that shearing of the rock mass can pinch and lock cone bolts in place, resulting in a much stiffer form of reinforcement. This may be reflected in the higher initial and plough stiffnesses. Of the 9 previously installed bolts, only 1 did not plough, indicating that the locking in these circumstances was generally not significant enough to prevent any movement of the bolt, but did appear to affect the displacement mechanism. Alternatively, it is also possible that the properties of the resin enveloping the bolt evolved with time, modifying the behaviour of the bolt as it ploughs. A significant shortcoming of this analysis is that the length of time that the previously installed bolts had been in the ground relative to one another is unknown. Further investigation could quantify how performance changes with time, as well as clarify what the underlying cause in the change of behaviour is Geology The Modified Cone Bolts were installed and tested in various lithologies. Figure 6.3 and Table 6.19 show how properties vary between bolts installed in ore, igneous/metaigneous and metasedimentary rocks. Previously installed bolts are highlighted in red, and omitted from statistical calculations. Table 6.19: Performance comparison of MCB33 bolts installed previously and on the day of testing Metric Lithology n x s ĉ v Ore kn/mm 2.4 kn/mm.38 Initial Stiffness Ign/Metaign kn/mm 3.8 kn/mm.34 Metased kn/mm 1.9 kn/mm.2 Ore kn 16.8 kn.25 Plough Point Ign/Metaign kn 22.3 kn.49 Metased kn 17.2 kn.28 Ore kn/mm.46 kn/mm.33 Plough Stiffness Ign/Metaign kn/mm.92 kn/mm.43 Metased kn/mm.74 kn/mm.35 Ore kn 4.4 kn.3 Secant Stiffness Ign/Metaign kn/mm 3.62 kn/mm.85 Metased kn/mm 3.4 kn/mm.76

123 PloughaStiffnessa(kN/mm) Secant Stiffness (kn/mm) InitialMStiffnessM(kN/mm) Plough Point (kn) Chapter 6. Factors Influencing Pull Test Performance DayMofMInstallation PreviouslyMInstalled Ore (Meta)Igneous Metaseds Ore (Meta)Igneous Metaseds (a) Initial stiffness (b) Plough point Ore (Meta)Igneous Metaseds Ore (Meta)Igneous Metaseds (c) Plough stiffness (d) Secant stiffness Figure 6.3: Performance comparison of MCB33 bolts installed in ore, igneous/metaigneous and metasedimentary lithologies While all comparisons between ore and host rock must be approached cautiously as only one MCB33 test campaign was undertaken in ore, it would appear that these bolts acted in a less stiff manner and with a higher plough point. It is interesting to note that this is the precise opposite effect of what was observed for the bolts installed on a date before the pull test, characterised by a low plough point and high stiffness. Without further data, it is difficult to posit an explanation of this behaviour, or to assess its reproducibility Inter-variable Relationships The five response variables described are not independent; secant stiffness is a function of the other four factors. As such correlations will be found between this factor and the others, but this is regarded as mathematically obvious and not particularly relevant to this thesis. No relationship is found between any of the other four factors, except for one, shown in Figure Although no trend is seen between plough load and plough stiffness for bolts immediately after being installed, a relationship seems to develop between the two factors as time progresses. As discussed in Section 6.3.1, this may be due to shearing of the bolt initialising plough (resulting in a lower plough point during testing), while simultaneously stiffening the system, or evolving properties of the resin. The low coefficient of determination is likely a result of the varying conditions to which the bolts in question were exposed. These not only include installation conditions, but also the time-related effects, such as ground movement or energy input from seismicity and blasting activities. While the analysis in this thesis is far from exhaustive, Figure 6.31 shows that not only static performance, but also mechanisms and correlations for the MCB33 are subject to change with time.

124 Plough/Stiffness/rkN/mmv Chapter 6. Factors Influencing Pull Test Performance Day of/installation Previously installed 4 R² = R² = Plough Load/rkNv Figure 6.31: Relationship between plough stiffness and plough point for the MCB Summary of Cone Bolt Findings A comparison of the performance of MCB33s installed on a date prior to pull testing with those installed on the date the pull tests was made. It appears as though as time elapses, plough may be initiated at lower loads, but the stiffness of the combined plough/elastic deformation response increases. This timedependent behaviour also appears when examining correlations between different metrics of performance; plough load and plough stiffness correlate with one another for the bolts that were installed before the test date, but not for those that were tested on the date of installation. The relationship between the plough load and plough stiffness metrics was also apparent when contrasting geologies. MCB33s installed in ore appear to plough in a less stiff manner, but with a higher plough point than those installed in waste rock. Although data is too limited to be conclusive on the influence of geology, it is interesting to note that these two metrics once again appear to be linked. Like the rebar rock bolt, metrics of stiffness (along with plough point) were the primary objectives of investigation of this chapter, the reasoning is very different. Stiffness is examined for the rebar to understand how load is distributed along the bolt, while load is assumed to be evenly distributed along the tendon of an MCB33. The metrics examined were indicators of how the cone and the resin interacted with each other, and what appears to govern their behaviour in quasi-static loading conditions. It is not clear how these results translate into the dynamic performance of a cone bolt. Gaudreau et al. (24) found that after cyclic dynamic loading until failure of two cone bolts in laboratory conditions, one had a cumulative plough of 127 mm through the resin grout, and the other 177 mm of plough. Even if a cone bolt were to be pulled until failure as opposed to yield, these plough displacements would not be observed in static conditions (Simser et al, 26). As such, it is difficult to assess from quasi-static testing whether a stiffer bolt response due to the cone not ploughing or ploughing at a reduced rate would result in a significantly different energy capacity of the bolt in dynamic scenarios, although it may be an effect worth investigating.

125 Chapter 6. Factors Influencing Pull Test Performance Summary The findings of this Chapter are summarized in Table 6.2. The reinforcement elements are shown with the parameter that influences performance, as measured by a specified performance indicator. The nature of the relationship is described, along with its strength and confidence in the findings. Generally, a weak relationship and low confidence is due to a lack or narrow range of data, while a strong relationship and high confidence is indicative of a large amount of data clearly showing a certain effect. Table 6.2: Summary of observed relationships of between various factors and performance indicators for each rock bolt type Rock Bolt Parameter Performance indicator Relationship Strength Confidence Nominal diameter Ultimate capacity None None High Length Ultimate capacity Positive Weak Medium Installation method Ultimate capacity Higher with bolter, Strong High lower with jackleg FRS Drive time Ultimate capacity Positive Strong High Drill bit diameter Ultimate capacity Negative Weak Medium Lithology Ultimate capacity None None Medium Rock mass quality Ultimate capacity Higher in good, Strong Medium lower in poor Length All indicators None None Medium Unloading stiffness Tangent stiffness Positive Strong High Embedment length Tangent stiffness, Positive Medium High secant stiffness Rebar Spin time All indicators None None Low Residence time Tangent stiffness, Negative Weak Low secant stiffness Lithology All indicators None None Medium Plough stiffness Plough load Negative (previously Medium Medium installed bolts) Residence time Plough load Negative Strong High MCB33 Plough stiffness Positive Strong Medium Lithology Plough load Higher in ore, Weak Low lower in waste Plough stiffness Lower in ore, Weak Low higher in waste Only the following relationships met the criterion of high confidence: FRS performance is independent of nominal diameter. An FRS has a higher ultimate capacity if installed with a bolter. An FRS has a higher ultimate capacity if it takes longer to install. The tangent stiffness of a rebar rock bolt correlates positively with its unloading stiffness. The stiffness of a rebar rock bolt correlates positively with grout coverage. The plough load of an MCB33 is lower for a bolt installed prior to the date of testing. This chapter has discussed performance metrics in terms of working capacity and various measures of stiffness, although these are single values representing potentially complex responses to load. In Chapter 7, the displacement response of rock bolts subject to a pull test is described in greater detail.

126 Chapter 7 Characterization of Rock Bolt Behaviour Having investigated rock bolt response to load and the influence of various parameters, this chapter describes bolt behaviour observed during a pull test, defining a distribution of displacement development with load. Bolt performance is compared to manufacturer specifications and recommendations, and to that of other bolt types. 7.1 Characterisations of Bolt Behaviour using Laboratory Pull Tests Comparisons of the behaviours and capacities of various rock bolts subject to laboratory testing have been published previously (Stillborg, 1993; Li et al, 214). Stillborg (1993) conducted a series of tests on various rock bolts, namely mechanical bolts, 2 mm rebar grouted in both cement and resin, resingrouted 22 mm fibreglass rods, 39 mm Split Sets, and EXL Swellex bolts. Figure 7.1 shows the results, and is a widely used reference in the design of underground support systems. Note that Figure 7.1 is a composite assembled by Hoek et al. (1995), who labelled the horizontal axis deformation. Stillborg (1993) distinguished between rock bolt displacement and deformation. Displacement induced during tests on rebar was attributed solely to deformation of the bolt. However, bolts including the SS39, Swellex EXL and expansion shell anchored bolt exhibited slip, so in these cases, the horizontal axis was labelled as displacement. Throughout this thesis, behaviour has been expressed in terms of displacement, as it was difficult to interpret whether displacement measured by the pull test apparatus was solely attributable to bolt deformation. Stillborg s (1993) testing program consisted of bolts 3 m in length installed across two 1.5 m blocks of high strength (UCS = 6 MPa) concrete. Holes were drilled using a percussion drill in an attempt to recreate the roughness of holes drilled in the field. 3 samples of each bolt were tested across a simulated joint by moving the two concrete blocks apart. The load-displacement behaviours shown in Figure 7.1 are the average of the results from the three tests (Stillborg, 1993). Li et al. (214) performed a review of more rock bolts tested using a similar apparatus at the Norwegian University of Science and Technology (NTNU). Bolts were installed in two 95 mm x 95 mm x 95 mm high-strength concrete 112

127 Chapter 7. Characterization of Rock Bolt Behaviour 113 blocks. Hydraulic jacks exerted a load which pushed the two blocks apart as displacement was measured (Li et al, 214). Doucet & Voyzelle (212) present the results of laboratory static pull tests performed at the CANMET-MMSL facility using the procedures and apparatus described in ASTM D741-8, discussed in Section Instead of concrete blocks, rock bolts were installed in a steel tube, and load applied to the bolt head, as in an in situ pull test. The results presented by Doucet & Voyzelle (212) are from tests performed between 26 and 211, and include the behaviours of a mechanical bolt, resin grouted rebar, the MCB33, the discontinued MCB38, the Durabar Yieldable Bolt, the discontinued Roofex Rx8D and Rx2D, 2 mm and 22 mm D-Bolts, the Yield Lok Dynamic and the Yield Lok Static. Figure 7.1: Average load displacement behaviours obtained by Stillborg in a laboratory setting (Stillborg, 1993; composited by Hoek et al, 1995) The use of a database of in situ pull tests has significant advantages over these laboratory campaigns as tools for design. The bolts in a database of in situ pull tests are installed in conditions found on a mine site. Instead of installing bolts using lab-scale equipment in blocks of high strength concrete or steel tubes, they are installed in holes drilled in rock by equipment and procedures used daily on site. In addition, the database assembled is a much larger dataset than the laboratory tests discussed. Figure 7.1 was made using 3 tests for each bolt type (Stillborg, 1993). Static test results presented by Doucet & Voyzelle (212) appear to be results from single tests. With more tests performed, a distribution of behaviour may be assembled and can be used for probabilistic design and analysis. There are two primary advantages to the laboratory tests discussed. The first is that bolts are

128 Chapter 7. Characterization of Rock Bolt Behaviour 114 pulled until failure. The second is the test configuration. Load may be applied across a simulated joint (as in Stillborg, 1993, and Li et al, 214), which is arguably a better analogue of the nature of loading a reinforcement element may be subject to in practice, rather than loading of the bolt head (as is performed during an in situ pull test). The test may be instrumented differently to provide more certainty of the results and interpretation. For example, displacement at the toe of the bolt may be measured to determine whether the bolt is slipping or if all displacement may be attributed to bolt deformation. Such a measurement cannot be easily performed in an in situ pull test. Further discussion of testing methods and how they relate to one another may be found in Chapters 2 and 4. Previously assembled pull test databases (Tomory et al, 1998; Soni, 2) have primarily analysed the performance of rock bolts in terms of load capacity. Tomory et al. (1998) did not include displacement of the Split Set in their analysis. Soni (2) used the displacement measured during the pull test of a Swellex bolt to determine whether a bolt had slipped or not, thus as a binary metric of performance as opposed to a continuous one. Displacement characteristics of rock reinforcement and ground support elements in general are critical to enact some methods of design, such as those that incorporate the ground reaction curve (Section 4.3.3). In this chapter, load-displacement behaviours of certain rock bolts are compiled across all pull testing campaigns in the database in order to define a distribution of behaviour for that bolt. The objective is to propose composite diagrams of load-displacement behaviours of various types of rock bolt as undertaken by Stillborg (1993), using a larger set of in situ pull test data. 7.2 Friction Rock Stabilizers Displacement data was recorded for 25 FA35 and 24 FA39 bolts. Using this data as well as the ultimate capacity measurements taken for all FRS bolts in the database, performance metrics and the behaviour of the FRS were characterised. All FA35s and FA39s for which displacement was measured were 1.68 m in length and were tested without a pre-load Characterization of FRS Performance Table 7.1 shows the 1 th, 25 th, 5 th, 75 th and 9 th percentiles of ultimate capacity and secant stiffness for the 49 FRS bolts for which displacement was recorded. Note that ultimate capacities are not normalized to anchorage length. Table 7.1: FA35 and FA39 performance percentiles Metric Bolt n P 1 P 25 P 5 P 75 P 9 Ultimate FA kn 58. kn 64.6 kn 71.4 kn 88.6 kn Capacity FA kn 51.2 kn 57.3 kn 69.7 kn 79.4 kn Total kn 52.8 kn 62. kn 7.7 kn 83.2 kn Secant FA kn/mm 6.9 kn/mm 15. kn/mm 22.7 kn/mm 39.7 kn/mm Stiffness FA kn/mm 5.9 kn/mm 9.5 kn/mm 13.5 kn/mm 18.8 kn/mm Total kn/mm 6.5 kn/mm 1.9 kn/mm 18.2 kn/mm 28.2 kn/mm The percentiles of performance metrics used for the two bolts are not identical. Although the 1 th and 75 th percentiles are quite similar between the two bolts, the FA35 dataset appears to generally have a higher ultimate capacity. The FA35 also appears to behave in a more stiff manner than the FA39, the disparity between the two bolts increasing at higher percentiles. However, it has been demonstrated in

129 Loadt(kN) Chapter 7. Characterization of Rock Bolt Behaviour 115 Chapter 5 that there is no difference in ultimate capacity between the two bolts when the entire dataset of ultimate capacities is examined, and that secant stiffness is not necessarily reflective of performance. As such, the two datasets are combined in Figure 7.2. In this figure, the data points represent the combination of load and displacement at which the ultimate capacity of a bolt was reached. Boxes are defined by percentiles of both ultimate capacity and secant stiffness. Median ultimate capacity and secant stiffness are indicated by solid black lines. 1 9 P P P 9 P 75 P 5 P 5 P 25 P 1 P 25 P 1 FA35 FA Displacementt(mm) Figure 7.2: Distributions of secant stiffness and ultimate capacity for a pull test on an FRS with an anchorage length of 1.52 m In the construction of Figure 7.2, not all data regarding the ultimate capacity of an FRS was used. Figure 7.3 and Table 7.2 compares the dispersion of ultimate capacities for all configurations of FRS discussed in this thesis, including the FA35/FA39 dataset for which displacements are measured, and the FRS dataset as a whole. As multiple lengths of bolt are represented, load is quantified in terms of kn/m of anchorage length. Table 7.2: Distribution of ultimate capacity across all FRS configurations Bolt n P 1 P 25 P 5 P 75 P 9 FA kn/m 32.1 kn/m 4.9 kn/m 47.6 kn/m 52.4 kn/m FA kn/m 33.2 kn/m 36.9 kn/m 41.5 kn/m 51.1 kn/m FA kn/m 32.3 kn/m 38.4 kn/m 46.2 kn/m 53.5 kn/m FB kn/m 32.1 kn/m 39.9 kn/m 49.6 kn/m 58.4 kn/m FB kn/m 29.2 kn/m 35. kn/m 49.6 kn/m 57.8 kn/m FB kn/m 29.2 kn/m 38.9 kn/m 46.2 kn/m 51.1 kn/m FA35+FA kn/m 34.6 kn/m 4.7 kn/m 46.4 kn/m 54.6 kn/m Grand Total kn/m 31.5 kn/m 38.7 kn/m 47.4 kn/m 53.5 kn/m

130 Loadt(kN/m) LoadV(kN/m) Chapter 7. Characterization of Rock Bolt Behaviour Median Mean P 9 P 75 P P 25 P 1 1 FA35 FA39 FA46 FB35 FB39 FB46 FRSVVariant Figure 7.3: Ultimate capacity per metre distributions for all FRS configurations Percentiles of ultimate capacity generally agree with each other across bolt type. The main exception is the FA39, the capacity of which appears to be relatively narrowly distributed between 3 and 4 kn/m. The FB35 and FB39 have the widest distributions, while those of the FB46, FA35 and FA46 are similar. The nature of the FA39 dataset carries over into the combined FA35/FA39 dataset. P 1 and P 25 are higher than the equivalent percentiles for other bolts, although the median, P 75 and P 9 better match the rest of the data. In Figure 7.4, percentiles of secant stiffness and ultimate capacity are generalized so they may be applied to various lengths of FRS. Load capacity is expressed in kn/m, displacement in percentage of anchorage length, and percentiles of ultimate capacity shown are for the entire FRS dataset. 7 6 P 9 5 P P 9 P 75 P 5 P 5 P 25 P 25 P 1 P Displacementt(9toftanchoragetlength) Figure 7.4: Distribution of secant stiffness and ultimate capacity for all FRS bolts tested

131 Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour 117 To use Figure 7.4 for design, a decision must be made regarding what percentiles of performance (i.e. ultimate capacity and secant stiffness) are to be used. As an example, a designer wishes to use the 25 th percentiles of both ultimate capacity and secant stiffness of an FRS with an anchorage length of 2 m. The intersect of the 25 th percentiles is at 33 kn/m and.32 % displacement, so the ultimate capacity for the 2 m bolt will be 66 kn at 6.4 mm at displacement. 25% of bolts will exhibit lower stiffness, 25% lower ultimate capacity, but only 6.25% will have a lower ultimate capacity with more displacement Characterization of FRS Behaviour Figure 7.5 shows the load-displacement relationships for all FRS pull tests for which displacement was recorded. Displacements at.5 ton (4.45 kn) intervals are shown (although data was logged digitally with variable load resolution), and behaviour after ultimate capacity is reached is not shown. FA35 bolts are depicted in black, and FA39 in red FA35 FA Displacement (mm) Figure 7.5: All FA35 and FA39 pull test load-displacement relationships Both bolts show erratic displacement behaviour as load develops, suggesting that in many cases the bolt slips multiple times before reaching its ultimate capacity. To construct a representative distribution of behaviour, percentiles of displacement were calculated at.5 t (4.45 kn) intervals. As load increases, the dataset diminishes as the ultimate capacities of bolts are surpassed. To address this, dummy data entries were used for these bolts so that the number of bolts used to construct the percentiles remained constant. While the behavioural range is calculated using displacement, this is interpreted as stiffness, i.e. the 9 th percentile of displacement is equal to the 1 th percentile of stiffness. Figure 7.6 illustrates both the percentiles of stiffness (calculated from displacement at load intervals), as well as the percentiles of ultimate capacity simultaneously. As once the ultimate capacity of a particular test has been exceeded no displacement data at higher loads may be recorded, the 1 th percentile of stiffness must end at the 1 th percentile of ultimate capacity, as only 9% of the data remains from which to calculate stiffness (mathematically, at this point the dummy entry influences the value of the percentile, so the percentile

132 Loadc(kN) Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour 118 is no longer shown). This should not be taken to indicate that bolts with low secant stiffness have proportionally low ultimate capacities; Figure 7.2 shows that this does not appear to be the case. Figure 7.7 shows a smoothed version of the same figure displaying the initial anchorage capacity suggested by the manufacturer (courtesy of Supplier A) P 9 7 P P 1 P 5 P Displacement (mm) Figure 7.6: Distribution of displacement measured during pull testing of FA35 and FA39 with anchorage lengths of 1.52 m P 9 P 75 P 5 P 25 P 1 4 InitialcCapacitycRange (FromcSupplier) Displacementc(mm) Figure 7.7: Conceptual displacement distribution of FRSs with anchorage lengths of 1.52 m subject to a pull test Figure 7.6 shows that in situ FRS behaviour may be quite variable; it is possible that next to no load

133 Loadt(kN/m) Chapter 7. Characterization of Rock Bolt Behaviour 119 is required to initialize displacement, or that no displacement may be observed until 35 kn of load. Load then increases with growing displacement before plateauing at a maximum value (ultimate capacity). In Figure 7.7, behaviour is extrapolated to be perfectly plastic. In reality, after reaching the ultimate capacity, load is continually built and released as the bolt slips out of the hole incrementally. Note that the percentiles of ultimate capacity displayed in this figure are for the FA35/FA39 dataset, and not the entire FRS dataset. Figure 7.8 shows a generalized version of Figure 7.7. Load is expressed in load per metre of anchorage, and displacement in terms of anchorage length. This characterisation of behaviour was performed only bolts with an anchorage length of 1.52 m, and its applicability to other bolt lengths has not been verified. 7 6 P P 75 P 5 P 25 P Displacementt(9toftanchoragetlength) Figure 7.8: Conceptual displacement distribution of pull tests performed on FRSs Initial anchorage values of 27 to 54 kn are claimed by Suppliers A and B for both 35 and 39 mm configurations of FRS. As seen in Figure 7.7, pull testing indicates that these values are generally met or exceeded. However, both suppliers claim anchorage values of 54 kn to 89 kn for their respective 46 mm FRSs (courtesy of Suppliers A and B). Table 7.3 compares manufacturer recommendations with observed pull test results in terms of both absolute and length normalized load. As both suppliers provide their load recommendations in terms of absolute load, the typical 35 mm and 39 mm anchorage length of m is used to normalize load to length for all bolts. For 35 and 39 mm bolts, almost all pull tests (97.6%) have ultimate capacities exceeding the lower bound of the suppliers claim, and around 6% exceed the upper bound. Suppliers claim a higher range of ultimate capacities for their 46 mm bolts between 54 kn and 89 kn. It has been found in this thesis that there is no difference in performance across different FRS diameters. 77.7% of the 46 mm FRS tests exceeded the lower bound of the claimed range, but only 15.5% exceeded the upper bound, a much lower pass rate than the 59.2% observed for the 35 mm and 39 mm FRSs. When loads are adjusted to a common anchorage length, the discrepancy in supplier specifications becomes more apparent. 59.4% of 35/39 mm FRSs pass their upper bound of 35.4 kn/m, compared to only 1.9% of 46 mm FRSs passing their upper bound of 58.4 kn/m. It is difficult to see how suppliers determine their capacity

134 Chapter 7. Characterization of Rock Bolt Behaviour 12 specifications, as it has been observed that no difference exists between the length normalized ultimate capacities of different FRS diameters. Assuming that the claimed ranges of capacity are determined consistently (i.e. use of equivalent bolt length, installation procedures, etc.), it is difficult to explain this variation. Table 7.3: FRS performance compared to supplier specifications Bolt Claimed Capacity % Passing Claimed Capacity % Passing FA35 27 kn 98.8% 17.7 kn/m 98.8 % 54 kn 66.7% 35.4 kn/m 69.1% FA39 27 kn 1.% 17.7 kn/m 1.% 54 kn 53.3% 35.4 kn/m 53.3% FA46 54 kn 84.% 35.4 kn/m 64.% 89 kn 14.% 58.4 kn/m 4.% FB35 27 kn 94.6% 17.7 kn/m 93.5% 54 kn 67.4% 35.4 kn/m 65.2% FB39 27 kn 98.9% 17.7 kn/m 97.6% 54 kn 46.2% 35.4 kn/m 45.8% FB46 54 kn 78.% 35.4 kn/m 54.7% 89 kn 16.% 58.4 kn/m.9% 35 mm + 39 mm 27 kn 97.6% 17.7 kn/m 96.9% Total 54 kn 59.2% 35.4 kn/m 59.4% 46 mm Total 54 kn 77.7% 35.4 kn/m 57.7% 89 kn 15.5% 58.4 kn/m 1.9% It is recognized that this discussion touches only on recently installed bolts. The 46 mm bolt has more constituent steel, as it is larger in diameter. This means that the load at which it will physically fail is greater than the smaller diameters of FRS discussed. If an FRS is sheared an locked in place, a larger cross-sectional area will result in a higher load required to break the bolt. Similarly, the shear strengths of the bolts may not be equivalent. However, if these failure modes are not observed at an operation, the FRS bolts appear to have the same capacity regardless of diameter. 7.3 Rebar Rock Bolts Displacement data from 51 fully encapsulated resin-grouted Rebar A and Rebar B between 1.8 m and 2.4 m in length was collected. Various pre-loads, between and 3 tons (26.7 kn) were applied during testing of these rock bolts. Displacement data from these bolts was used to characterize the performance and behaviour of rebar. When working capacity is quantified, partially encapsulated bolts are incorporated into the analysis.

135 LoadA(kN) Chapter 7. Characterization of Rock Bolt Behaviour Characterization of Rebar Rock Bolt Performance In Chapter 5, it was found that rebar from both Suppliers A and B had similar working capacities and secant stiffness. Table 7.4 elaborates on this similarity, showing that distributions of both performance metrics for fully encapsulated bolts as quantified by percentiles are comparable. As such, the two datasets are combined in Figure 7.9, depicting the combination of load and displacement at which the rebar reached their working capacities. Most of the rebar was pull tested after pre-loading the bolt to 2 tons (17.8 kn). The points in Figure 7.9 are thus depicted with a simulated 17.8 kn pre-load. Table 7.4: Percentiles of performance metrics for rebar rock bolts Metric Supplier n P 1 P 25 P 5 P 75 P 9 A kn kn kn kn kn Working capacity B kn kn kn kn kn Total kn kn kn kn kn A kn/mm 15.3 kn/mm 28.9 kn/mm 43.8 kn/mm 46.9 kn/mm Secant stiffness B kn/mm 15.3 kn/mm 27.9 kn/mm 35.2 kn/mm 46.6 kn/mm Total kn/mm 15.8 kn/mm 28.1 kn/mm 39.5 kn/mm 45.9 kn/mm P 9 P 75 P 5 P 25 P 9 P 75 P 5 P 25 P 1 P RebarAA RebarAB DisplacementA(mm) Figure 7.9: Distributions of secant stiffness and working capacity for a pull test with a pre-load of 17.8 kn on a rebar rock bolt Characterisation of Rebar Rock Bolt Behaviour Figure 7.1 shows all pull tests performed. Tests performed with a pre-load of are shown in red, a pre-load of 17.8 kn in grey, and 26.7 kn in black.

136 Load (kn) Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour t (26.7 kn) Pre-load 2 t (17.8 kn) Pre-load Pre-load Displacement (mm) Figure 7.1: All rebar pull test load-displacement relationships It can be seen in Figure 7.1 that large displacements at low loads may occur if the bolt is not preloaded. Figure 7.11 shows the load displacement data neglecting the first 17.8 kn of all pull tests. Figure 7.12 incorporates displacement data below 17.8 kn and correcting the percentiles of the entire data set to match those of the tests performed without pre-load at 17.8 kn. As pull tests are often performed only until the rebar yields, if not before, data limitations result in the 75 th and 9 th percentiles not clearly showing yield behaviour P 25 P 1 P 5 P 75 P Displacement (mm) Figure 7.11: Distribution of displacement measured during pull testing of 2 mm rebar 1.8 m to 2.4 m in length with a pre-load of 17.8 kn A comparison of Figures 7.11 and 7.12 illustrates the objective of using a pre-load during pull testing;

137 LoadW(kN) Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour P 1 P P 5 P 75 P Displacement (mm) Figure 7.12: Distribution of displacement measured during pull testing of 2 mm rebar 1.8 m to 2.4 m in length without a pre-load large displacements occur at low loads as the rock mass and testing apparatus tighten up, before a stiffer linear response to load is developed. As the stiffness and working capacity of rebar subject to a pull test are independent of one another, they are separated in Figure The behaviour shown in Figure 7.12 is extrapolated to P 9 of working capacity. Median working capacity is shown in red, and other percentiles of working capacity are listed. Note that partially encapsulated rebar were incorporated into the dataset used to calculate percentiles of working capacity P 1 P 25 P 5 P 75 P PercentilesWofWWorkingWCapacity P 9 131WkN P WkN P 5 124WkN P WkN P 1 113WkN DisplacementW(mm) Figure 7.13: Conceptual distribution of displacement during pull testing of 2 mm rebar without a pre-load

138 Chapter 7. Characterization of Rock Bolt Behaviour 124 Figure 7.13 shows that less than 2 mm of displacement may be observed when pull testing a 2 mm rebar before working capacity is reached and the rebar begins to yield. About 5 mm of displacement represents median behaviour if the bolt is not pre-loaded, and in some conditions over 14 mm of displacement may be measured. Table 7.5 compares the rebar strength claimed by the supplier to the observed working capacities. Table 7.5: Rebar performance compared to supplier specifications Metric Specification % Passing Minimum thread yield strength 89 kn (Supplier A), 86 kn (Supplier B) 1% Minimum thread tensile strength 134 kn (Supplier A), 116 kn (Supplier B) 1% All rebar reached their working capacities above the supplier-specified thread yield strength. While not all rebar were tested until the 134 kn minimum thread tensile strength specified by Supplier A, thread failure was not reported in any pull test. 7.4 Modified Cone Bolt The behaviour of the MCB33 was described by four metrics combining to be represented by one (secant stiffness). Secant stiffness and steel yield strength are the two primary metrics used to describe bolt performance, although all metrics contribute to the overall behaviour of the MCB Characterisation of Modified Cone Bolt Performance Figure 7.14 shows the combinations of load and displacement at which the bolts began to yield (quantified by secant stiffness), assuming a pre-load of 2 tons (17.8 kn). This is a limited representation of bolt behaviour that is elaborated on in Table 7.6, which shows percentiles of each metric. All cone bolts tested were the same length, so no length normalization was performed. Table 7.6: Distributions of MCB33 performance metrics Metric n P 1 P 25 P 5 P 75 P 9 Initial Stiffness kn/mm 7.5 kn/mm 9.7 kn/mm 11.4 kn/mm 15.2 kn/mm Plough Load kn 35.4 kn 53.4 kn 68.4 kn 91.6 kn Plough Stiffness kn/mm 1.4 kn/mm 2.1 kn/mm 2.8 kn/mm 3.7 kn/mm Yield Strength kn kn kn 13.3 kn kn Secant Stiffness kn/mm 1.9 kn/mm 2.8 kn/mm 4.1 kn/mm 6.3 kn/mm Figure 7.14 shows that cone bolt performance is more variable than that of rebar, in terms of both yield strength and displacement. While the MCB33 yield strengths observed are similar to those of the rebar, the displacement at which yield occurs is much larger. When both elements are installed in a support system, this would result in the stiffer rebar bearing more load than the less stiff cone bolt. As such, the yield strengths presented in Table 7.6 should not be used in a design methodology based solely on the load demand on and capacity of a support system, but one that also incorporates displacement into the calculation.

139 Loadm(kN) Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour Displacement (mm) P 9 P 75 P 5 P 9 P 75 P 5 P 25 P 25 P 1 P Displacement (mm) Figure 7.14: Distributions of secant stiffness and working capacity for a pull test with a pre-load of 17.8 kn on an MCB Characterisation of Modified Cone Bolt Behaviour In Figure 7.15, 58 pull tests on the MCB33 are shown, sorted by pre-load. The behaviour comprised of elastic deformation, plough and plastic deformation of the bolt as described in Section is seen for most of the pull tests. As the objective of pull testing cone bolts is to demonstrate their large displacement capacities as opposed to load capacity, tests may be stopped when plough has been demonstrated but not yield, or alternatively after significant plastic deformation has been incurred mtm(35.6mkN)mPre-load 3mtm(26.7mkN)mPre-load 2mtm(17.8mkN)mPre-load Displacementm(mm) Figure 7.15: 56 MCB33 pull tests collected from Vale s Sudbury operations

140 Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour 126 Percentiles of displacement are presented in Figure As many pull tests are stopped before the bolt has yielded, data becomes increasingly limited at higher loads. Figure 7.16 was created by decreasing the size of the dataset used to calculate percentiles (from n = 45 to n = 32) at a load where a large number of tests are stopped (12 tons; 17 kn) by eliminating dummy displacements from the percentile calculations at 17 kn. Above 17 kn, dummy displacements were reintroduced to maintain the n = 32 dataset as it decreased in size at higher values of load P 1 P 25 P 5 12 P 75 P Displacement (mm) Figure 7.16: Distribution of displacement measured during pull testing of a 2.4 m MCB33 with a pre-load of 17.8 kn Figure 7.16 clearly shows the two distinct behaviours of the cone bolt described in Section 5.4: the 1 th and 25 th percentiles are comprised of the bolts that do not develop a linear plough response before yielding, while a linear plough response is seen in the larger percentiles. The development of this behaviour is the principal controlling factor over the displacement that may be observed before the bolt yields. Figure 7.17 shows a conceptual version of Using the initial stiffness of the MCB33s in the database, percentiles are extrapolated to load as if no pre-load were recorded. While it is acknowledged that larger displacements may be observed at low loads as is the case for the rebar, these displacements are still relatively small compared to those that occur as the bolt ploughs through the resin. In Figure 7.17, it can be seen that the median displacement observed before an MCB33 begins to yield is 35 mm, although can be as much as 1 mm. Extreme cases of low plough fall below P 1, although a decreasing stiffness is observed. In Table 7.7, Mansour s specifications regarding yield strength are compared with the observations made in the database. Table 7.7: MCB33 performance compared to supplier specifications Metric Specification % Passing Minimum thread yield strength 98.5 kn 1% Typical thread yield strength kn 81%

141 Loadf(kN) Chapter 7. Characterization of Rock Bolt Behaviour P 1 P 25 P 5 P 75 P PercentilesfoffYieldfLoad P 9 137fkN P 75 13fkN P 5 125fkN P fkN P 1 17fkN Displacementf(mm) Figure 7.17: Conceptual distribution of displacements for a pull test of a 2.4 m MCB33 without a pre-load The minimum thread yield strength (98.5 kn) appears to accurately reflect a minimum, as all of the 47 bolts with a yield strength recorded begin to yield above these values. Additionally, 81% begin yielding above the typical thread yield strength. As such, pull test findings agree with the specifications claimed by the supplier. However, it must be reiterated that because of the large displacement the MCB33 undergoes even in static loading conditions, design methodologies that incorporate the static reinforcement offered by the Modified Cone Bolt must consider displacement as well as load. 7.5 D-Bolt Because of the high strength of the D-Bolt, testing personnel usually used pull testing as a verification tool and did not yield the bolt in every testing campaign due to safety concerns associated with bolt failure during a pull test. As a result, very few working capacities were recorded in comparison to values of secant stiffness. As such, stiffness (and thus displacement) are primarily evaluated Characterisation of D-Bolt Performance Figures 7.18 and 7.19 show percentiles of stiffness plotted with minimum and typical yield loads of the 2 mm and 22 mm D-Bolts according to Normet (214). These appear to agree fairly well with the limited working capacity data that does exist (displayed as black points in both figures). Table 7.8 shows the percentiles of secant stiffness for both D-Bolt configurations. Table 7.8: Secant stiffness distributions of 2 mm and 22 mm D-Bolts Bolt n P 1 P 25 P 5 P 75 P 9 2 mm D-Bolt kn/mm 8.5 kn/mm 23.3 kn/mm 35.8 kn/mm 59.5 kn/mm 22 mm D-Bolt kn/mm 5.3 kn/mm 11.1 kn/mm 24.3 kn/mm 26.1 kn/mm

142 Loady(kN) Loady(kN) Chapter 7. Characterization of Rock Bolt Behaviour P 1 P 25 P 5 P 75 P 9 TypicalyYieldyLoad MinimumyYieldyLoad Displacementy(mm) Figure 7.18: Distribution of secant stiffness for a pull test with a pre-load of 17.8 kn on a 2 mm D-Bolt 25 2 P 1 P 25 P 5 P 75 TypicalyYieldyLoad P 9 MinimumyYieldyLoad Displacementy(mm) Figure 7.19: Distribution of secant stiffness for a pull test with a pre-load of 17.8 kn on a 22 mm D-Bolt Figures 7.18 and 7.19 show that the pull tests performed on both D-Bolts resulted in variable displacements and behaviours, with the 22 mm bolt acting in a less stiff manner than a 2 mm bolt. As discussed in Section 4.3.4, much of the 22 mm D-Bolt testing was performed in very fractured ground conditions, potentially influencing the degree of displacement observed. Despite this, both D-Bolt configurations perform stiffly compared to the Modified Cone Bolt. As the stiffness is so variable, site specific pull testing is recommended in order to fully understand displacement development with load, and how

143 Load (kn) Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour 129 to implement the D-Bolt into a ground support standard Characterisation of D-Bolt Behaviour Figures 7.2 and 7.21 show the pull tests collected for the 2 mm and 22 mm D-Bolts. It can be seen that very few of the bolts yield during testing Displacement (mm) Figure 7.2: 2 mm D-Bolt pull tests collected from Vale s Sudbury operations Displacement (mm) Figure 7.21: 22 mm D-Bolt pull tests collected from Vale s Sudbury operations Figures 7.2 and 7.21 serve to highlight the wide range of behaviour observed for both diameters of D-Bolt when subject to a pull test. However, a cluster of similarly behaving 22 mm D-Bolts with

144 Load (kn) Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour 13 relatively high stiffness is observed in Figure All of these bolts were tested in one campaign, demonstrating that the D-Bolt may exhibit consistent behaviour within some testing campaigns, but not in others. Figures 7.22 and 7.23 show the distributions of behaviour for either diameter of D-Bolt P 1 P 25 P P 75 P Displacement (mm) Figure 7.22: Load displacement behaviour of a 2 mm D-Bolt with a pre-load of 17.8 kn 25 P 1 P 25 P 5 2 P P Displacement (mm) Figure 7.23: Load displacement behaviour of a 22 mm D-Bolt with a pre-load of 17.8 kn The D-Bolt has the potential to exhibit high displacement at low loads as seen for P 75 and P 9 of displacement for the 22 mm D-Bolt. This is similar to the behaviour observed for rebar in Figure 7.12, where large initial displacements may be attributed to compression of the rock mass and testing apparatus. The effect in Figure 7.23 continues to higher loads than the equivalent effect for rebar, but

145 Loady(kN) Loady(kN) Chapter 7. Characterization of Rock Bolt Behaviour 131 this is thought to be the result of a large proportion of pull tests on the D-Bolts begin performed in poor ground. Figures 7.24 and 7.25 show conceptual versions of D-Bolt behaviour with manufacturer specifications (Normet, 214). Behaviours are extrapolated to kn P 1 P 25 P 5 P 75 P 9 TypicalyYieldyLoad MinimumyYieldyLoad Displacementy(mm) Figure 7.24: Conceptual distribution of displacements for a pull test of a 2 mm D-Bolt without a pre-load P 1 P 25 P 5 P 75 P 9 TypicalyYieldyLoad MinimumyYieldyLoad Displacementy(mm) Figure 7.25: Conceptual distribution of displacements for a pull test of a 22 mm D-Bolt without a pre-load While too little working capacity data was collected to reliably verify that the D-Bolt performs at or above the supplier s specifications in a pull test, it should be acknowledged that when working

146 Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour 132 capacity/yield was captured during the test, it always occurred above the minimum yield load specified. This includes 3 2 mm D-Bolts yielding above 14 kn, and 3 22 mm D-Bolt yielding above 17 kn. 7.6 Expandable bolts As discussed in Section 4.3.5, the expandable bolt generally begins to yield before it develops a consistent slipping response, i.e. working capacity is defined by yield and not slip. As such, load is not length normalized. Because of the highly variable manner in which expandable bolts appear to react to load in a pull test and the unclear relationship between stiffness and length, stiffness is also expressed in absolute terms. In Section 5.6 it was found that very little difference existed between the Manganese and Premium lines of Swellex bolt (Mn and Pm) in terms of stiffness and insufficient data was available to fully investigate the working capacity (while Atlas Copco claim equal minimum yield loads between the Pm12 and Mn12, the Pm24 has a yield load 2 kn greater than that of the Mn24; Atlas Copco, 212). Additionally, the behaviour of the two types of bolt mainly differ after yield; the Mn line may exhibit significantly more plastic deformation than an equivalent Pm bolt before failing (Scolari, 25). As behaviour up until yield (i.e. working capacity) is examined, Mn12 and Pm12 Swellex configurations are grouped together, as are the Mn24 and Pm24 configurations Characterisation of Expandable Bolt Performance Figure 7.26 shows the combinations of load and displacement that characterise the capacities of the Pm12 and Mn12 bolts, summarized with secant stiffness data for the Pm24 and Mn24 in Table 7.9. Note that the 25 th and 5 th percentiles of Pm12/Mn12 working capacity are both 89 kn. As discussed in Section 4.3.5, pull tests on Pm24 and Mn24 bolts were generally not performed until the bolt yielded. Only data from bolts that were fully embedded in ore or rock (i.e. not back fill) was used for stiffness calculations P 75 P 5 P 25 P 9 P 75 P 25 /P 5 P Displacement (mm) Figure 7.26: Distributions of secant stiffness and working capacity for a pull test on Swellex Pm12 and Mn12 without a pre-load

147 Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour 133 Table 7.9: Swellex secant stiffness percentiles Metric Bolt n P 1 P 25 P 5 P 75 P 9 Working capacity Pm12 & Mn kn 89 kn 89 kn 93.4 kn 99.6 kn Secant Stiffness Pm12 & Mn12 8 N/A 5. kn/mm 5.8 kn/mm 1.1 kn/mm N/A Pm24 & Mn kn/mm 7.8 kn/mm 18.1 kn/mm 22.1 kn/mm 28.3 kn/mm Little available data only permitted a limited analysis of Swellex performance compared to the other bolts discussed. Figure 7.26 shows that the Pm12 and Mn12 bolts reach their working capacities consistently at loads of about 9 kn. In Table 7.9 it can be seen that the Pm24 and Mn24 have the potential to react in a very stiff, as well as a very soft manner. This variable behaviour is not as apparent with the Pm12 and Mn12 bolts, although less data is available Characterisation of Expandable Bolt Behaviour Figures 7.27 and 7.28 show pull test data collected for the four configurations of Swellex Displacement (mm) Figure 7.27: Swellex Pm12 and Mn12 pull tests collected from Vale s Sudbury operations

148 Load (kn) Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour Displacement (mm) Figure 7.28: Swellex Pm24 and Mn24 pull tests collected from Vale s Sudbury operations In Figure 7.27 it can be seen that pull tests performed on Pm12/Mn12 may exhibit several behaviours. In some cases, very little displacement is initially observed, although stiffness decays with load. In other cases, a linear response to load is observed. The Pm24 and Mn24 bolts act in a more consistent manner, usually responding linearly to load, although two pull tests exhibit low initial stiffness. The distributions of these behaviours are constructed in Figures 7.29 and 7.3. Insufficient data was available to calculate 1 th and 9 th percentiles of displacement for the Pm12 and Mn12 bolts. All pull tests shown for these configurations had no pre-load. A pre-load of 4 tons (35.6 kn) is depicted for the Pm24 and Mn P 25 P 5 P Displacement (mm) Figure 7.29: Load displacement behaviour of Pm12 and Mn12 bolts

149 Load (kn) Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour P 1 P 25 P 5 P 75 P Displacement (mm) Figure 7.3: Load-displacement behaviour of Pm24 and Mn24 bolts Figures 7.29 and 7.3 show that Swellex expandable bolts appear to react linearly to load, although may bear low loads with little to no displacement. The Pm24 and Mn24 bolts exhibit a consistently stiff response to load, however there does appear to exist a relatively small possibility of a much softer response (note that this may be in part attributable to the reaction of the rock mass and testing apparatus). As a result of a lack of displacement data at high loads, neither figure clearly shows yielding behaviour. Figures 7.31 and 7.32 show smoothed relationships and depict the supplier s specifications regarding yield load P 25 P 5 P 75 Minimum Yield Load (Pm12, Mn12) Displacement (mm) Figure 7.31: Conceptual load displacement behaviour of Pm12 and Mn12 bolts subject to a pull test

150 Load (kn) Chapter 7. Characterization of Rock Bolt Behaviour P 1 P 25 P 5 P 75 P 9 2 Minimum Yield Load (Pm24) Minimum Yield Load (Mn24) Displacement (mm) Figure 7.32: Conceptual load displacement behaviour of Pm24 and Mn24 bolts subject to a pull test Only sufficient working capacity data was collected for the Pm12 and Mn12 to compare with supplier specifications, shown in Table 7.1. As displacement data was collected in 1 ton (8.9 kn) load increments, one of the data bins is the 89 kn kn. This means any bolt that yields between 89 kn and 97.9 kn, it is said to have yielded at 89 kn using procedures enacted throughout this thesis. This is problematic, as a large portion of the data (47%) falls in this range, and the supplier specified minimum yield load for the Pm12 and Mn12 is 9 kn. As such, bolts with a recorded working capacity of 89 kn are assumed to have passed the 9 kn threshold specified by the supplier. Table 7.1: Swellex Pm12 and Mn12 performance compared to supplier specifications Metric Specification % Passing Minimum yield strength (Pm12, Mn12) 9 kn 88% Two of the Swellex pull tested did not comply with manufacturer specifications (both of them Pm12). It is possible this is due to a difference in the definition of working capacity or its method of determination used in this thesis, and the definition of yield strength used by Atlas Copco. Higher load resolution is desirable to determine with more certainty whether bolts in the 89 kn to 97.9 kn range yielded above or below 9 kn. Additionally, the supplier s specifications for the Pm24 and Mn24 should be verified by pulling these bolts to loads above 2 kn and 18 kn respectively. 7.7 Summary An important part of designing a support system is the selection of the appropriate reinforcement elements required to stabilise an excavation (Thompson et al, 212). Figure 7.33 shows the conceptual behaviours of all bolts discussed on similar scales. Some bolts (MCB33, both diameters of D-Bolt, and Swellex Pm24/Mn24) have their behaviours extrapolated to load as a result of a lack of data.

151 Load (kn) Load (kn) Load (kn) Load (kn) Load (kn) Load (kn) Loady(kN) Chapter 7. Characterization of Rock Bolt Behaviour MedianyWorkingyCapacity Displacement (mm) (a) FA35 and FA Displacementy(mm) (b) Rebar Displacement (mm) (c) MCB Typical Yield Load Displacement (mm) 25 (d) 2 mm D-Bolt Displacement (mm) 25 (e) 22 mm D-Bolt Minimum Yield Load (Pm24) Minimum Yield Load (Mn24) 1 5 Minimum Yield Load Displacement (mm) (f) Swellex Pm12 and Mn Displacement (mm) (g) Swellex Pm24 and Mn24 Figure 7.33: Conceptual load displacement behaviour extrapolated to load for various rock bolts subject to a pull test

152 Loadg8kN( Chapter 7. Characterization of Rock Bolt Behaviour 138 Figure 7.34 shows the median behaviours of all bolts discussed in this chapter extrapolated to a load of kn. 25, 2,, 22gmmgDFBolt SwellexgPm24 SwellexgMn24 MediangWorkinggCapacity TypicalgYieldgLoad MinimumgYieldgLoad 95, 2,gmmgDFBolt 2,gmmgRebar MCB33 9,, SwellexgPm923Mn92 FA35sgFA39 5,,, 2, 4, 6, 8, 9,, Displacementg8mm( Figure 7.34: Median load displacement behaviour for all bolts with no pre-load

153 Chapter 7. Characterization of Rock Bolt Behaviour 139 Figure 7.34 shows that the 22 mm D-Bolt appears to have the greatest load capacity, although the minimum yield loads shown for the Swellex Pm24 and Mn24 are minima specified by Atlas Copco. The stiffest reinforcement element is the 2 mm rebar, likely due to limited load propagation down the bolt. The least stiff bolts (in terms of secant stiffness) appear to be the MCB33 and the Swellex Pm12 and Mn12. While the MCB33 is a yielding bolt, and the ploughing mechanism is the cause of the large displacement, it is unclear why the Pm/Mn12 acted in so soft a manner as continuously frictionally coupled bolts. The bolts with the lowest working capacity are the two FRS A diameters shown, approximately representative of all FRS configurations tested. If the working capacity of the MCB33 is defined by the onset of plough, it has quite a similar median capacity to that of the FRS A bolts. However, the load at which plough occurs is much more variable than the ultimate capacities achieved by FRSs as a whole. Two bolts in Figure 7.34 were examined by Stillborg (1993), specifically resin-grouted 2 mm rebar and a 39 mm nominal diameter FRS. The rebar tested by Stillborg (1993) behaves in a slightly stiffer manner than the median depicted in Figure This is to be expected comparing laboratory with field testing, and the difference is only 3-4 mm. As Stillborg (1993) tested bolts across a simulated joint, the yield response measured is for the bolt tendon, while the working capacity shown in Figure 7.34 is the result of the threads yielding; plastic deformation starts around 15.3 tonnes (15 kn) for the rebar tested by Stillborg (1993), while median rebar working capacity is kn. Stillborg (1993) depicts the SS39 as undergoing no displacement until its capacity is reached, at which stage it slips out of the hole at a constant load. Pull tests that constituted the database showed that bolts usually slip multiple times with growing displacement before reaching ultimate capacity; the median behaviour shows a displacement of 11 mm before ultimate capacity is reached. The SS39 tested by Stillborg (1993) was 3 m in length, and had a capacity of approximately 5 tonnes (49 kn). However, as Stillborg (1993) conducted testing across a simulated joint, in reality two 1.5 m lengths of Split Set were being tested. This means that a length-normalized capacity of 32.6 kn/m was recorded. In comparison, the median ultimate capacity for 1.68 m FA35/39s was 62 kn (and the overall median was 38.7 kn/m), however Tomory et al. (1998) recorded an average Split Set capacity of 31.9 kn/m, likely as the result of jackleg installations. While the Swellex tested by Stillborg (1993) is no longer marketed, its overall behaviour may still be compared to modern Swellex bolts. In Stillborg s (1993) test, the EXL Swellex dowel undergoes a very stiff response before slipping in a manner comparable to an FRS at 11 tonnes (18 kn). In situ pull tests on modern Swellex bolts show quite a different behaviour. Although a softer linear response is observed, the modern Swellex is not intended to slip when installed in hard rock, and was generally shown to yield in the database, excluding partial embedment tests and paste/sand fill installations. When using the figures in this section, it must be kept in mind that these results represent pull tests on bolts, and may not necessarily reflect performance when subject to in situ loading mechanisms. In Chapter 8, the limitations of this thesis, including the pull tests as a method of assessing reinforcement element performance, are discussed. Recommendations are also provided to improve the pull testing procedures observed during this thesis.

154 Chapter 8 Conclusions In the final chapter, the contributions made by this thesis are outlined. recommendations made and the path forward laid out. Limitations are discussed, 8.1 Contributions In order to conduct this thesis, a large database of pull tests performed on various rock bolts was assembled. This database includes 985 individual pull tests performed between 23 and 215 on friction rock stabilizers, rebar rock bolts, Modified Cone Bolts, D-Bolts and expandable bolts, amongst others. The data was collected from 6 Vale operations in and around the Sudbury Basin. Although databases of pull tests have been assembled previously, this database incorporates a wider range of bolt types. In addition, collecting pull test data from mines within relatively close proximity to one another under one company provides consistency in procedures and conditions for bolt installation and testing, as well as consistency in the ground control products used on and delivered to the 6 mine sites. A review was performed on methods of testing rock bolts. This included the ASTM standard (ASTM D ) and ISRM suggested methods for pull testing. These were compared with the methods and apparatuses used for pull testing in practice, and shortcomings of the standard, the suggested method, and the practised methods of pull testing outlined. The measurements taken from a pull test, metrics that may be calculated and the significance of the measurements and metrics was also discussed. An analysis of the data collected followed. Theoretical bolt behaviour of rock bolts subject to loading was compared to behaviour observed during the in situ pull tests. Various metrics of bolt performance describing load and displacement behaviour, primarily ultimate/working capacity and several measures of stiffness, were quantified and their distributions characterised for different types of bolt. Subsequently, the influence of factors associated with bolt installation, the rock mass and the bolt itself on rock bolt performance were investigated. The findings of this analysis built on and was compared to previous work when possible. Using the collected load-displacement behaviour of friction rock stabilizers, rebar, the MCB33, D- Bolts and Swellex, a distribution of behaviour was built for each bolt. Although the characterisation of rock bolt behaviour has been previously performed in a laboratory setting, the analysis contained herein uses a larger dataset than these analyses. Additionally, as the pull tests are performed in situ, they better represent installation conditions found at a mine site and as a result are more applicable to the 14

155 Chapter 8. Conclusions 141 design of underground excavations in hard rock mines. 8.2 Load Capacities of Reinforcement Elements The analyses contained in this thesis have allowed for the development of summary statistics for the capacities of different rock bolts from pull tests. Ultimate capacity for the FRS is quantified in load per anchorage length, assumed to be.15 m less than the length of the element. The average ultimate capacities of all FRS diameters were found to be approximately equivalent, ranging between 37.8 kn/m for the FB39 and 4. kn/m for both the FA35 and FB35. A dataset composed of all FRS data points gives a normal distribution of working capacity, with an average of 38.9 kn/m and standard deviation of 11.7 kn/m. This proved to be the reinforcement element investigated with the largest coefficient of variation describing load capacity. Although the working and ultimate capacities of reinforcement elements that physically yield or fracture as opposed to slip are dependent primarily on the mechanical properties of their constituent steel, the capacity of the FRS is more strongly dictated by how the bolt interacts with the rock mass in which it is installed. The two suppliers of FRS bolts provide a range of initial capacities for the bolts. The capacity ranges for both the 35 mm and 39 mm nominal diameters are between 27 kn and 54 kn (note these loads are not normalized to anchorage length). 97.6% of the FRS bolts with nominal diameters of 35 and 39 mm had capacities greater than 27 kn, and 59.2% had capacities greater than 54 kn. In comparison, both suppliers claim an initial capacity range of 54 kn to 89 kn for their 46 mm FRS. 77.7% of 46 mm bolts had capacities greater than 54 kn (note that the 46 mm bolts were generally.3 m longer than the 35 and 39 mm bolts tested), and only 15.5% passed the upper bound of 89 kn. As a result, there appears to be a difference in the way the ranges of capacities are calculated for different FRS diameters for both of the suppliers. Rebar rock bolts from two suppliers were examined. A small discrepancy in working capacity was observed between the two, however this may be the result of differing testing apparatuses and procedures. Rebar from Supplier A had an average working capacity of kn with a standard deviation of 7.5 kn. Rebar B had an average working capacity of 12.9 kn with a standard deviation of 7.4 kn. While the pull tests conducted by Supplier A were generally performed using digital data collection, Supplier B s tests often recorded data manually at load increments of one ton. As such, working capacities are rounded down to the nearest ton, while in the case of tests on Rebar A the working capacity will be rounded down to the nearest measurement of load. Despite this, all rebar exceed the minimum thread yield and tensile strengths specified by their respective suppliers. Two metrics of load capacity were defined and quantified for the MCB33: plough point and yield strength. The plough point appears to be normally distributed when only examining bolts tested immediately after installation, with an average value of 53.2 kn and standard deviation of 22.7 kn. The average MCB33 yield strength was kn, with a standard deviation of 16.3 kn. All of the MCB33s tested had working capacities greater than the minimum thread yield strength specified by Mansour, and 81% passed the typical thread yield strength. A review of the performance of these and other rock bolts (including the D-Bolt and Swellex) has allowed the development of charts depicting the behaviour of the bolts in a pull test. It was recognised that different metrics of performance may only be applicable to a certain set of bolt types. This is an important observation given that the ISRM suggested methods were developed prior to the introduction

156 Chapter 8. Conclusions 142 of many rock bolt types, and although the ASTM pull testing standard (D ) was developed more recently, it does not distinguish between types of rock bolt beyond grouted versus mechanically anchored bolts. 8.3 Limitations The limitations of this thesis can be broadly divided into three classes; the limitations of a pull test as a method of assessing performance, those associated with the pull test as implemented in practice, and the limitations of the data present in the database. The primary conceptual limitation of a pull test is that it is only representative of a very specific loading scenario as bolts are pulled axially at the head/thread. Loading in an underground environment may be much more complex; reinforcement elements may have multiple axial, shear or rotational loads applied at any location on their lengths, and these forces may evolve with time. Additionally, loading is only applied in a quasi-static manner during a pull test. The response to dynamic load is not measured, although may be of interest for yielding reinforcement elements. The most prevalent difficulty encountered when evaluating the performance of the rock bolts tested in the database was the presence of displacements attributed to the response of the excavation surface as opposed to that of the bolt alone. While the measurement of this response may be applicable to certain scenarios, it introduced not only variability, but also bias into the data, as certain elements (such as the D-Bolt) tended to be tested in poorer quality rock masses than others. ASTM D provides a methodology that limits the measurement of the rock mass response by using an apparatus that measures displacement relative to a stationary datum, however it was not enacted by pull testing personnel. Additionally, methods of data recording varied in the database between manual and digital displacement logging. Manual logging is performed in loading increments which may result in poor data resolution, while digital logging was often performed at a resolution that does not reflect the capabilities of this method. Shortcomings existed with the way in which pull tests were reported. Many parameters (for example rock mass quality or borehole diameter) were inconsistently or seldom noted. Pull tests were almost always done on rock bolts that had been installed the same day as the pull test, or if not, at some unnoted previous date. A significant shortcoming of the database is the lack of data describing bolt failure during a pull test. It is acknowledged that this was to preserve the safety of the personnel conducting the test, but it is an advantage of laboratory testing that bolts may be failed in a controlled and safe manner. 8.4 Recommendations Attempts have previously been made to introduce a standardized pull test data sheet, in which various parameters are to be recorded. For example, Soni (2) presents a sheet on which various testing parameters may be recorded. These include rock strength parameters, rock mass classification(s), the type and length of bolt, the diameter of the drillbit and drillhole, inflation pressure, inflation or drive time, installation date, residence time, and grout type, length and collar depth. In the pull test reports collected, these parameters were noted with varying frequency, although all are pieces of information that may affect performance. In addition to these, there are further recommendations to be made with

157 Chapter 8. Conclusions 143 regards to information recorded and pull test practices. In addition to those outlined by Soni (2), the following aspects of a pull test should be clearly reported. The objective and type of pull test should be stated. For example: partial embedment/encapsulation, bolt slip, bolt yield, achieve a load of x kn. If a test is ended prematurely, it should be explained why. Record the bolt name including configuration and presence or lack of possible modifications, for example a plastic coating. Bolt length (as well as anchorage length) and diameter should be clear. Report equipment used for the pull test. This includes the installation equipment, the loading system, displacement measurement system and the logging equipment. If an electric pump is used to apply load, record loading rate. If an electronic logger is used, report logging frequency. It should be noted whether the displacement measurement system is mounted on the pull test apparatus itself, or if it is stationary relative to the excavation. It should be explicitly stated what units are used to record load and displacement. recorded in tons, then short, long, or metric should be specified. If load is If the bolt is pre-loaded before recording displacement, that pre-load should be stated. If applicable, a bolt s spin time should be recorded. For these scenarios, temperature should also be measured. Appendix A presents pull test informations sheets from ASTM and ISRM, as well as modified data sheets, including one for the overall campaign and one to log each pull test individually. In terms of pull test procedure, modifications could be made without significantly affecting the amount of time required to perform them. They are as follows. If recording data manually, use a maximum load increment of 5 kn or.5 tons once load has been built to within 2 kn or 2 tons of the expected working capacity to obtain a more precise pull test result. If recording data electronically, the additional load resolution possible with this method must be exercised. If load is not built despite an observed increase in displacement, take two or more recordings of displacement to make it clear slip or perfectly plastic behaviour was observed. If a reinforcement element such as an FRS slips, the test should continue until a pre determined displacement (e.g. 15 mm) has occurred. This ensures that maximum recorded loads are at least comparable between tests, and the criterion used to dictate the end of a test is consistent. When using a method of measuring displacement that incorporates the rock mass response, unloading of the reinforcement element should be recorded to assess elastic deformation, and thus estimate how much of the displacement response may be attributed to the rock mass reaction. Displacement should be recorded at the highest load achieved. Pump pressure should then be released until a load equal to or greater than the pre-load is reached (if no pre-load is used, at

158 Chapter 8. Conclusions 144 least 15 kn is recommended), and displacement should be recorded again. This is to minimize the effect of rock mass and pull test apparatus relaxation. The measurement of the stiffness of the unloading response gives an idea of how load attenuates down the bolt length from the point of load application, and may be used as a relative measure of performance. In some cases, the analysis presented was hampered by a lack of data for specific types of bolt or parameters recorded. These limitations may be overcome by expanding the database. One of the advantages of the database assembled for this thesis over previously constructed pull test databases is that all pull tests were performed in and around the Sudbury Basin, reducing variability in installation conditions and procedures. If possible, this aspect of the database should be preserved, including data only collected from hard rock mines, and should be tested in a similar or superior manner to the pull tests presented herein. 8.5 Implications and Path Forward The database assembled presents a large number of pull tests performed on different reinforcement elements. Having reviewed and compared the standard and practical execution of testing methods, an analysis of the database was performed, quantifying various metrics of performance for most types of reinforcement element present in the database. Bolt performance across different conditions was examined, and a thorough review of the load displacement response of rock bolt installed in situ to a pull test was presented with comparisons to measures of performance dictated by the corresponding supplier. Although there are fundamental limitations of an in situ pull test, as well as weaknesses resulting from its practical execution, great value may be obtained at relatively little cost and disruption to mine site activities. The implications of these limitations are observable throughout this thesis, and ways in which future pull testing may be improved are discussed. Future work should entail expanding the database to build on and/or verify the analyses presented herein. If possible, more precise methods of pull testing should be enacted in order to reduce the subjectivity of results obtained from a pull test. The results of this thesis may be used in the process of designing a ground support system. The working capacity of a reinforcement element is a critical aspect of support system strength that should be known with a high degree of confidence. The displacement response of an element to load is a fundamental aspect of some design methodologies that must also be quantified. Understanding how different conditions may affect the performance of individual reinforcement elements is crucial information in a mine where various ground conditions or methods of bolt installation are used, or for an organization that operates several different mines. The ability to objectively compare reinforcement element performance, both to specifications provided by the bolt suppliers and to other types of element, allows for the selection of the correct tool for a design problem. With higher confidence in the performance of reinforcement elements, ground support systems that are not only more cost effective, but that are also safer, may be designed and implemented.

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165 Appendices 151

166 Appendix A Pull Testing Forms A.1 ASTM D Sample Form Figure A.1: Rock bolt pull test sample form (ASTM D4435, 213) 152