Analysis and Interpretation of In Situ Rock Bolt Pull Tests in ...

Analysis and Interpretation of In Situ Rock Bolt Pull Tests in HardRock Mines

by

Luke Nicholson

A thesis submitted in conformity with the requirementsfor the degree of Masters of Applied ScienceGraduate Department of Civil Engineering

University of Toronto

© Copyright 2016 by Luke Nicholson

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

2016

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

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 test-

ing, 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.

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Contents

1 Introduction 1

1.1 Problem Definition . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1

1.2 Significance . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2

1.3 Objectives . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2

1.4 Methodology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2

1.5 Structure of the Thesis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2

2 Testing of Rock Bolts 4

2.1 Rock Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4

2.2 Laboratory Testing Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5

2.2.1 Tension Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5

2.2.2 Wedge Tension Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6

2.2.3 Bend Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7

2.2.4 Tests of Expandable Rock Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7

2.2.5 Laboratory Rock Anchor Capacity Pull Test . . . . . . . . . . . . . . . . . . . . . 8

2.2.6 Laboratory Drop Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

2.3 In Situ Pull Test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 9

2.4 Discussion of Testing Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12

2.5 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13

3 Composition of the Database 14

3.1 Pull Test Setting . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14

3.1.1 Regional Geology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15

3.1.2 Coleman Mine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16

3.1.3 Copper Cliff Mine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16

3.1.4 Creighton Mine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 17

3.1.5 Garson Mine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18

3.1.6 Stobie Mine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18

3.1.7 Totten Mine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18

3.2 Reinforcement Elements in the Pull Test Database . . . . . . . . . . . . . . . . . . . . . . 19

3.2.1 Friction Rock Stabilizers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19

3.2.2 Rebar Rock Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21

3.2.3 Modified Cone Bolt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

3.2.4 D-Bolt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

iv

3.2.5 Expandable Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23

3.3 Database . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24

3.3.1 Database Description . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24

3.3.2 Comparison to Other Pull Test Databases . . . . . . . . . . . . . . . . . . . . . . . 26

3.3.3 Specific and General Limitations of the Pull Test Database . . . . . . . . . . . . . 26

3.4 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27

4 Review of Implemented Pull Test Methods 28

4.1 Implementation of Pull Tests in Practice . . . . . . . . . . . . . . . . . . . . . . . . . . . . 28

4.2 Deviations from ASTM Standard D4435-13 and ISRM Suggested Methods for Rockbolt

Testing . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29

4.2.1 Deviations in Apparatus . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30

4.2.2 Deviations in Procedure . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30

4.2.3 Practical Considerations in Pull Testing . . . . . . . . . . . . . . . . . . . . . . . . 31

4.3 Pull Test Data . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 32

4.3.1 Working Capacity and the Measurement of Load . . . . . . . . . . . . . . . . . . . 32

4.3.2 Recording Displacement During a Pull Test . . . . . . . . . . . . . . . . . . . . . . 34

4.3.3 Rock Bolt Stiffness . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 36

4.3.4 Limitations of the Metrics Measured in a Rock Bolt Pull Test . . . . . . . . . . . . 39

4.4 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 39

5 Summary Statistics and Interpretation of Pull Test Data 40

5.1 Summary Statistics and Statistical Techniques . . . . . . . . . . . . . . . . . . . . . . . . 40

5.1.1 Summary Statistics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 40

5.1.2 Statistical Techniques . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41

5.2 Friction Rock Stabilizers . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 43

5.2.1 Theoretical Behaviour of a Frictional Rock Stabilizer . . . . . . . . . . . . . . . . . 43

5.2.2 Observed Behaviour of Friction Rock Stabilizers . . . . . . . . . . . . . . . . . . . 45

5.2.3 Characterization of Performance Metrics for Friction Rock Stabilizers . . . . . . . 46

5.3 Rebar Rock Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50

5.3.1 Theoretical Behaviour of a Rebar Rock Bolt . . . . . . . . . . . . . . . . . . . . . . 50

5.3.2 Observed Behaviour of Rebar Rock Bolts . . . . . . . . . . . . . . . . . . . . . . . 51

5.3.3 Characterization of Performance Metrics for Rebar Rock Bolts . . . . . . . . . . . 52

5.4 Modified Cone Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 56

5.4.1 Theoretical Behaviour of a Modified Cone Bolt . . . . . . . . . . . . . . . . . . . . 56

5.4.2 Observed Behaviour of Modified Cone Bolts . . . . . . . . . . . . . . . . . . . . . . 57

5.4.3 Proposed Interpretation of Modified Cone Bolt Behaviour . . . . . . . . . . . . . . 59

5.4.4 Characterization of Performance Metrics for Modified Cone Bolts . . . . . . . . . . 61

5.5 D-Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63

5.5.1 Theoretical Behaviour of a D-Bolt . . . . . . . . . . . . . . . . . . . . . . . . . . . 63

5.5.2 Observed Behaviour of D-Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 64

5.5.3 Characterization of Performance Metrics for D-Bolts . . . . . . . . . . . . . . . . . 67

5.6 Expandable Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 69

5.6.1 Theoretical Behaviour of an Expandable Bolt . . . . . . . . . . . . . . . . . . . . . 69

v

5.6.2 Observed Behaviour of Expandable Bolts . . . . . . . . . . . . . . . . . . . . . . . 69

5.6.3 Characterization of Performance Metrics for Expandable Bolts . . . . . . . . . . . 71

5.7 Other Reinforcement Elements . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 73

5.7.1 Yield-Lok . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 73

5.7.2 Fibreglass Rebar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 74

5.7.3 DS Bolt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75

5.7.4 Other Expandable Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75

5.7.5 MD Bolt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75

5.8 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 76

6 Factors Influencing Pull Test Performance 78


6.1.1 Influence of Length . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 78

6.1.2 Installation Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83

6.1.3 Influence of Drive Time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85

6.1.4 Influence of Drill Bit Diameter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88

6.1.5 Influence of Bolt Diameter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 91

6.1.6 Geology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 93

6.1.7 Rock Mass Quality . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 97

6.1.8 Summary of Investigation on FRS Pull Tests . . . . . . . . . . . . . . . . . . . . . 99

6.2 Rebar Rock Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 100

6.2.1 Rebar Rock Bolt Length . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 100

6.2.2 Encapsulation Length . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 101

6.2.3 Spin Time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103

6.2.4 Residence Time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 104

6.2.5 Geology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 105

6.2.6 Summary of Investigation on Rebar Rock Bolt Pull Tests . . . . . . . . . . . . . . 106

6.3 Modified Cone Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 107

6.3.1 Residence Time . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 107

6.3.2 Geology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 108

6.3.3 Inter-variable Relationships . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109

6.3.4 Summary of Cone Bolt Findings . . . . . . . . . . . . . . . . . . . . . . . . . . . . 110

6.4 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111

7 Characterization of Rock Bolt Behaviour 112

7.1 Characterisations of Bolt Behaviour using Laboratory Pull Tests . . . . . . . . . . . . . . 112


7.2.1 Characterization of FRS Performance . . . . . . . . . . . . . . . . . . . . . . . . . 114

7.2.2 Characterization of FRS Behaviour . . . . . . . . . . . . . . . . . . . . . . . . . . . 117

7.3 Rebar Rock Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 120

7.3.1 Characterization of Rebar Rock Bolt Performance . . . . . . . . . . . . . . . . . . 121

7.3.2 Characterisation of Rebar Rock Bolt Behaviour . . . . . . . . . . . . . . . . . . . . 121

7.4 Modified Cone Bolt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 124

7.4.1 Characterisation of Modified Cone Bolt Performance . . . . . . . . . . . . . . . . . 124

vi

7.4.2 Characterisation of Modified Cone Bolt Behaviour . . . . . . . . . . . . . . . . . . 125

7.5 D-Bolt . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 127

7.5.1 Characterisation of D-Bolt Performance . . . . . . . . . . . . . . . . . . . . . . . . 127

7.5.2 Characterisation of D-Bolt Behaviour . . . . . . . . . . . . . . . . . . . . . . . . . 129

7.6 Expandable bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 132

7.6.1 Characterisation of Expandable Bolt Performance . . . . . . . . . . . . . . . . . . 132

7.6.2 Characterisation of Expandable Bolt Behaviour . . . . . . . . . . . . . . . . . . . . 133

7.7 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 136

8 Conclusions 140

8.1 Contributions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 140

8.2 Load Capacities of Reinforcement Elements . . . . . . . . . . . . . . . . . . . . . . . . . . 141

8.3 Limitations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 142

8.4 Recommendations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 142

8.5 Implications and Path Forward . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 144

List of References 144

Appendices 151

A Pull Testing Forms 152

A.1 ASTM D4435-13 Sample Form . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 152

A.2 ISRM Suggested Method for Pull Testing Data Sheet . . . . . . . . . . . . . . . . . . . . . 153

A.3 Proposed Pull Test Information Sheets . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 154

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List of Figures

2.1 Types of rock bolt (Hadjigeorgiou & Charette, 2001) . . . . . . . . . . . . . . . . . . . . . 4

2.2 Tensile test specimen with a reduced section (ASTM A370, 2012) . . . . . . . . . . . . . . 6

2.3 Apparatus for a tensile test on an FRS (ASTM F432, 2013) . . . . . . . . . . . . . . . . . 6

2.4 Bolt head configuration for a wedge tension test (ASTM F606, 2013) . . . . . . . . . . . . 7

2.5 Two alternative configurations for the ferrule test (ASTM F432, 2013) . . . . . . . . . . . 8

2.6 Apparatus for the Laboratory Rock Anchor Capacity Pull Test (ASTM D7401, 2008) . . 8

2.7 Apparatus for the Laboratory Drop Test (ASTM D7401, 2008) . . . . . . . . . . . . . . . 9

2.8 Apparatus for a rock bolt anchor pull test (ASTM D4435, 2013) . . . . . . . . . . . . . . 10

2.9 Conceptual load versus bolt head deflection curve for a rock bolt pull test (ASTM D4435,

2013) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 11

2.10 Apparatus for a pull test (ISRM, 1981) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 12

3.1 Map of Sudbury area showing locations of relevant mines (after Eckstrand & Hulbert, 2007) 14

3.2 Longitudinal section of Coleman Mine (Morissette et al, 2014) . . . . . . . . . . . . . . . . 16

3.3 Longitudinal section of Copper Cliff Mine (Chinnasane et al, 2014) . . . . . . . . . . . . . 17

3.4 Longitudinal section of Creighton Mine (Snelling et al, 2013) . . . . . . . . . . . . . . . . 17

3.5 Longitudinal section of Stobie Mine . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18

3.6 Longitudinal section of Totten Mine (After Sudbury Platinum Corporation, 2015) . . . . 19

3.7 FRS A schematic (Courtesy of Supplier A) . . . . . . . . . . . . . . . . . . . . . . . . . . 20

3.8 FRS B schematic (Courtesy of Supplier B) . . . . . . . . . . . . . . . . . . . . . . . . . . . 20

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) . . . 21

3.10 MCB33 (Courtesy of Mansour) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

3.11 D-Bolt schematic (Normet, 2014) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

3.12 VersaBolt schematic (Courtesy of Mansour) . . . . . . . . . . . . . . . . . . . . . . . . . . 23

3.13 Schematic of a Python bolt, and cross sections before (1) and after (2) inflation (Courtesy

of Jennmar) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 23

4.1 Normet’s pull test apparatus mounted on a 22 mm D-Bolt . . . . . . . . . . . . . . . . . . 29

4.2 Static laboratory pull test on a 22 mm D-Bolt 2.1 m in length (left; Doucet & Voyzelle,

2012) and dynamic impact test on a 22 mm D-Bolt 1.5 m in length (right; Li & Doucet,

2012) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 32

4.3 Methods of determining yield strength: halt of the pointer method (left) and offset method

(right; ASTM E6, 2009) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 33

viii

4.4 Determination of working capacity from a pull test . . . . . . . . . . . . . . . . . . . . . . 34

4.5 Measurement of displacement for a pull test on a generic reinforcement element . . . . . . 35

4.6 Generic reinforcement system (Thompson et al, 2012) . . . . . . . . . . . . . . . . . . . . 35

4.7 Measurement of displacement for axial loading of a point anchored rock bolt . . . . . . . . 36

4.8 Design of support systems using the ground reaction curve (Brady & Brown, 2006) . . . . 37

4.9 Calculation of secant and tangent stiffness from a pull test . . . . . . . . . . . . . . . . . . 38

4.10 Pull test campaign results for partially encapsulated Rebar A from November 16th, 2012

at Coleman Mine (Mainville et al, 2012) . . . . . . . . . . . . . . . . . . . . . . . . . . . . 38

5.1 Shear stress distribution along a frictionally coupled bolt subject to axial load (Li &

Stillborg, 1999) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44

5.2 Shear stress and axial load along a Swellex rock bolt (Li & Stillborg, 1999) . . . . . . . . 44

5.3 Pull test performed on an FA39 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45

5.4 Ultimate capacity per unit length distributions for FRSs with nominal diameters of 35,

39 and 46 mm . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47

5.5 Stiffness metrics for an FRS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48

5.6 Stiffness distributions for the FA35 and FA39 . . . . . . . . . . . . . . . . . . . . . . . . . 49

5.7 Model of the shear stress profile in a grouted rock bolt (Li & Stillborg, 1999) . . . . . . . 50

5.8 Model of tensile load and shear stress profile for a rebar (Li & Stillborg, 1999) . . . . . . 51

5.9 Pull test performed on a 20 mm Rebar B . . . . . . . . . . . . . . . . . . . . . . . . . . . 51

5.10 Pull test performed on a 20 mm Rebar A . . . . . . . . . . . . . . . . . . . . . . . . . . . 52

5.11 Working capacity distributions for rebar . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53

5.12 Secant stiffness for 20 mm rebar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54

5.13 Tangent stiffness for 20 mm Rebar . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 54

5.14 Comparison of Rebar A unloading stiffness to secant (a) and tangent (b) stiffness . . . . . 56

5.15 Laboratory pull test of an MCB33 (Simser et al, 2006) . . . . . . . . . . . . . . . . . . . . 57

5.16 Conceptual load–displacement behaviour of a cone bolt subject to quasi-static loading

(Simser et al, 2006) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57

5.17 Pull test performed on an MCB33 with plough . . . . . . . . . . . . . . . . . . . . . . . . 58

5.18 Pull test performed on an MCB33 without a linear plough response . . . . . . . . . . . . . 58

5.19 Close-up of the bolt response shown in Figure 5.18 . . . . . . . . . . . . . . . . . . . . . . 59

5.20 Amended conceptual load–displacement behaviour of a cone bolt . . . . . . . . . . . . . . 60

5.21 Performance metrics measured from a cone bolt pull test . . . . . . . . . . . . . . . . . . . 60

5.22 Load metric distributions for the MCB33 . . . . . . . . . . . . . . . . . . . . . . . . . . . 61

5.23 Stiffness metric distributions for the MCB33 . . . . . . . . . . . . . . . . . . . . . . . . . . 62

5.24 Apparatus for a simulated joint laboratory test on a D-Bolt (Li, 2012) . . . . . . . . . . . 63

5.25 Results of simulated joint laboratory tests on 20 mm D-Bolts (Li, 2012) . . . . . . . . . . 63

5.26 Pull tests performed on 22 mm D-Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . 65

5.27 Pull tests performed on 20 mm D-Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . 66

5.28 Pull tests performed on 22mm D-Bolts . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 67

5.29 Distributions of stiffness for 20 and 22 mm D-Bolts . . . . . . . . . . . . . . . . . . . . . . 68

5.30 Pull tests performed on Pm12 and Mn12 expandable bolts . . . . . . . . . . . . . . . . . . 69

5.31 Working capacities of Swellex Pm12 and Mn12 . . . . . . . . . . . . . . . . . . . . . . . . 71

5.32 Secant stiffness of Swellex variants . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 72

ix

5.33 Secant stiffness of Swellex sorted by installation medium . . . . . . . . . . . . . . . . . . . 73

6.1 Relationship between ultimate capacity and length for the FA35 . . . . . . . . . . . . . . 79

6.2 Relationship between ultimate capacity and length for the FB35 . . . . . . . . . . . . . . 80

6.3 Relationship between ultimate capacity and length for the FB39 . . . . . . . . . . . . . . 80

6.4 Relationship between ultimate capacity and anchorage length for all FRS bolts . . . . . . 82

6.5 Comparison of ultimate capacities between jackleg and bolter installations of the FA35 . . 83

6.6 Comparison of ultimate capacities between jackleg and bolter installations of all FRS bolts 84

6.7 Relationship between drive time and ultimate capacity for all FRS bolts . . . . . . . . . . 85

6.8 Relationship between drive time and ultimate capacity for FB35, distinguishing between

installation methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 86

6.9 Relationship between drive time and absolute ultimate capacity for all FRS configurations

installed using a bolter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87

6.10 Relationship between installation time and ultimate capacity for all bolter-installed FRS

fit with linear and power functions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88

6.11 Relationship between drill bit diameter and ultimate capacity for all FRS configurations . 89

6.12 Relationship between drill bit diameter and ultimate capacity for two testing campaigns

performed on the FB46 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90

6.13 Relationship between drill bit diameter and ultimate capacity for FA35 and FB35, sepa-

rated by installation method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 90

6.14 Bolt diameter to drill bit diameter ratios for all FRS variants . . . . . . . . . . . . . . . . 91

6.15 Relationship between bolt diameter to drill bit diameter ratio and ultimate capacity . . . 92

6.16 Relationship between bolt diameter to drill bit diameter ratio and ultimate capacity for

all FRS configurations by testing campaign . . . . . . . . . . . . . . . . . . . . . . . . . . 93

6.17 FRS ultimate capacity by lithology . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 94

6.18 Relationship between UCS and ultimate capacity . . . . . . . . . . . . . . . . . . . . . . . 95

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 . . . . . . . . . . . . . . . . . . 96

6.20 Ultimate capacities recorded for FA46 bolts installed in poor and good quality ground . . 97

6.21 Ultimate capacities recorded for FRSs installed in poor and good quality ground . . . . . 98

6.22 FA39s pulled at Garson, 9/20/2013 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 98

6.23 Rebar performance metrics by length . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 101

6.24 Relationships between stiffness and grout length : rebar length for Suppliers A and B . . 102

6.25 Unloading stiffness and grout length : rebar length for Rebar A . . . . . . . . . . . . . . . 103

6.26 Relationship between stiffness and resin spin time for Rebar A . . . . . . . . . . . . . . . 104

6.27 Stiffness comparison of Rebar A installed on the day of testing versus previously . . . . . 104

6.28 Stiffness comparison between lithologies for rebar . . . . . . . . . . . . . . . . . . . . . . . 105

6.29 Performance comparison of MCB33s installed prior to and on the day of testing . . . . . . 107

6.30 Performance comparison of MCB33 bolts installed in ore, igneous/metaigneous and metased-

imentary lithologies . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109

6.31 Relationship between plough stiffness and plough point for the MCB33 . . . . . . . . . . . 110

7.1 Average load–displacement behaviours obtained by Stillborg in a laboratory setting (Still-

borg, 1993; composited by Hoek et al, 1995) . . . . . . . . . . . . . . . . . . . . . . . . . . 113

x

7.2 Distributions of secant stiffness and ultimate capacity for a pull test on an FRS with an

anchorage length of 1.52 m . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 115

7.3 Ultimate capacity per metre distributions for all FRS configurations . . . . . . . . . . . . 116

7.4 Distribution of secant stiffness and ultimate capacity for all FRS bolts tested . . . . . . . 116

7.5 All FA35 and FA39 pull test load-displacement relationships . . . . . . . . . . . . . . . . . 117

7.6 Distribution of displacement measured during pull testing of FA35 and FA39 with anchor-

age lengths of 1.52 m . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118

7.7 Conceptual displacement distribution of FRSs with anchorage lengths of 1.52 m subject

to a pull test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 118

7.8 Conceptual displacement distribution of pull tests performed on FRSs . . . . . . . . . . . 119

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 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 121

7.10 All rebar pull test load-displacement relationships . . . . . . . . . . . . . . . . . . . . . . . 122

7.11 Distribution of displacement measured during pull testing of 20 mm rebar 1.8 m to 2.4 m

in length with a pre-load of 17.8 kN . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 122

7.12 Distribution of displacement measured during pull testing of 20 mm rebar 1.8 m to 2.4 m

in length without a pre-load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 123

7.13 Conceptual distribution of displacement during pull testing of 20 mm rebar without a

pre-load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 123

7.14 Distributions of secant stiffness and working capacity for a pull test with a pre-load of

17.8 kN on an MCB33 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 125

7.15 56 MCB33 pull tests collected from Vale’s Sudbury operations . . . . . . . . . . . . . . . . 125

7.16 Distribution of displacement measured during pull testing of a 2.4 m MCB33 with a

pre-load of 17.8 kN . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 126

7.17 Conceptual distribution of displacements for a pull test of a 2.4 m MCB33 without a

pre-load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 127

7.18 Distribution of secant stiffness for a pull test with a pre-load of 17.8 kN on a 20 mm D-Bolt128

7.19 Distribution of secant stiffness for a pull test with a pre-load of 17.8 kN on a 22 mm D-Bolt128

7.20 20 mm D-Bolt pull tests collected from Vale’s Sudbury operations . . . . . . . . . . . . . 129

7.21 22 mm D-Bolt pull tests collected from Vale’s Sudbury operations . . . . . . . . . . . . . 129

7.22 Load–displacement behaviour of a 20 mm D-Bolt with a pre-load of 17.8 kN . . . . . . . . 130

7.23 Load–displacement behaviour of a 22 mm D-Bolt with a pre-load of 17.8 kN . . . . . . . . 130

7.24 Conceptual distribution of displacements for a pull test of a 20 mm D-Bolt without a

pre-load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131

7.25 Conceptual distribution of displacements for a pull test of a 22 mm D-Bolt without a

pre-load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131

7.26 Distributions of secant stiffness and working capacity for a pull test on Swellex Pm12 and

Mn12 without a pre-load . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 132

7.27 Swellex Pm12 and Mn12 pull tests collected from Vale’s Sudbury operations . . . . . . . . 133

7.28 Swellex Pm24 and Mn24 pull tests collected from Vale’s Sudbury operations . . . . . . . . 134

7.29 Load–displacement behaviour of Pm12 and Mn12 bolts . . . . . . . . . . . . . . . . . . . . 134

7.30 Load-displacement behaviour of Pm24 and Mn24 bolts . . . . . . . . . . . . . . . . . . . . 135

7.31 Conceptual load–displacement behaviour of Pm12 and Mn12 bolts subject to a pull test . 135

xi

7.32 Conceptual load–displacement behaviour of Pm24 and Mn24 bolts subject to a pull test . 136

7.33 Conceptual load–displacement behaviour extrapolated to 0 load for various rock bolts

subject to a pull test . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 137

7.34 Median load–displacement behaviour for all bolts with no pre-load . . . . . . . . . . . . . 138

A.1 Rock bolt pull test sample form (ASTM D4435, 2013) . . . . . . . . . . . . . . . . . . . . 152

A.2 Rock bolt pull test data sheet (After ISRM, 1981) . . . . . . . . . . . . . . . . . . . . . . 153

A.3 Proposed pull test campaign information sheet . . . . . . . . . . . . . . . . . . . . . . . . 155

A.4 Proposed pull test data sheet . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 156

xii

List of Tables

3.1 Intact rock UCS values in MPa by mine site . . . . . . . . . . . . . . . . . . . . . . . . . . 15

3.2 Comparison of FRS supplier specifications (Courtesy of Suppliers A and B) . . . . . . . . 21

3.3 Steel properties for a 20 mm threaded rebar (Courtesy of Suppliers A, B and C) . . . . . 21

3.4 MCB33 mechanical properties (Courtesy of Mansour) . . . . . . . . . . . . . . . . . . . . 22

3.5 VersaBolt and D-Bolt mechanical properties (Courtesy of Mansour, Normet) . . . . . . . 23

3.6 Summary of Swellex, Omega and Python bolt mechanical properties (Atlas Copco, 2012;

Courtesy of Jennmar, DSI) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 24

3.7 Number of pull tests by bolt type . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25

5.1 General One-Way ANOVA table (NIST-SEMATECH,2003) . . . . . . . . . . . . . . . . . 43

5.2 Statistics regarding the ultimate capacities of FRSs . . . . . . . . . . . . . . . . . . . . . . 46

5.3 FRS stiffness summary statistics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50

5.4 Summary statistics for the working capacity of rebar . . . . . . . . . . . . . . . . . . . . . 53

5.5 Summary statistics for the stiffness of rebar . . . . . . . . . . . . . . . . . . . . . . . . . . 55

5.6 Comparison of stiffness between partially and fully encapsulated rebar . . . . . . . . . . . 55

5.7 Summary statistics of MCB33 load metrics . . . . . . . . . . . . . . . . . . . . . . . . . . 61

5.8 Summary statistics of MCB33 stiffness metrics . . . . . . . . . . . . . . . . . . . . . . . . 62

5.9 D-Bolt stiffness by anchor stability . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 64

5.10 Unloading stiffness calculated for pull tests performed on 22 mm D-Bolts . . . . . . . . . 66

5.11 Working capacities obtained from all D-Bolt pull tests . . . . . . . . . . . . . . . . . . . . 67

5.12 D-Bolt summary statistics . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 68

5.13 Swellex behaviour breakdown . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 70

5.14 Coupling strength of partially embedded and slipped Swellex pull tests . . . . . . . . . . . 70

5.15 Summary statistics of Swellex Pm12 and Mn12 working capacity . . . . . . . . . . . . . . 71

5.16 Swellex secant stiffness summary statistics . . . . . . . . . . . . . . . . . . . . . . . . . . . 72

5.17 Secant stiffness summary statistics on Swellex sorted by installation medium . . . . . . . 73

5.18 8’ (2.44 m) Yield-Lok pull test result summary . . . . . . . . . . . . . . . . . . . . . . . . 74

5.19 DS Bolt campaign summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 75

5.20 Summary of expandable bolt campaigns not including Swellex . . . . . . . . . . . . . . . . 75

5.21 Summary of working capacities for all bolts pull tested . . . . . . . . . . . . . . . . . . . . 77

6.1 Statistics on the relationship between ultimate capacity and length for the FA35 . . . . . 79

6.2 Statistics on the relationship between ultimate capacity and length for the FB35 . . . . . 80

6.3 Statistics on the relationship between ultimate capacity and length for the FB39 . . . . . 81

xiii

6.4 Single factor ANOVA performed on the relationship between ultimate capacity and length

for the FB39 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 81

6.5 Breakdown of bolts contributing to major sets of anchorage length . . . . . . . . . . . . . 81

6.6 Comparison of 1.52 m and 1.83 m of anchorage length for all FRS bolts . . . . . . . . . . 82

6.7 Ultimate capacity statistics for jackleg and bolter installations of the FA35 . . . . . . . . 83

6.8 Ultimate capacity statistics for jackleg and bolter installation of all FRS bolts . . . . . . . 84

6.9 Description of relationships between drive time and ultimate capacity for FRSs installed

with a MacLean Bolter . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 86

6.10 Summary statistics for bolt diameter to drill bit diameter ratios for all FRS variants . . . 91

6.11 Comparison of average ultimate capacities for pull tests performed in ore and waste rock

in the same campaign . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95

6.12 Comparison of FRS bolts installed in ore and waste rock . . . . . . . . . . . . . . . . . . . 96

6.13 Comparison of FA46 bolts installed in poor and good quality ground . . . . . . . . . . . . 97

6.14 Comparison of all FRSs installed in poor and good quality ground . . . . . . . . . . . . . 99

6.15 Comparison of partial and full encapsulation test statistics for Rebar A and B . . . . . . . 102

6.16 Statistics regarding residence time for Rebar A . . . . . . . . . . . . . . . . . . . . . . . . 104

6.17 Comparison of stiffness across different lithologies . . . . . . . . . . . . . . . . . . . . . . . 106

6.18 Performance comparison of MCB33s installed prior to and on the day of testing . . . . . . 107

6.19 Performance comparison of MCB33 bolts installed previously and on the day of testing . 108

6.20 Summary of observed relationships of between various factors and performance indicators

for each rock bolt type . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 111

7.1 FA35 and FA39 performance percentiles . . . . . . . . . . . . . . . . . . . . . . . . . . . . 114

7.2 Distribution of ultimate capacity across all FRS configurations . . . . . . . . . . . . . . . 115

7.3 FRS performance compared to supplier specifications . . . . . . . . . . . . . . . . . . . . . 120

7.4 Percentiles of performance metrics for rebar rock bolts . . . . . . . . . . . . . . . . . . . . 121

7.5 Rebar performance compared to supplier specifications . . . . . . . . . . . . . . . . . . . . 124

7.6 Distributions of MCB33 performance metrics . . . . . . . . . . . . . . . . . . . . . . . . . 124

7.7 MCB33 performance compared to supplier specifications . . . . . . . . . . . . . . . . . . . 126

7.8 Secant stiffness distributions of 20 mm and 22 mm D-Bolts . . . . . . . . . . . . . . . . . 127

7.9 Swellex secant stiffness percentiles . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133

7.10 Swellex Pm12 and Mn12 performance compared to supplier specifications . . . . . . . . . 136

xiv

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, 2012). 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 (Hadjige-

orgiou & Charette, 2001). 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 10) 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

Chapter 1. Introduction 2

1.2 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

2011 and 2014.

2. 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 method-

ology 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 pre-

sented. Standardized testing methods used to characterize rock bolt behaviour are identified, including

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 per-

formance 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 limita-

tions of the current work are identified, and recommendations are made for future work.

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, 2001). 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).

MechanicalBolts

Resin Bolts

Paddle bolt

Posimix

Solid deformed

Wriggle bolt

Chemicalanchor

Friction bolts

Split set

Swellex

Figure 2.1: Types of rock bolt (Hadjigeorgiou & Charette, 2001)

4

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, 2001).

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, 2014). 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, 2012).

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.

2.2.1 Tension Test

The American Society for Testing and Materials (ASTM) Standard F432-13 is ASTM’s primary rock

bolt testing standard, describing the “chemical, mechanical and dimensional requirements for roof and

rock bolts and accessories” (ASTM F432, 2013). It references and specifies a range of testing methods,

the first of which is the tension test. ASTM F432-13 dictates that “test bars for the manufacture of bolts

and threaded and threaded slotted bars” (ASTM F432, 2013) must be in accordance with ASTM A370-

12, which describes procedures and definitions for “the mechanical testing of steels, stainless steels, and

related alloys” (ASTM A370, 2012). 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.


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 A370, 2012)

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, 2013).

Figure 2.3: Apparatus for a tensile test on an FRS (ASTM F432, 2013)

2.2.2 Wedge Tension Test

The wedge tension test, described in ASTM Standard F606-13, is to be used to test bolts and threaded

bars in conjunction with the nut intended for practical use. A 10◦ 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


yield strength is obtained using the “drop of the beam” method (ASTM D4435, 2013), 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 F606, 2013).

Figure 2.4: Bolt head configuration for a wedge tension test (ASTM F606, 2013)

2.2.3 Bend Test

The bend test as described by ASTM Standard F432-13 is only applicable to notched bendable bolts.

Bolts are subject to one bending cycle, i.e. are bent 90◦ 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 A615-15 corresponding

to the steel grade and bar diameter, plus 6000 lbf (26.7 kN). After fracture of the bolt in the tensile test

(which must occur above a load of 23000 lbf; 102.3 kN), there should be no evidence of prior fracture

induced by the bend test (ASTM F432, 2013).

2.2.4 Tests of Expandable Rock Bolts

ASTM Standard F432-13 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


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, 2013)

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 A370-12.

2.2.5 Laboratory Rock Anchor Capacity Pull Test

The laboratory rock anchor capacity pull test is one of two tests described in ASTM Standard D7401-08.

It may be applied to elements with “mechanical, resin or other similar anchor systems” (ASTM D7401,

2008). 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 D7401, 2008).

Figure 2.6: Apparatus for the Laboratory Rock Anchor Capacity Pull Test (ASTM D7401, 2008)


2.2.6 Laboratory Drop Test

The drop test is the second test described in ASTM Standard D7401-08. 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 D7401, 2008).

Figure 2.7: Apparatus for the Laboratory Drop Test (ASTM D7401, 2008)

2.3 In Situ Pull Test

ASTM Standard D4435-13 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 D4435-13: Standard Test Method for Rock Bolt Anchor Pull Test

ASTM D4435-13 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. 10 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 ca-

pacity to fail the bolt and induce at least 50 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


be measured to an accuracy of at least 0.025 mm and resolution of 0.013 mm. The displacement trans-

ducer must be supported from a point no closer than 0.9 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 0.25 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 0.5◦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, 2013)

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, 10 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 20 equal load increments, or increments of 2.2 kN, whichever is less. Displacement is recorded

to the nearest 0.013 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, 2013).

As per ASTM D4435-13, 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,

2013).


Figure 2.9: Conceptual load versus bolt head deflection curve for a rock bolt pull test (ASTM D4435,2013)

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 50 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 0.05 mm, to be mounted on “firm rock” (ISRM, 1981). Figure 2.10 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 0 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 5-10 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 40 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).


Figure 2.10: 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-nut-

plate 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 D7401-08 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, 2014). 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


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 D4435-13

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.

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.

NickelOffset

Milnet N

North RangeShaft

WhistleZ AZkm

Capreol

Chelmsford

Longvac

StrathconaColeman

Fecunis

Fecunis LakeLevack

Levack WestBoundary

HardyWindy

Lw

Clarabelle Copper Cliff SouthCopper Cliff North

Little StobieStobie

FroodSUDBURY

Garson

LEGEND

Granophyre

Quartzxrich gabbro

Norite

SublayerSudburyIgneaous

Complex Chemsford Formation

Onwatin Formation

Onaping Formation

CreightonI Murray granites

Quartzite

GreywackeI volcanic rocks

South Range Shear Zone FaultOlivine diabase dykes

Norduna

RamseyLake

FalconbridgeEast Falconbridge

L w

L w

L w

L w

L w

Fraser

McCreedy

Granite gneiss and plutonsArchean

Proterozoic

NixCuxPGE deposits

Murray

WorthingtonVictoria

ChicagoSultana

Collins

Trillabelle

Creighton

Totten

Figure 3.1: Map of Sudbury area showing locations of relevant mines (after Eckstrand & Hulbert, 2007)

14

Chapter 3. Composition of the Database 15

3.1.1 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 30 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, 2009), resulting in a

200 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, 2009).

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 700 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, 2009). Vale’s Sudbury mines, described in Sections 3.1.2 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 TottenOre 130 110-150 130 91 155 170Granite Gneiss 240Mafic Gneiss 280Granite Breccia 210Sudbury Breccia 300Olivene Diabase DykeQuartz Diorite 150 270Metasediments 150-170 133Granite 240 240Norite 150 155 191Trap Dyke 200-240 260Amphibolite 200-240Greenstone 209Metabasalt 248Conglomerate 223Metagabbro 190.4


3.1.2 Coleman Mine

Located in the North Range, 30 km north-west of Sudbury, Coleman Mine entered production in 1970

and is now one of Vale’s most productive mines. Four ore bodies (Main Ore Body, West Ore Body,

153 Ore Body and 170 Ore Body) are currently being exploited (Vale, 2011), extracting a combined

1,515,000 metric tons of ore in 2013 (Vale, 2013). Several faults and an olivine-diabase dyke contribute

to the challenge of managing seismicity on site. Up until 2010, 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, 2011). Figure 3.2 shows a longitudinal section of the mine. The 170 orebody

extends below 5700 feet (1740 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 2010 and March 2015.

153 OB MOB

170 OB

Figure 3.2: Longitudinal section of Coleman Mine (Morissette et al, 2014)

3.1.3 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 1960,

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, 2014)

mined with methods including blasthole, vertical retreat, slot and slash and uppers retreat (Vale, 2010).

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, 2014). 164 pull tests were obtained from Copper Cliff Mine between

October, 2008 and March, 2015. Figure 3.3 shows a longitudinal section of Copper Cliff Mine.


Figure 3.3: Longitudinal section of Copper Cliff Mine (Chinnasane et al, 2014)

3.1.4 Creighton Mine

Creighton Mine is one of the oldest operating mines in Ontario, having entered production in 1901 (Vale,

2014). It is located on the outer rim of the South Range. Currently, the majority of production comes

from the 7400 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, 2014). 294 pull tests were collected from Creighton

Mine, performed between January 2003 and January 2015. Figure 3.4 shows a longitudinal section of

Creighton Mine.

9LShaft11LShaft5LShaft3LShaft7LShaft

3 ShaftRamp

6 Shaft

8Shaft

461Orebody

Surface NWSE

0 2000m

20

00

N

800

0N

10000LL 3048Lm

5800LL 1768Lm

7000LL 2134Lm

7680LL 2341Lm

8200LL 2500Lm

3800LL 1158Lm

1000LL 305Lm

Footwall Orebodies

Orebodies

Hangingwall Orebodies

GraniteIGabbro Complex

Sudbury Igneous Complex

Shear Zone

Figure 3.4: Longitudinal section of Creighton Mine (Snelling et al, 2013)


3.1.5 Garson Mine

Garson Mine consists of two operations; the surface ramp, and the shaft-accessed main mine which ex-

tends down 5200 feet. Four main orebodies are present, coincident with four shear zones (Mukwakwami

et al, 2013). 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 600 m, and dip at about 70◦ (Abdellah et al, 2014). 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 2011 and March 2015.

3.1.6 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., 2003). Vale suspended operations at the Frood section of the

mine in 2012 (Carmichael, 2012), although Stobie remains a significant producer at present. 150 pull

tests were collected from Stobie Mine, performed between December 2011 and December 2014. Figure

3.5 shows a long section of Stobie Mine.

Figure 3.5: Longitudinal section of Stobie Mine

3.1.7 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 64,000 tons of ore in

2013 and will ramp up to full production in 2016 (Vale, 2015). The sulphide ore is found in an offset


quartz diorite dike dipping at 80◦, the geological model of which is often compared to that of the Copper

Cliff Offset (Lightfoot & Farrow, 2002). 113 tests from between April 2010 and December 2014 were

obtained.

Figure 3.6: Longitudinal section of Totten Mine (After Sudbury Platinum Corporation, 2015)

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.

3.2.1 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.


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 zinc-

based 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 2014). 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 2012, 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 2014). 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.


Table 3.2: Comparison of FRS supplier specifications (Courtesy of Suppliers A and B)

Minimum Ultimate Tensile Strength Initial Load Capacity Nominal Bit SizeFA33 71 kN 27 to 54 kN 30 to 33 mmFB33 89 kN 27 to 54 kN 31 to 33 mmFA35 71 kN 27 to 54 kN 32 to 35 mmFB35 89 kN 27 to 54 kN 31.8 to 33.3 mmFA39 89 kN 27 to 54 kN 35 to 38 mmFB39 102 kN 27 to 54 kN 35 to 38 mmFA46 133 kN 54 to 89 kN 41 to 44 mmFB46 145 kN 54 to 89 kN 41 to 45 mm

3.2.2 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 20 mm in diameter, with a limited number of tests on 22 mm

bolts. Both Suppliers A and C manufacture their rebar using Grade 60 steel (minimum yield strength

of 420 MPa; ASTM A615, 2015), while Supplier B uses grade 400W steel (minimum yield strength of

400 MPa; CSA G30.18, 2009). Figure 3.9 shows schematics of the three rebar, and Table 3.3 compares

the strengths of 20 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 20 mm threaded rebar (Courtesy of Suppliers A, B and C)

Gr. 60 Gr. 400 WMinimum Yield Strength 89 kN 86 kNMinimum Ultimate Tensile Strength 134 kN 116 kN

As shown in Table 3.3, the Grade 60 rebar (Suppliers A and C) has a marginally higher yield strength

as well as higher tensile strength than Supplier B’s Grade 400 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 10 UNC threads.


3.2.3 Modified Cone Bolt

The Modified Cone Bolt (Figure 3.10) 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, 2010). The MCB33

is manufactured using a modified C1055 carbon steel. Table 3.4 shows the minimum and typical yield

and ultimate loads at the thread of the bolt.

Figure 3.10: MCB33 (Courtesy of Mansour)

Table 3.4: MCB33 mechanical properties (Courtesy of Mansour)

Minimum TypicalYield Strength 98.5 kN 113.6 kNUltimate Tensile Strength 151.5 kN 163.6 kN

3.2.4 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, 2011). The D-Bolt is manufactured with a smooth bar diameter of either 20 mm or 22 mm,

with pull tests on both diameters of bolt present in the database.

Figure 3.11: D-Bolt schematic (Normet, 2014)


Mansour manufacture a similar, 20.5 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 (20 mm) D-Bolt (22 mm)Typical yield strength 138 kN 150 kN 190 kNTypical ultimate tensile strength 192 kN 210 kN 250 kN

3.2.5 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 (Courtesyof 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, 2012). 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


3.6 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, 2012;Courtesy of Jennmar, DSI)

Type Variant Breaking Load (kN) Elongation Material Thickness (mm)

Swellex Pm12 110 10% 2Mn12 110 20% 2

Omega 12 Tonnes Standard 120 10% 212 Tonnes Plus 115 20% 2

Python Standard 120 25% 2

Swellex Pm16 160 10% 2Mn16 150 20% 2

Omega 16 Tonnes Plus 150 20% 2Python Midi 160 25% 2

Swellex Pm24 240 10% 3Mn24 220 20% 3

Omega 24 Tonnes Standard 240 10% 324 Tonnes Plus 220 20% 3

Python Super 240 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, 2014).

3.3 Database

In total, pull test data for at least 26 bolt configurations from 7 suppliers across 5 classes of bolt was

collected.

3.3.1 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


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 NumberFA35 92

FRS A Supplier A FA39 30FA46 52FB35 100

FRS FRS B Supplier B FB39 101FB46 154

FRS C Supplier C FC35 6FC46 6

MD Bolt Sandvik N/A 10TOTAL 555Rebar A Supplier A 20 mm 52Rebar B Supplier B 20 mm 54

Grouted Bolts Rebar C Supplier C 20 mm 18(Static) Fibreglass Rebar FiReP N/A 10

Unspecified rebar 7TOTAL 141

Pm12 57Swellex Atlas Copco Pm24 36

Mn12 17Expandable Bolts Mn24 20

Omega DSI 12t 424t 1

Python Jennmar 16t 5TOTAL 140

Modified Cone Bolt MCB33 78MCB38 8

D-Bolt Normet 20 mm 15Yielding Bolts 22 mm 20

DS-Bolt Mansour 20.5 mm 4Yield-Lok Jennmar N/A 10TOTAL 135

Other Eyebolts in FA39 Supplier A N/A 11Mechanical 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 non-

galvanized. Given the size of the non-galvanized sample, it proved to be impossible to separate the


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.

3.3.2 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 900 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 (2000) collected 309 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 (2000) 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 50 mines in the analysis of Tomory et al; 1998, and

mine sites across North America and Europe in the case of Soni; 2000), 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 (2000) 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.

3.3.3 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 char-

acterization. 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


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 spec-

ifications 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.

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

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 cali-

brated 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 10 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 con-

tribute 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 displace-

ment 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 D4435-13 and ISRM Sug-

gested Methods for Rockbolt Testing

ASTM D4435-13 (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


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.

4.2.1 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 D4435-13, and have travel ranges exceeding

the specified 50mm, 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,

2013). 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 0.9 m from the reaction frame if supported by the face

(ASTM D4435, 2013). 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.

4.2.2 Deviations in Procedure

Both ASTM and ISRM require bolt installations representative of typical operational procedures. ASTM

suggests 10 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 20. 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.


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 10 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 20 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 40 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, 2013c) 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.

4.2.3 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 10. 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 500 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


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 proper-

ties 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.

4.3.1 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.

0

50

100

150

200

250

300

0 20 40 60 80 100 120 140 160 180Plate End Displacement (mm)

Loa

d (k

N)

0

50

100

150

200

250

300

0 50 100 150 200 250

Lo

ad

(kN

)

Displacement (mm)

Impact load

Plate load

Figure 4.2: Static laboratory pull test on a 22 mm D-Bolt 2.1 m in length (left; Doucet & Voyzelle,2012) and dynamic impact test on a 22 mm D-Bolt 1.5 m in length (right; Li & Doucet, 2012)


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, 2014). 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 elon-

gation of the material has commenced” (ASTM E6, 2009). Figure 4.3 shows two methods of determining

yield strength according to ASTM Standard E6-09b.

Figure 4.3: Methods of determining yield strength: halt of the pointer method (left) and offset method(right; ASTM E6, 2009)

In Figure 4.3, the yield strength (Sy; 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 0.5, 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


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.

Load

Displacement

Working Capacity

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.

4.3.2 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 dis-

placement 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 demon-

strated in Figure 4.5.

db 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. ds

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,

ds is a result of fracture compression and generation, and movement of the resulting blocks. da is the

apparent displacement, equal to the summation of db and ds. 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 D4435-13 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 db).


db

dads

db = displacement attributable to bolt deformation or movement

ds = displacement induced on the surface of the excavation

da = db + ds = apparent displacement

Figure 4.5: Measurement of displacement for a pull test on a generic reinforcement element

The measurement of da 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, 2012).

Figure 4.6: Generic reinforcement system (Thompson et al, 2012)

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 ds relative to the rest of the wedge’s surface. The bolt deforms, resulting in a

displacement db at the bolt head. The summation of these two displacements, da, 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 da of a pull test may reflect the da of such an axial loading

scenario, depending on factors such as plate size relative to the footprint of the pull test apparatus on


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 ds), and the length of bolt that will deform (influencing db).

dbds

da

da = db + ds

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.

4.3.3 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, 2008). A conceptual example is provided by Brady & Brown (2006), 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


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, 2006)

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.


0

20

40

60

80

100

120

140

160

0 1 2 3 4 5 6 7 8 9 10

Load

(kN

)

Displacement (mm)

SecantcStiffness

TangentcStiffness

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

4.10); 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 0 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 4.10 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 4.10: Pull test campaign results for partially encapsulated Rebar A from November 16th, 2012at Coleman Mine (Mainville et al, 2012)


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.

4.3.4 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.

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.

5.1.1 Summary Statistics

The mean of a dataset is the arithmetical average, expressed as x for a data set {x1, x2, ..., xi}, and is

calculated as

x = 1n

n∑i=1

xi

where n is the number of samples. The sample standard deviation (s) is also routinely quantified to

describe dispersion, calculated as

sx =

√∑ni=1(xi − x)2

n− 1

Both of these metrics are used to calculate the coefficient of variation for the sample (cv) as a way

of describing the size of the sample standard deviation relative to the mean of the data (Baecher &

Christian, 2003). This allows for the direct comparison of dispersion between variables with different

magnitudes of mean.

40

Chapter 5. Summary Statistics and Interpretation of Pull Test Data 41

cv =sxx

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 (G1).

G1 =n

(n− 1)(n− 2)

n∑i=1

(xi − xs

)3

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 G1. A normal distribution is

described by G1 = 0 (Doane & Seward, 2011). The second statistic used to describe distribution shape

is a measure of excess kurtosis, referred to as kurtosis herein, calculated as

Kurtosis =

(n(n+ 1)

(n− 1)(n− 2)(n− 3)

n∑i=1

(xi − xs

)4)− 3(n− 1)2

(n− 2)(n− 3)

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 0; kurtosis greater than 0 indicates a peakedness of the distribution, and kurtosis

below 0 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 (Yp) is calculated by setting

p(N + 1) = k + d

where k is an integer value and d the remaining decimal. There are three ways Yp is subsequently

calculated:

1. If 0 < k < N , Yp = Y|k| + d(Y|k+1| − Y|k|)2. If k = 0 , Yp = Y|1|

3. If k >= N , Yp = Y|N |

Note that the 50th percentile is the median of the data (NIST-SEMATECH, 2003).

5.1.2 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

yi = β0 + β1xi + εi

where

β1 =Sxy

Sxx=

∑ni=1(xi − x)(yi − y)∑n

i=1(xi − x)2

and

β0 = y − β1x


and εi 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

(R2) may be calculated for a linear regression, as

R2 =RSS

TSS=

∑ni=1(yi − yi)2∑ni=1(yi − y)2

where yi is the fitted value for the term β0 +β1xi. The coefficient of determination assesses the level

of correlation between y and x; R2 = 1 describes a perfectly linear relationship, and R2 = 0 indicates

no linear relationship is present (Fox, 2008).

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 (µ1 = µ2).

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 Y1 and Y2 with equal variance, and the t-test for populations with assumed

unequal variance. For the former, the t-statistic is calculated as

t =Y1 − Y2

sY1Y2

√1n1

+ 1n2

where

sY1Y2 =

√(n1 − 1)s2Y1

+ (n2 − 1)s2Y2

n1 + n2 − 2

Accounting for the degrees of freedom, calculated as

v = n1 + n2 − 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 (tcrit), the null hypothesis is rejected, suggesting a significant difference

between population means. In addition, a p-value is estimated. If t > tcrit, the p-value will be less than

the significance level chosen (NIST-SEMATECH, 2003).

If samples are assumed to have unequal variance, the same procedure is followed but the calculation

of t and v are as follows:

t =Y1 − Y2√

s21/n1 + s22/n2

v =(s21/n1 + s22/n2)2

(s21/n1)2/(n1 − 1) + (s22/n2)2/(n2 − 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.


Table 5.1: General One-Way ANOVA table (NIST-SEMATECH,2003)

Source Sum of Squares DoF Mean Square F0

Factor SSF = J∑

(yi. − y..)2 I − 1 MSF = SSF /(I − 1) MSF/MSE

Residual SSE =∑∑

(yij − yi.)2 I(J − 1) MSE = SSE/(I(J − 1))

Total SST =∑∑

(yij − y..)2 IJ − 1

where

yi. =1

J

J∑j=1

yij

and

y.. =1

IJ

I∑i=1

J∑j=1

yij

Similar to the t-test, the F0 value calculated is compared to the F -distribution, accounting for degrees

of freedom (DoF) and significance level. If F0 > Fcrit, the null hypothesis is rejected and a difference

between means is suggested at a predetermined level of significance (NIST-SEMATECH, 2003).

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.

5.2.1 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).


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 (x2), beyond which the shear

stress on the interface decreases following function τb(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.


5.2.2 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.

0

10

20

30

40

50

60

70

80

90

0 1 2 3 4 5 6 7 8 9 10

Load

m(kN

)

Displacementm(mm)

FE

DCBA

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 manu-

ally 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.


5.2.3 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” (0.152 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 cv Skewness Kurtosis Average Diameter

FA35 81 40.0 kN/m 10.3 kN/m 0.26 −0.18 −0.49 36.02 mmFA39 30 38.3 kN/m 7.2 kN/m 0.19 0.97 −0.05 39.11 mmFA46 50 39.1 kN/m 11.1 kN/m 0.28 0.12 0.88 45.23 mmFB35 92 40.0 kN/m 13.7 kN/m 0.34 −0.21 −0.24 34.89 mmFB39 83 37.9 kN/m 13.3 kN/m 0.35 0.46 −0.29 38.30 mmFB46 106 37.8 kN/m 10.8 kN/m 0.28 −0.36 −0.22 45.99 mm

ALL 442 38.9 kN/m 11.7 kN/m 0.30 0.00 -0.04 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 = 0.745 and 0.405 respectively), as

well as for the ANOVA of all configurations (p = 0.695). 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


100%

0

2

4

6

8

10

12

14

16

18

20

Cum

ulat

ive5

Fre

que

ncy

Fre

quen

cy

Ultimate5Capacity5(kN/m)

FrequencyCumulative

0 10 20 30 40 50 60 70 80

80%

60%

40%

20%

0%

(a) FA35

0

2

4

6

8

10

12

14

16 100%

Cum

ulat

ive5

Fre

que

ncy

80%

60%

40%

20%

0%

Ultimate5Capacity5(kN/m)0 10 20 30 40 50 60 70 80

Fre

quen

cy

(b) FB35

0%

20%

40%

60%

80%

100%

0

2

4

6

8

10

12

Cum

ulat

ive5

Fre

que

ncy

Fre

quen

cy


(c) FA39

0

2

4

6

8

10

12

14

16

18

Fre

quen

cy

0%

20%

40%

60%

80%

100%

Cum

ulat

ive5

Fre

que

ncy


(d) FB39

0

2

4

6

8

10

12

0%

20%

40%

60%

80%

100%

Cum

ulat

ive5

Fre

que

ncy


Fre

quen

cy

(e) FA46

0

5

10

15

20

25

0%

20%

40%

60%

80%

100%

Cum

ulat

ive3

Fre

que

ncy


Fre

quen

cy

(f) FB46

Figure 5.4: Ultimate capacity per unit length distributions for FRSs with nominal diameters of 35, 39and 46 mm

normal, with very low values of both skewness and kurtosis.

Although the FA39 has the lowest cv, 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. 30 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,


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 0.88, indicating

a peakedness of the distribution. In Figure 5.4e, two outliers are observed: one in the 5-10 kN/m bin,

and one in the 70-75 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 900

pull tests was 1.09 tons/ft, or 31.9 kN/m, with cv = 0.42. 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 50 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.

0

10

20

30

40

50

60

70

0 2 4 6 8 10 12 14

Load

m(kN

)

Displacementm(mm)

SecantmStiffness

FirstmDropmStiffness

Figure 5.5: Stiffness metrics for an FRS


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.

0/

20/

40/

60/

80/

100/

0

1

2

3

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6

10 20 30 40 50 60 70 80 90 100

Cum

ulat

iveS

Fre

que

ncy

Fre

quen

cy

FirstSDropSStiffnessS(kN/mm)0

FrequencyCumulative

(a) FA35 first drop stiffness

7

0/

20/

40/

60/

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100/

0

1

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3

4

5

6

10 20 30 40 50 60 70 80 90 100

Cum

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iveS

Fre

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Fre

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FirstSDropSStiffnessS(kN/mm)0

(b) FA39 first drop stiffness

0%

20%

40%

60%

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0

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10 20 30 40 50 60 70 80 90 100

Cum

ulat

ivek

Fre

que

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Fre

quen

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SecantkStiffnessk(kN/mm)0

(c) FA35 secant stiffness

Fre

quen

cy

0%

20%

40%

60%

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100%

0

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5

6

10 20 30 40 50 60 70 80 90 100

Cum

ulat

ivek

Fre

que

ncy


7

8

9

10

(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 distri-

butions are very large relative to the means of the data (cv > 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


Table 5.3: FRS stiffness summary statistics

Stiffness Bolt n x s cv Skewness KurtosisFirst drop FA35 25 28.8 kN/mm 26.0 kN/mm 0.90 1.32 0.97

FA39 24 25.3 kN/mm 26.0 kN/mm 1.03 1.88 2.72Secant FA35 25 19.7 kN/mm 18.1 kN/mm 0.92 2.66 3.14

FA39 24 9.9 kN/mm 4.7 kN/mm 0.47 0.72 0.00

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

5.3.1 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 (x0) is fully decoupled from the rock mass. Between x0 and x1, the

bolt is partially decoupled with the shear stress along the interface between the bolt and the grout (τb)

equal to the residual strength of the interface (sr). Beyond point x1, the grout is less damaged and has

higher strength, with the peak strength (sp) occurring at point x2. Beyond this point, the full strength of

the grout is not mobilized, and shear stress attenuates down the remainder of the bolt following function

τb(x). Figure 5.8 shows a similar shear stress profile, but shows function P (x), representing tensile load

in the bolt (Li & Stillborg, 1999).


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 x2 (the location of sp, 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 x2 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 = 0, i.e. at the bolt’s threads.

5.3.2 Observed Behaviour of Rebar Rock Bolts

Figure 5.9 shows the load–displacement relationship for a pull test performed on a 20 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.

0

20

40

60

80

100

120

140

160

0 1 2 3 4 5 6 7 8 9 10

Load

(kN

)

Displacement (mm)

A B

Figure 5.9: Pull test performed on a 20 mm Rebar B

The stiffness of a 1.8 m 400W steel bar with a diameter of 20 mm and an elastic modulus of 200


GPa (as is used to manufacture Rebar B; Lamothe, November 2014) 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 115.6

kN (R2 = 0.998) 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 5.10 shows the same relationship for a 20 mm Rebar A, and perhaps a more typical pull test,

where adjustments and rock fracturing are observable on the load–displacement graph.

0

20

40

60

80

100

120

140

160

0 1 2 3 4 5 6 7 8 9 10

Load

(kN

)

Displacement (mm)

Figure 5.10: Pull test performed on a 20 mm Rebar A

A tangent stiffness of 59.3 kN/mm is calculated for the linear portion of the graph, circled in red

(R2 = 0.999). The minimum elastic modulus of Rebar A, like Rebar B, is 200 GPa (Mainville-Beach,

October 2014), 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; Martın et al, 2010), 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 (x2 in Figure 5.7) than by the overall length of

bolt.


5.3.3 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.

0%

20%

40%

60%

80%

100%

0

2

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6

8

10

12

105 110 115 120 125 130 135 140 145C

umul

ativ

e5F

req

uenc

y

Fre

quen

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Working5Capacity5(kN)

FrequencyCumulative

100

(a) 20 mm Rebar A

0

2

4

6

8

10

12

14

Fre

quen

cy

0%

20%

40%

60%

80%

100%

105 110 115 120 125 130 135 140 145

Cum

ulat

ive5

Fre

que

ncy

Working5Capacity5(kN)100

(b) 20 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 133.5 kN), thus

the lack of entries in the 125 kN to 130 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 cv Median Skewness Kurtosis

A 34 123.8 kN 7.5 kN 0.06 124.1 kN 0.12 -0.12B 23 120.9 kN 7.4 kN 0.06 123.1 kN -0.79 1.48

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.


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.

0p

20p

40p

60p

80p

100p

0

2

4

6

8

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12

10 20 30 40 50 60 70 80

Cum

ulat

iveN

Fre

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Fre

quen

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SecantNStiffnessN(kN/mm)

PartiallyNEncapsulated

FullyNEncapsulated

Cumulative

0

(a) Rebar A secant stiffness

0

1

2

3

4

5

6

7

Fre

quen

cy

0%

20%

40%

60%

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100%

10 20 30 40 50 60 70 80

Cum

ulat

iveN

Fre

que

ncy

SecantNStiffnessN(kN/mm)0

(b) Rebar B secant stiffness

Figure 5.12: Secant stiffness for 20 mm rebar

0)

20)

40)

60)

80)

10)

0

1

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7

8

9

Cum

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ives

Fre

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TangentsStiffnessspkN/mmd

PartiallysEncapsulated

FullysEncapsulated

Cumulative

0 10 20 30 40 50 60 70 80

(a) Rebar A tangent stiffness

09

209

409

609

809

1009

0

1

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7

10 20 30 40 50 60 70 80 90

Cum

ulat

ive(

Fre

que

ncy

Fre

quen

cy

Tangent(Stiffness((kN/mm)0

(b) Rebar B tangent stiffness

Figure 5.13: Tangent stiffness for 20 mm Rebar


Table 5.5: Summary statistics for the stiffness of rebar

Stiffness Supplier n x s cv Skewness KurtosisSecant A 43 27.87 kN/mm 13.40 kN/mm 0.48 1.18 3.09

B 21 27.38 kN/mm 14.19 kN/mm 0.52 0.52 0.37Tangent A 36 43.43 kN/mm 25.68 kN/mm 0.59 2.90 11.05

B 21 30.90 kN/mm 10.23 kN/mm 0.33 0.62 0.20

70% of the rebar pull tests in the data base were performed on 1.8 m, 20 mm diameter rebar with

an elastic modulus of 200 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 cvA Full 23 29.2 kN/mm 17.1 kN/mm 0.59

Secant stiffness Partial 27 23.4 kN/mm 9.0 kN/mm 0.39B Full 21 27.4 kN/mm 14.2 kN/mm 0.52

Partial 7 15.3 kN/mm 4.2 kN/mm 0.28A Full 21 50.6 kN/mm 31.2 kN/mm 0.62

Tangent stiffness Partial 22 29.9 kN/mm 9.3 kN/mm 0.31B Full 21 30.9 kN/mm 10.2 kN/mm 0.33

Partial 8 17.6 kN/mm 4.7 kN/mm 0.27

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 0.59 and 0.52). However, Rebar A have greater values of tangent stiffness (an average of 50.6

kN/mm compared to 30.9 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


and pull test apparatus tighten up. 80% 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 10 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.

ym=m0.199xmrm13.34R²m=m0.2884

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120

Sec

antm

Stif

fnes

smuk

N/m

mp

UnloadingmStiffnessmukN/mmp

Fullmencapsulation

Partialmencapsulation

Fullmencapsulationmregression

Partialmencapsulationmregression

All

(a) Secant stiffness

y/=/0.3272x/+/7.2054R²/=/0.9949

y/=/0.5356x/+/1.543R²/=/0.7528

y/=/0.511x/+/1.4277R²/=/0.7503

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120

Tan

gent

/Stif

fnes

s/(k

N/m

m)

Unloading/Stiffness/(kN/mm)

(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

5.4.1 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, 2006). This is illustrated in Figure 5.15.


20

18

16

14

12

10

8

6

4

2

00 50 100 150 200 250

(mm)

(tonnes)

6 mm cone plow

17 mmconeplow

Figure 5.15: Laboratory pull test of an MCB33 (Simser et al, 2006)

Around 50 mm of displacement occurred before strain hardening is induced. Only 17 mm of this was

attributed to cone plough. Of the remaining 170 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 5.16.

Eventual failureof the bolt

Limited plow of thecone (~ 20-40 mm)

Steel stretching

Load

Deformation

Figure 5.16: Conceptual load–displacement behaviour of a cone bolt subject to quasi-static loading(Simser et al, 2006)

5.4.2 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 30-40 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


magnitude occurring between yield and strain hardening may be observed for a carbon steel, depending

on composition and manufacturing process (ASM International, 2002). It should be noted that testing

on the bolt shown in Figure 5.17 was stopped before failure occurred.

0

2

0

0

20

40

60

80

100

120

140

160

0 10 20 30 40 50 60 70 80

Lo

ad

(kN

)

Displacement (mm)

A B C D

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.

0

20

40

60

80

100

120

140

160

0 10 20 30 40 50 60 70 80

Load

(kN

)

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 (200 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 (R2 = 0.9992). 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


becoming more damaged at an accelerating rate, but is competent enough to resist a linear plouging

response until the bolt tendon yields.

R2t=t0.9992

0

20

40

60

80

100

120

140

160

0 1 2 3 4 5 6 7 8

Load

(kN

)

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 signif-

icant 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).

5.4.3 Proposed Interpretation of Modified Cone Bolt Behaviour

Figure 5.20 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, 2009; Doucet & Voyzelle, 2012) 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


Load

Displacement

Elastic deformation of bolt tendon

Plough begins

Elastic deformation and plough

Yield of bolt tendon

Plastic deformation of tendon, limited plough

Strain hardeningTo failure of bolt

Figure 5.20: 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.

0

20

40

60

80

100

120

140

160

0 10 20 30 40 50 60 70 80

Load

(kN

)

Displacement (mm)

PloughkStiffness

SteelkYieldkStrength

PloughkPoint

InitialkStiffness

SecantkStiffness

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


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.

5.4.4 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.

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Figure 5.22: Load metric distributions for the MCB33

Table 5.7: Summary statistics of MCB33 load metrics

Variable n x s cv Skewness KurtosisPlough Point 63 53.2 kN 22.7 kN 0.43 0.48 −0.45Yield Strength 52 117.2 kN 16.3 kN 0.14 −1.51 2.55

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 2006 which contributed 5 of the 7 values in the 80 - 90 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 20 and 30 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 5.23. Summary

statistics are presented in Table 5.8.


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(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 cv Skewness KurtosisInitial Stiffness 51 10.4 kN/mm 4.1 kN/mm 0.40 1.63 4.64Plough Stiffness 50 2.23 kN/mm 0.90 kN/mm 0.41 0.59 −0.37Secant Stiffness 58 4.95 kN/mm 3.51 kN/mm 0.71 1.49 1.90

The stiffness of a 200 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


Hadjigeorgiou & Potvin, 2011), or using passive monitoring in situ (as in Morissette et al, 2014). The

MCB33 has been demonstrated to perform significantly differently when subject to impact loading versus

quasi-static (Doucet & Voyzell, 2012), as such the results of one loading rate may not indicate a similar

effect for another.

5.5 D-Bolts

5.5.1 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, 2012). 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, 2012)

Figure 5.25: Results of simulated joint laboratory tests on 20 mm D-Bolts (Li, 2012)

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, 2012. 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


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 0.12 m long anchor sections, at 0.34

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, 0.34 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 0.9 m in length; Li, 2012). 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 20 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 200-210 GPa (Charette, 2014).

Table 5.9: D-Bolt stiffness by anchor stability

Stable Anchor Length 20 mm D-Bolt 22 mm D-Bolt1st 0.34 m 184 kN/mm 227 kN/mm2nd 1.14 m 55 kN/mm 68 kN/mm3rd 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.

5.5.2 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).


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Creighton,C2014

CopperCCliff,C2009

BeforeCfirstCanchorC(0.34Cm)

BeforeCthirdCanchorC(2.24Cm)

BeforeCsecondCanchorC(1.14Cm)

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 (2012; Figure 5.25) exhibited slightly more than 10 mm of perfectly plastic deformation between

yield and strain hardening on a smooth section of 0.8 m, corresponding to a strain of around 1.3%. The

bolt pulled at Creighton in 2014 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 (2012), 2.5 mm of deformation representing a strain of 1.3%

corresponds to a deformation length of about 190 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 3.0 thread (Normet, 2014). 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 3.0 thread is 20.7 mm for flat form threads (ASME, 2005). Assuming the threading process

does not significantly alter the yield strength of the bolt material, the lowest diameter part of the bolt

(the 20.7 mm minor thread diameter) will yield first, about 20 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 5.25. 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 20 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 30◦ off of perpendicular from the wall face. It is not clear how this


affects the results, although the appearance of a working capacity of about 120 kN was attributed to

movement of the testing rig (the minimum yield load of a 20 mm D-Bolt is 140 kN).

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Boltl1

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Beforelfirstlanchorl(0.34lm)

Beforelthirdlanchorl(2.24lm)

Beforelsecondlanchorl(1.14lm)

Figure 5.27: Pull tests performed on 20 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 20 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 near-

identical 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 5.28. 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

5.10 shows the unloading stiffness calculated for the three pull tests.

Table 5.10: Unloading stiffness calculated for pull tests performed on 22 mm D-Bolts

Test Stiffness1 635 kN/mm2 54 kN/mm3 36 kN/mm


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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.

5.5.3 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 5.11.

Table 5.11: Working capacities obtained from all D-Bolt pull tests

Diameter Working Capacity142 kN

20 mm 160 kN142 kN196 kN

22 mm 196 kN196 kN


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.

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Figure 5.29: Distributions of stiffness for 20 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 20 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 cv Skewness KurtosisTangent 20 mm 12 31.2 kN/mm 22.3 kN/mm 0.72 1.18 1.34

22 mm 15 19.2 kN/mm 8.3 kN/mm 0.43 -0.10 -1.57Secant 20 mm 12 24.6 kN/mm 18.1 kN/mm 0.73 1.20 1.58

22 mm 19 13.4 kN/mm 8.8 kN/mm 0.65 0.53 -1.35

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 20 mm D-Bolts as generally

acting stiffer than the 22 mm equivalent. This is likely misleading; two thirds of the 20 mm D-Bolt pull

tests were performed at Copper Cliff Mine, on 3550 and 3710 Levels (1,080 and 1130 m), while over half

of the 22 mm D-Bolt pull tests were performed at Creighton Mine at twice that depth, between 7680

and 7940 Levels (2340 and 2420 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.


5.6 Expandable Bolts

5.6.1 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.

5.6.2 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 5.30 shows examples of three, with ideal elastic

stiffness shown for bolts of equivalent length, cross–sectional area and elastic modulus (200 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, 2005).

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1.8WmWStiffness

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WorkingWCapacity

Figure 5.30: 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


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 .3048 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 VariablePm12 19 42% 32% 5% 21%Mn12 8 12.5% 25% 25% 37.5%Pm24 15 13% 27% 12% 47%Mn24 9 22% 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 AveragePartial Rock or ore 11 131 kN/m 263 kN/m 228 kN/m

Full Sandfill or paste 12 14 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 40

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 (2000) 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


the database assembled for this thesis are known to have slipped. Soni’s database included 173 “slipped”

tests out of 304 entries in the data set (Soni, 2000). 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 (2000). 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.

5.6.3 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 5.15. 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.

0P

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Pm12

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0 20 40 60 80 100 120 140 160 180

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 cvPm12 15 88.4 kN 5.5 kN 0.06Mn12 5 96.1 kN 7.4 kN 0.08

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


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.

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Secant3Stiffness3(kN/mm)

Mn12Pm12Cumulative

0

(a) Pm12, Mn12

0

1

2

3

4

5

6

7

Fre

quen

cy

Mn24Pm24Cumulative

0%

20%

40%

60%

80%

100%

4 8 12 16 20 24 28 32

Cum

ulat

iveN

Fre

que

ncy

SecantNStiffnessN(kN/mm)0

(b) Pm24, Mn24

Figure 5.32: Secant stiffness of Swellex variants

Table 5.16: Swellex secant stiffness summary statistics

Variant n x s cv Skewness Kurtosis Cross-sectional areaPm12 37 7.60 kN/mm 2.71 kN/mm 0.36 0.14 -0.67 2.45 ∗ 10−4mm2

Mn12 10 6.54 kN/mm 3.31 kN/mm 0.51 0.77 -0.36 2.45 ∗ 10−4mm2

Pm24 24 12.40 kN/mm 7.22 kN/mm 0.58 0.76 0.15 4.81 ∗ 10−4mm2

Mn24 16 12.37 kN/mm 7.17 kN/mm 0.58 0.80 -0.02 4.81 ∗ 10−4mm2

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 20.1 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 200

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


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.

2

4

6

8

10

12

14

0

5

10

15

20

25

30

35

Sec

ant S

tiffn

ess

(kN

/mm

)

Pm12Mn12

Sandfill Rock/Ore

(a) Pm12 and Mn12

0

5

10

15

20

25

30

35

Pm24Mn24

0

5

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15

20

25

30

35

Sec

ant S

tiffn

ess

(kN

/mm

)

Sandfill Rock/Ore

(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 cv pPm12 & Mn12 Sandfill 34 7.3 kN/mm 2.5 kN/mm 0.34

Rock/ore 10 6.5 kN/mm 3.3 kN/mm 0.51 0.524Pm24 & Mn24 Sandfill 19 10.1 kN/mm 5.7 kN/mm 0.56

Rock/ore 14 16.2 kN/mm 8.7 kN/mm 0.53 0.027

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

0.524, 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=0.027). 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 2011 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.


5.7.1 Yield-Lok

Two pull test campaigns were performed on Yield-Lok bolts by Vale’s Sudbury operations between 2011

and 2014, 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 5.18.

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, 2011 151 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, 2013 N/A 36.6 kN/mm

N/A 18.1 kN/mm

N/A 16.0 kN/mm

Although only three working capacities were recorded, they were all around 150 kN. The 2013 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.

5.7.2 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 345.6 kN and 371.1 kN, averaging 358.2 kN, corresponding to an average

ultimate tensile strength of 1035 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, 2001). 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 200 GPa, the laboratory

testing of the FRP rebar resulted in an average elastic modulus of 57.7 GPa. For a 71” (1.80 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 20.1 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


19 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.

5.7.3 DS Bolt

One testing campaign was performed on 4 DS Bolts (now known as the VersaBolt; Lamothe, September

2014). Comparable to the D-Bolt, it is a smooth bar punctuated by oval anchors along its length. 20.5

mm bolts were pulled at Totten on April 19th, 2013. Load and displacement data was recorded for 3

of these tests. Working capacity and tangent and secant stiffness are shown in Table 5.19. 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 Stiffness2 124.3 kN 22.0 kN/mm 19.5 kN/mm3 116.4 kN 36.9 kN/mm 24.5 kN/mm4 125.7 kN 22.8 kN/mm 18.3 kN/mm

The typical thread and bar yield strengths of the 20.5 mm DS Bolt are 125 kN and 138 kN respectively

(courtesy of Mansour), indicating yield of the thread.

5.7.4 Other Expandable Bolts

Although Swellex were by far the most tested expandable bolts, three other configurations were tested.

One campaign of Jennmar’s Midi (160 kN ultimate capacity) Python bolt was performed at Creighton

in 2010, 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 2013. The results of the pull tests are shown in Table 5.20.

Table 5.20: Summary of expandable bolt campaigns not including Swellex

Bolt Configuration n Embedment Working Capacity Secant StiffnessPython Midi 5 N/A N/A N/AOmega 24t 1 Full 178 kN 10.2 kN/mmOmega 12t 4 Partial N/A N/A

Table 5.20 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.

5.7.5 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.


5.8 Summary

Performance metrics have different degrees of variability for the reinforcement elements discussed. Met-

rics 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 40 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 equiva-

lent 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.


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 D4435-13. 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 cv

FA35 81 40.0 kN/m 10.3 kN/m 0.26

FA39 30 38.3 kN/m 7.2 kN/m 0.19

FA46 50 39.1 kN/m 11.1 kN/m 0.28

FRS* FB35 92 40.0 kN/m 13.7 kN/m 0.34

FB39 83 37.9 kN/m 13.3 kN/m 0.35

FB46 106 37.8 kN/m 10.8 kN/m 0.29

All 442 38.9 kN/m 11.7 kN/m 0.30

Grouted (Static) Rebar A (20 mm) 34 123.8 kN 7.5 kN 0.06

Rebar B (20 mm) 23 120.9 kN 7.4 kN 0.06

MCB33* 52 117.2 kN 16.3 kN 0.14

D-Bolt (20 mm) 3 148 kN N/A N/A

Grouted (Yielding) D-Bolt (22 mm) 3 196 kN N/A N/A

Versabolt 3 122 kN N/A N/A

Yield-Lok 3 149 kN N/A N/A

Expandable Swellex Pm12 15 88.4 kN 5.5 kN 0.06

Swellex Mn12 5 96.1 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.

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.


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 param-

eters 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.

6.1.1 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 6.1.

78

Chapter 6. Factors Influencing Pull Test Performance 79

0

20

40

60

80

100

120

0.0 0.5 1.0 1.5 2.0 2.5

Ulti

mat

e C

apa

city

(kN

)

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 mn 10 71x 49.82 kN 61.68 kNs 13.17 kN 15.82 kNcv 0.26 0.26

t0 −2.595p 0.0111

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 = 0.05). As the p-value is 0.0111, 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 (10 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 0.15 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 = 0.05) 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).


0

20

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60

80

100

120

0.0 0.5 1.0 1.5 2.0 2.5

Ulti

mat

e C

apa

city

(kN

)


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 mn 68 20 4x 63.4 kN 69.4 kN 40.0 kNs 22.6 kN 10.2 kN 17.4 kNcv 0.36 0.15 0.43t0 1.681p 0.049

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.

0

20

40

60

80

100

120

0.0 0.5 1.0 1.5 2.0 2.5

Ulti

mat

e C

apa

city

(kN

)


Figure 6.3: Relationship between ultimate capacity and length for the FB39


Table 6.3: Statistics on the relationship between ultimate capacity and length for the FB39

Length 1.52 m 1.65 m 2.13 mn 64 10 11x 61.8 kN 56.9 kN 46.4 kNs 20.1 kN 10.4 kN 20.5 kNcv 0.33 0.18 0.44

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 lengthfor the FB39

Source of Variation SS df MS F p-value F critBetween Groups 4614.4 2 2307.2 6.1628 0.0032 3.1079Within Groups 30699 82 374.37

Total 35313 84

A p-value of 0.0032 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 5.2.3 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 mFA35 30.5% 0%FA39 12.9% 0%FA46 0% 31%FB35 29.2% 12.7%FB39 27.5% 0%FB46 0% 56.3%


y/=/14.947x/h/39.486R²/=/0.0207

0a 10a 20a 30a 40a 50a

0

20

40

60

80

100

120

140

160

0.0 0.5 1.0 1.5 2.0 2.5

Distribution/p10/kN/m/binsc

Ulti

mat

e/C

apac

ity/p

kNc

Anchorage/Length/pmc

L=1.52m

L=1.83m

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 = 0.05).

Table 6.6: Comparison of 1.52 m and 1.83 m of anchorage length for all FRS bolts

Anchorage Length n x s cv t tcrit p1.52 m 233 61.8 kN 18.7 kN 0.301.83 m 158 70.3 kN 18.6 kN 0.26 -4.413 1.649 0.0000

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 475 39 mm Split Sets, over a much wider range of

installation conditions (over 50 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.


6.1.2 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 = 0.05).

0

10

20

30

40

50

60

70

Ulti

mat

edC

apa

city

d(kN

/m)

Bolter-installed

Copper Cliff 04/27/2012

Copper Cliff 05/11/2012

Totten 04/22/2010

Stobie 05/28/2014

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 Boltern 31 44x 32.60 kN/m 46.25 kN/ms 8.49 kN/m 7.47 kN/mcv 0.26 0.16t0 −7.2tcrit 2.001p 6 ∗ 10−10

There is a significant difference between the means of the two data sets, with bolter-installed FRSs

exhibiting almost 50% more capacity than a jackleg-installed bolt. The jackleg bolts represent four

testing campaigns in different conditions, and 90% 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.


0%

5%

10%

15%

20%

25%

30%

5 10 15 20 25 30 35 40 45 50 55 60 65 70 75

Frequency

Ultimate Capacity (kN/m)

MacLean

Jackleg

0

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 Boltern 39 289x 31.8 kN/m 39.8 kN/ms 7.83 kN/m 11.59 kN/mcv 0.25 0.29t0 5.56tcrit 1.67p 0.0000

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


the modern methods of bolt installation used at Vale’s Sudbury operations.

6.1.3 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, 2008), 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+I38.407R²I=I0.5252

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

eIC

apac

ityI(

kN)

InstallationITimeI(s)

(a) FA35

yn=n0.4436xn+n56.484R²n=n0.03

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

enC

apac

ityn(

kN)

InstallationnTimen(s)

(b) FB35

yI=I1.119xI+I46.415R²I=I0.1513

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

eIC

apac

ityI(

kN)


(c) FA39

yI=I2.4483xI+I25.055R²I=I0.3946

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

eIC

apac

ityI(

kN)


(d) FB39

yI=I1.5842xI+I40.991R²I=I0.5517

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

eIC

apac

ityI(

kN)


(e) FA46

yn=n1.2072xn+n48.476R²n=n0.2298

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

enC

apac

ityn(

kN)


(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


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+n44.69R²n=n0.4931

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

enC

apac

itynu

kNw

InstallationnTimenusw

Bolter

Jacklegnornunknown

Figure 6.8: Relationship between drive time and ultimate capacity for FB35, distinguishing betweeninstallation methods

Separating the MacLean-installed FB35 bolts results in a clearer relationship (R2 = 0.493, compared

to 0.030 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 installedwith a MacLean Bolter

Variant n Average Ultimate Capacity Average Drive Time Regression Coefficient R2

FA35 36 73.6 kN 19.5 s 0.86 kN/s 0.70FA39 23 60.5 kN 11.7 s 0.98 kN/s 0.11FA46 38 65.7 kN 16.7 s 1.53 kN/s 0.41FB35 58 71.6 kN 11.9 s 2.25 kN/s 0.49FB39 34 65.8 kN 15.0 s 3.05 kN/s 0.55FB46 19 68.5 kN 18.1 s 1.80 kN/s 0.54


yI=I0.8646xI+I56.744R²I=I0.6984

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

eIC

apac

ityI(

kN)


(a) FA35

yn=n2.2248xn+n44.69R²n=n0.4931

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

enC

apac

ityn(

kN)


(b) FB35

yI=I1.119xI+I46.415R²I=I0.1513

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

eIC

apac

ityI(

kN)


(c) FA39

yI=I3.052xI+I19.913R²I=I0.5455

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

eIC

apac

ityI(

kN)


(d) FB39

yI=I1.5291xI+I40.25R²I=I0.4127

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

eIC

apac

ityI(

kN)


(e) FA46

y)=)1.7974x)+)36.01R²)=)0.5383

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

e)C

apac

ity)(

kN)

Installation)Time)(s)

(f) FB46

Figure 6.9: Relationship between drive time and absolute ultimate capacity for all FRS configurationsinstalled 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 0.86

kN/s for the FA35 to 3.05 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 10 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

6.10 shows a power function fit to the data from all FRS configurations.


y = 23.661x0.3904

R² = 0.4217

0

20

40

60

80

100

120

140

160

0 20 40 60 80 100

Ulti

mat

enC

apa

city

n+kN

)

InstallationnTimen+s)

FB35

FB39

FB46

FA35

FA39

FA46

y = 1.4801x + 45.695R² = 0.3641

Figure 6.10: Relationship between installation time and ultimate capacity for all bolter-installed FRSfit with linear and power functions

Figure 6.10 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 6.10, possibly due to the fact

that Tomory et al. were primarily investigating bolts installed using a jackleg, while the data presented

in Figure 6.10 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 10 seconds, it may be

prudent to implement a denser bolting pattern. Conversely, a 30 second drive time for any bolt suggests

the bolt will perform well in a loading scenario analogous to a pull test.

6.1.4 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


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=D0.093

0

20

40

60

80

29.0 31.0 33.0 35.0 37.0

Ulti

mat

eDC

apac

ityD(

kN/m

)

DrillDBitDDiameterD(mm)

(a) FA35

R²D=D0.0053

0

20

40

60

80

29.0 31.0 33.0 35.0 37.0

Ulti

mat

eDC

apac

ityD(

kN/m

)


(b) FB35

R²D=D0.1082

0

20

40

60

80

31.0 33.0 35.0 37.0 39.0

Ulti

mat

eDC

apac

ityD(

kN/m

)


(c) FA39

R²D=D0.0009

0

20

40

60

80

31.0 33.0 35.0 37.0 39.0

Ulti

mat

eDC

apac

ityD(

kN/m

)


(d) FB39

R²D=D0.0809

0

20

40

60

80

39.0 41.0 43.0 45.0 47.0

Ulti

mat

eDC

apac

ityD(

kN/m

)


(e) FA46

R²D=D0.3651

0

20

40

60

80

39.0 41.0 43.0 45.0 47.0

Ulti

mat

eDC

apac

ityD(

kN/m

)


(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.


R²D=D0.8044

R²D=D0.2482

0

10

20

30

40

50

60

70

80

39.0 40.0 41.0 42.0 43.0 44.0 45.0 46.0 47.0

Ulti

mat

eDC

apa

city

D(kN

/m)


CreightonD9/12/2005

CreightonD4/28/2008

Figure 6.12: Relationship between drill bit diameter and ultimate capacity for two testing campaignsperformed 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=D0.0342

R²D=D0.4411

0

20

40

60

80

29.0 31.0 33.0 35.0 37.0

Ulti

mat

eDC

apa

city

DgkN

/m)

DrillDBitDDiameterDgmm)

MacLeanDBolterJackleg

(a) FA35

R²D=D0.2354

R²D=D0.0169

0

20

40

60

80

29.0 31.0 33.0 35.0 37.0

Ulti

mat

eDC

apa

city

D(kN

/m)


MacLeanDBolterUnknown

(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 (R2 = 0.4411). 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

(R2 = 0.2354). 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


a universally applicable model of FRS performance dependant solely on the diameter of the drill bit

used.

6.1.5 Influence of Bolt Diameter

There is insufficient variation in bolt diameter for specific FRS sizes for its effects to be reliably deter-

mined. 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 6.10. The average ultimate

capacities in Table 6.10 are calculated from the same data set used for the calculation of the bolt to drill

bit diameter ratios.

0.98

1.00

1.02

1.04

1.06

1.08

1.10

1.12

1.14

1.16

1.18

Bol

t D

iam

eter

/Bit

Dia

met

er

FA35 FB35 FA39 FB39 FA46 FB46Bolt Configuration

Figure 6.14: Bolt diameter to drill bit diameter ratios for all FRS variants

Table 6.10: Summary statistics for bolt diameter to drill bit diameter ratios for all FRS variants

Variant n x s cv Average Ultimate CapacityFA35 58 1.109 0.024 0.022 41.4 kN/mFB35 26 1.068 0.012 0.011 41.6 kN/mFA39 25 1.100 0.025 0.023 39.0 kN/mFB39 28 1.061 0.023 0.021 35.5 kN/mFA46 43 1.061 0.039 0.037 41.0 kN/mFB46 48 1.057 0.036 0.034 34.1 kN/m

Figure 6.14 and Table 6.10 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 6.1.4 this does not necessarily

translate directly into performance. Table 6.10 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


relationship between bolt to drill bit diameter ratio and ultimate capacity.

R²o=o0.0945

0

20

40

60

80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

eoC

apa

city

o(kN

/m)

Bolto/oDrilloBitoDiameter

(a) FA35

R²D=D0.0306

0

20

40

60

80

0.98 1.03 1.08 1.13

Ulti

mat

eDC

apa

city

D(kN

/m)

BoltD/DDrillDBitDDiameter

(b) FB35

R²D=D0.0019

0

20

40

60

80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

eDC

apa

city

D(kN

/m)


(c) FA39

R²D=D0.0091

0

20

40

60

80

0.98 1.03 1.08 1.13

Ulti

mat

eDC

apa

city

D(kN

/m)


(d) FB39

R²o=o0.0253

0

20

40

60

80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

eoC

apa

city

o(kN

/m)

Bolto/oDrilloBitoDiameter

(e) FA46

R²D=D0.2289

0

20

40

60

80

0.98 1.03 1.08 1.13

Ulti

mat

eDC

apa

city

D(kN

/m)


(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 10 or more data points.


R²B=B0.4567

0

20

40

60

80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

eBC

apa

city

BhkN

/mO

BoltB/BDrillBBitBDiameter

Totten,B04/22/2010Totten,B4/4/2014Creighton,B4/29/2014BhOreOCreighton,B4/29/2014BhNoriteOTotten,B9/23/2010

(a) FA35

0

20

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80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

enC

apa

city

n(kN

/m)

Boltn/nDrillnBitnDiameter

Totten,n2014/11/19

(b) FB35

0

20

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60

80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

esC

apa

city

s(kN

/m)

Bolts/sDrillsBitsDiameter

Garson,s9/30/2013

Coleman,s4/2/2014s(Ore)

Coleman,s4/2/2014s(Granite)

(c) FA39

0

20

40

60

80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

enC

apa

city

n(kN

/m)

BoltnDiametern:nDrillnBitnDiameter

Coleman,n3/7/2014

(d) FB39

0

20

40

60

80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

egC

apa

city

gXkN

/m)

BoltgDiameterg:gDrillgBitgDiameter

Creighton,g6/10/2014gXOre)

Creighton,g6/10/2014gXSubX)

(e) FA46

0

20

40

60

80

0.98 1.03 1.08 1.13 1.18

Ulti

mat

efC

apa

city

f(kN

/m)

BoltfDiameterf:fDrillfBitfDiameter

CopperfCliff,f2014/1/31f(QuartzfDiorite)CopperfCliff,f2014/1/31f(Ore)Totten,f2014/12/3

(f) FB46

Figure 6.16: Relationship between bolt diameter to drill bit diameter ratio and ultimate capacity forall 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, 4/22/2010) 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.

6.1.6 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), brec-

cias (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 cate-

gories. 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.


0

20

40

60

80

Ulti

mat

eUC

apa

city

U(kN

/m)

Metased

MetaigneousOreIgneous

BrecciaSandfill

LargeUdrillbitUdiameterU(relative)U

SmallUdrillbitUdiameterU(relative)

(a) FA35

0

20

40

60

80

Ulti

mat

ekC

apa

city

k(kN

/m)

Metased


BrecciaSandfill

(b) FB35

0

20

40

60

80

Metased


BrecciaSandfill

Ulti

mat

ekC

apa

city

k(kN

/m)

(c) FA39

0

20

40

60

80

Ulti

mat

ekC

apa

city

k(kN

/m)

Metased


BrecciaSandfill

(d) FB39

0

20

40

60

80

Ulti

mat

ekC

apa

city

k(kN

/m)

Metased


BrecciaSandfill

(e) FA46

0

20

40

60

80U

ltim

atek

Ca

paci

tyk(

kN/m

)

Metased


BrecciaSandfill

(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 under-

perform 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.


R²P=P0.3039

0

20

40

60

80

0 100 200 300

Ulti

mat

ePC

apac

ityP(

kN/m

)

UCSP(MPa)

(a) FA35

R² = 0.0201

0

20

40

60

80

0 100 200 300

Ulti

mat

e C

apac

ity (

kN/m

)

UCS (MPa)

(b) FB35

R²P=P0.0356

0

20

40

60

80

0 100 200 300

Ulti

mat

ePC

apac

ityP(

kN/m

)

UCSP(MPa)

(c) FA39

R² = 0.0262

0

20

40

60

80

0 100 200 300

Ulti

mat

e C

apac

ity (

kN/m

)

UCS (MPa)

(d) FB39

R²P=P0.0027

0

20

40

60

80

0 100 200 300

Ulti

mat

ePC

apac

ityP(

kN/m

)

UCSP(MPa)

(e) FA46

R² = 0.0611

0

20

40

60

80

0 100 200 300

Ulti

mat

e C

apac

ity (

kN/m

)

UCS (MPa)

(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 inthe same campaign

Campaign UCSore Element nOre nWaste xOre xWaste

FA35 3 3 36.4 kN/m 27.2 kN/mFA39 3 3 34.2 kN/m 34.6 kN/m

Creighton, 5/16/2011 130 MPa FA46 4 5 44.1 kN/m 48.9 kN/mFB35 3 3 31.2 kN/m 30.2 kN/mFB46 3 3 42.5 kN/m 32.8 kN/m

Garson, 11/24/2011 91 MPa FB39 5 5 26.5 kN/m 27.9 kN/mStobie, 10/15/2012 155 MPa FB39 5 5 34.3 kN/m 31.2 kN/mCopper Cliff, 1/31/2014 110-150 MPa FB46 4 5 30.2 kN/m 28.3 kN/mColeman, 4/2/2014 130 MPa FA39 5 5 37.4 kN/m 36.9 kN/mCreighton, 4/29/2014 130 MPa FA35 5 5 43.8 kN/m 43.9 kN/mCreighton, 6/10/2014 130 MPa FA46 6 4 30.9 kN/m 29.9 kN/mCopper Cliff, 3/3/2015 110-150 MPa FB35 5 5 44.9 kN/m 53.12 kN/m

FB39 5 5 54.9 kN/m 58.4 kN/m


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 tcrit and t values are two-tailed (α = 0.05).

0% 10% 20% 30% 40% 50% 60%

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

1.6

1.8

Distribution

NormalizedUltimateCapacity

Ore Waste

Figure 6.19: Distributions of FRS bolts installed in ore and waste rock normalized to the campaignaverage 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 cv t tcrit pOre 56 0.995 0.206 0.207

Waste 56 1.000 0.124 0.124 0.167 1.987 0.868Ore 56 1.000 0.181 0.181

Waste 56 1.021 0.173 0.169 0.618 1.982 0.538

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 170 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.


6.1.7 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 80, although may be as low as 45 and as high as 90 in some lithologies. Bieniawski

(1989) distinguishes between “Good” and “Fair” rock at an RMR of 60; as such, it is assumed that

testing performed in rock masses labelled as Poor have an RMR that falls below 60, and Good above

60.

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 6.20 and Table 6.13. The t-test

performed is two-tailed (α = 0.05) assuming unequal variances.

0

10

20

30

40

50

60

70

80

UltimateCapacity(kN/m)

Poor GoodRock Mass Quality

Figure 6.20: 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 cv t tcrit pGood 26 39.3 kN/m 12.3 kN/m 0.31Poor 9 30.6 kN/m 6.3 kN/m 0.21 2.704 2.048 0.012

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

30% 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


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.

0, 10, 20, 30, 40, 50, 60,

0

10

20

30

40

50

60

70

80

90

Distribution R5 kN/m binsM

Ulti

mat

eCC

apa

city

CRkN

/mM Garson, 9/20/2013

PoorGood

Poor GoodRockCMassCQuality:

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, 9/20/2013, 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 6.22. 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, 9/20/2013


Table 6.14: Comparison of all FRSs installed in poor and good quality ground

Ground n x s cv t tcrit pGood 80 39.2 kN/m 9.5 kN/m 0.24

Poor (including Garson, 9/20/2013) 25 34.9 kN/m 9.8 kN/m 0.28 1.926 2.023 0.061Poor (omitting Garson, 9/20/2013) 19 30.8 kN/m 5.4 kN/m 0.18 5.157 2.011 0.000

If the Garson FA39 test is included, the p-value of 0.061 is calculated, resulting in a marginal t-test

result; strictly speaking, the null hypothesis is not rejected if α = 0.05, 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 30% 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.

6.1.8 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


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.


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 12” – 24” 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.

6.2.1 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.


0

50

100

150

200

0.0 1.0 2.0 3.0

Tan

gen

tbStif

fnes

sb(k

N/m

m)

BoltbLengthb(m)

FullbEmbedment

PartialbEmbedment


0

50

100

150

200

0.0 1.0 2.0 3.0

Tan

gen

t Stif

fnes

s (k

N/m

m)

Bolt Length (m)


0

50

100

150

200

0.0 1.0 2.0 3.0

Sec

ant S

tiffn

ess

(kN

/mm

)

Bolt Length (m)

(c) Rebar A secant stiffness

0

50

100

150

200

0.0 1.0 2.0 3.0

Sec

ant S

tiffn

ess

(kN

/mm

)

Bolt Length (m)

(d) Rebar B secant stiffness

0

40

80

120

160

0.0 1.0 2.0 3.0

Wor

king

Cap

acity

(kN

)

Bolt Length (m)

(e) Rebar A working capacity

0

40

80

120

160

0.0 1.0 2.0 3.0

Wor

king

Cap

acity

(kN

)

Bolt Length (m)

(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.

6.2.2 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.


0

10

20

30

40

50

60

0.0 0.2 0.4 0.6 0.8 1.0

Tan

gen

thStif

fnes

sh(k

N/m

m)

GrouthLengthh:hBolthLength


0

10

20

30

40

50

60

0.0 0.2 0.4 0.6 0.8 1.0

Tan

gen

thStif

fnes

sh(k

N/m

m)



0

10

20

30

40

50

60

0.0 0.2 0.4 0.6 0.8 1.0

Sec

anthS

tiffn

essh

(kN

/mm

)


(c) Rebar A secant stiffness

0

10

20

30

40

50

60

0.0 0.2 0.4 0.6 0.8 1.0

Sec

anthS

tiffn

essh

(kN

/mm

)



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 cvTangent Partial 22 29.9 kN/mm 9.3 kN/mm 0.31

A Full 10 40.6 kN/mm 13.7 kN/mm 0.34Secant Partial 27 23.4 kN/mm 9.0 kN/mm 0.31

Full 9 34.0 kN/mm 13.7 kN/mm 0.33Tangent Partial 8 17.6 kN/mm 4.7 kN/mm 0.27

B Full 12 25.4 kN/mm 5.2 kN/mm 0.20Secant Partial 8 15.3 kN/mm 4.2 kN/mm 0.28

Full 16 21.7 kN/mm 9.6 kN/mm 0.44

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 70% 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.


0

20

40

60

80

100

120

140

160

180

200

0% 20% 40% 60% 80% 100%

UnloadingStiffness(kN/mm)

Grout coverage1.8 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, corre-

sponding 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 0% grout coverage, the maximum stiffness is equivalent to the stiffness of the entire length

of rebar, and at 100% 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

20 mm bolt (314 mm2) plus a 28 mm resin cartridge (616 mm2; combined total of 930 mm2) is roughly

equivalent to a 34 mm hole (908 mm2) 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.

6.2.3 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.


0

10

20

30

40

50

60

70

0 2 4 6 8 10 12SpinETimeE(s)

1.8EmEFullEEncapsulation1.8EmEPartialEEncapsulation

Tan

gen

tEStif

fnes

sE(k

N/m

m)

(a) Tangent stiffness

0

10

20

30

40

50

60

70

0 2 4 6 8 10 12

Sec

ant S

tiffn

ess

(kN

/mm

)

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 10 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.

6.2.4 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.

0

25

50

75

100

125

150

175

200

Tan

gen

tlStif

fnes

sl(k

N/m

m)l

1.8lm2.4lm

InstallationDaylofltest Previous

(a) Tangent stiffness

0

25

50

75

100

125

150

175

200

Sec

ant S

tiffn

ess

(kN

/mm

)

InstallationDay of test Previous

(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 cvDay of installation 2.4 m 5 77.0 kN/mm 49.9 kN/mm 0.65

Tangent 2.4 m & 1.8 m 15 52.8 kN/mm 33.9 kN/mm 0.64Previously installed 2.4 m 6 45.3 kN/mm 25.2 kN/mm 0.56Day of installation 2.4 m 5 38.8 kN/mm 24.0 kN/mm 0.62

Secant 2.4 m & 1.8 m 14 35.7 kN/mm 17.3 kN/mm 0.49Previously installed 2.4 m 9 19.0 kN/mm 11.2 kN/mm 0.59


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.

6.2.5 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 6.17.

1.8 m Full1.8 m Partial2.4 m Full

0

20

40

60

80

100

120

140

160

Tan

gen

t S

tiffn

ess

(kN

/mm

)

Ore Waste


0

20

40

60

80

100

120

140

160

Sec

ant S

tiffn

ess

(kN

/mm

)

Ore Waste

(b) Rebar A secant stiffness

0

20

40

60

80

100

120

140

160

Tan

gen

t Stif

fnes

s (k

N/m

m)

Ore Waste

(c) Rebar B tangent stiffness

0

20

40

60

80

100

120

140

160

Sec

ant S

tiffn

ess

(kN

/mm

)

Ore Waste


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.


Table 6.17: Comparison of stiffness across different lithologies

Manufacturer Stiffness Configuration Rock Unit n x s cv1.8 m Full Ore 0

Host 10 40.6 kN/mm 13.7 kN/mm 0.34Tangent 1.8 m Partial Ore 15 33.4 kN/mm 8.5 kN/mm 0.25

Host 7 22.4 kN/mm 6.5 kN/mm 0.292.4 m Full Ore 7 68.0 kN/mm 44.4 kN/mm 0.65

A Host 4 45.2 kN/m 30.1 kN/mm 0.671.8 m Full Ore 0

Host 9 34.0 kN/mm 13.7 kN/mm 0.40Secant 1.8 m Partial Ore 20 26.3 kN/mm 7.4 kN/mm 0.28

Host 7 15.0 kN/mm 8.4 kN/mm 0.562.4 m Full Ore 8 36.6 kN/mm 18.5 kN/mm 0.51

Host 6 12.1 kN/mm 4.9 kN/mm 0.41Tangent 1.8 m Full Ore 5 20.9 kN/mm 3.7 kN/mm 0.17

Host 7 28.6 kN/mm 3.5 kN/mm 0.121.8 m Partial Ore 3 15.7 kN/mm 3.5 kN/mm 0.22

B Host 5 18.8 kN/mm 5.3 kN/mm 0.28Secant 1.8 m Full Ore 7 22.3 kN/mm 11.9 kN/mm 0.53

Host 9 21.2 kN/mm 8.1 kN/mm 0.381.8 m Partial Ore 3 13.3 kN/mm 3.7 kN/mm 0.28

Host 4 16.7 kN/mm 4.4 kN/mm 0.26

6.2.6 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 D4435-13 that

measures only bolt head displacement and excludes rock mass compression is recommended.


6.3 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.

6.3.1 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 = 0.05).

0

5

10

15

20

25

30

Initi

al S

tiffn

ess

(kN

/mm

)


(a) Initial stiffness

0

20

40

60

80

100

120

Plo

ugh

Poi

nt (

kN)


(b) Plough point

0

1

2

3

4

5

6

7

Plo

ughy

Stif

fnes

sy(k

N/m

m)

InstallationDayyofytest Previous

(c) Plough stiffness

0

2

4

6

8

10

12

14

Sec

ant S

tiffn

ess

(kN

/mm

)


(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 cv t tcrit pInitial Stiffness Day-of 32 10.0 kN/mm 3.2 kN/mm 0.32

Previous 8 11.2 kN/mm 8.2 kN/mm 0.73 0.382 1.860 0.712Plough Point Day-of 36 61.8 kN 21.0 kN 0.34

Previous 14 45.1 kN 23.9 kN 0.53 2.292 2.080 0.032Plough Stiffness Day-of 31 2.00 kN/mm 0.69 kN/mm 0.35

Previous 12 3.16 kN/mm 1.56 kN/mm 0.49 2.479 2.160 0.028Secant Stiffness Day-of 33 3.97 kN/mm 2.63 kN/mm 0.66

Previous 12 4.39 kN/mm 2.06 kN/mm 0.47 0.567 2.060 0.575


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 = 0.032 and p = 0.028 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, 2004), potentially initialising the ploughing process. Simser et al. (2006) 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.

6.3.2 Geology

The Modified Cone Bolts were installed and tested in various lithologies. Figure 6.30 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 cvOre 4 6.4 kN/mm 2.4 kN/mm 0.38

Initial Stiffness Ign/Metaign 14 11.1 kN/mm 3.8 kN/mm 0.34Metased 7 9.7 kN/mm 1.9 kN/mm 0.20

Ore 4 67.0 kN 16.8 kN 0.25Plough Point Ign/Metaign 14 45.3 kN 22.3 kN 0.49

Metased 9 61.9 kN 17.2 kN 0.28Ore 4 1.40 kN/mm 0.46 kN/mm 0.33

Plough Stiffness Ign/Metaign 12 2.12 kN/mm 0.92 kN/mm 0.43Metased 8 2.13 kN/mm 0.74 kN/mm 0.35

Ore 4 131.2 kN 4.4 kN 0.03Secant Stiffness Ign/Metaign 15 4.28 kN/mm 3.62 kN/mm 0.85

Metased 11 4.46 kN/mm 3.40 kN/mm 0.76


0

5

10

15

20

25

30

Initi

alMS

tiffn

essM

(kN

/mm

) DayMofMInstallationPreviouslyMInstalled

Ore (Meta)Igneous Metaseds

(a) Initial stiffness

0

20

40

60

80

100

120

Plo

ugh

Poi

nt (

kN)


(b) Plough point

0

1

2

3

4

5

6

7

Plo

ugha

Stif

fnes

sa(k

N/m

m)


(c) Plough stiffness

0

2

4

6

8

10

12

14

16

Sec

ant S

tiffn

ess

(kN

/mm

)


(d) Secant stiffness

Figure 6.30: Performance comparison of MCB33 bolts installed in ore, igneous/metaigneous andmetasedimentary 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.

6.3.3 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 6.31.

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.


R² = 0.0009

R² = 0.3153

0

1

2

3

4

5

6

7

0 20 40 60 80 100 120

Plo

ugh/

Stif

fnes

s/rk

N/m

mv

Plough Load/rkNv

Day of/installation

Previously installed

Figure 6.31: Relationship between plough stiffness and plough point for the MCB33

6.3.4 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 time-

dependent 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 con-

trasting 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. (2004) 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, 2006). 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.


6.4 Summary

The findings of this Chapter are summarized in Table 6.20. 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 6.20: Summary of observed relationships of between various factors and performance indicatorsfor each rock bolt type

Rock Bolt Parameter Performance indicator Relationship Strength ConfidenceNominal diameter Ultimate capacity None None High

Length Ultimate capacity Positive Weak MediumInstallation method Ultimate capacity Higher with bolter, Strong High

lower with jacklegFRS Drive time Ultimate capacity Positive Strong High

Drill bit diameter Ultimate capacity Negative Weak MediumLithology Ultimate capacity None None Medium

Rock mass quality Ultimate capacity Higher in good, Strong Mediumlower in poor

Length All indicators None None MediumUnloading stiffness Tangent stiffness Positive Strong HighEmbedment length Tangent stiffness, Positive Medium High

secant stiffnessRebar Spin time All indicators None None Low

Residence time Tangent stiffness, Negative Weak Lowsecant stiffness

Lithology All indicators None None MediumPlough stiffness Plough load Negative (previously Medium Medium

installed bolts)Residence time Plough load Negative Strong High

MCB33 Plough stiffness Positive Strong MediumLithology Plough load Higher in ore, Weak Low

lower in wastePlough 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.

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, 2014). Stillborg (1993) conducted a series of tests

on various rock bolts, namely mechanical bolts, 20 mm rebar grouted in both cement and resin, resin-

grouted 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 = 60 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. (2014) 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 950 mm x 950 mm x 950 mm high-strength concrete

112

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, 2014). Doucet & Voyzelle (2012) present the results of laboratory static pull tests performed

at the CANMET-MMSL facility using the procedures and apparatus described in ASTM D7401-08,

discussed in Section 2.2.5. 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 (2012)

are from tests performed between 2006 and 2011, 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 Rx20D, 20 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 (2012) 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


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, 2014), 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, 2000) 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 (2000) 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.


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.

7.2.1 Characterization of FRS Performance

Table 7.1 shows the 10th, 25th, 50th, 75th and 90th 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 P10 P25 P50 P75 P90

Ultimate FA35 25 47.7 kN 58.0 kN 64.6 kN 71.4 kN 88.6 kNCapacity FA39 24 47.6 kN 51.2 kN 57.3 kN 69.7 kN 79.4 kN

Total 49 48.0 kN 52.8 kN 62.0 kN 70.7 kN 83.2 kNSecant FA35 25 5.1 kN/mm 6.9 kN/mm 15.0 kN/mm 22.7 kN/mm 39.7 kN/mmStiffness FA39 24 4.4 kN/mm 5.9 kN/mm 9.5 kN/mm 13.5 kN/mm 18.8 kN/mm

Total 49 4.8 kN/mm 6.5 kN/mm 10.9 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 10th

and 75th 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


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.

0

10

20

30

40

50

60

70

80

90

100

0 2 4 6 8 10 12 14 16 18 20

Load

t(kN

)

Displacementt(mm)

P90

P10

P75

P50

P25P10

P 25

P 50P 75P90

FA35

FA39

Figure 7.2: Distributions of secant stiffness and ultimate capacity for a pull test on an FRS with ananchorage 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 P10 P25 P50 P75 P90

FA35 81 25.1 kN/m 32.1 kN/m 40.9 kN/m 47.6 kN/m 52.4 kN/mFA39 30 30.9 kN/m 33.2 kN/m 36.9 kN/m 41.5 kN/m 51.1 kN/mFA46 50 24.1 kN/m 32.3 kN/m 38.4 kN/m 46.2 kN/m 53.5 kN/mFB35 92 21.7 kN/m 32.1 kN/m 39.9 kN/m 49.6 kN/m 58.4 kN/mFB39 83 23.1 kN/m 29.2 kN/m 35.0 kN/m 49.6 kN/m 57.8 kN/mFB46 106 24.3 kN/m 29.2 kN/m 38.9 kN/m 46.2 kN/m 51.1 kN/m

FA35+FA39 49 31.5 kN/m 34.6 kN/m 40.7 kN/m 46.4 kN/m 54.6 kN/mGrand Total 442 24.0 kN/m 31.5 kN/m 38.7 kN/m 47.4 kN/m 53.5 kN/m


0

10

20

30

40

50

60

70

80

Load

V(kN

/m)

FB35 FB39 FB46FA35 FA39 FA46

MedianMean

FRSVVariant

P90

P10

P75

P50

P25

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 30 and 40 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. P10 and P25 are

higher than the equivalent percentiles for other bolts, although the median, P75 and P90 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.

0

10

20

30

40

50

60

70

0.09 0.29 0.49 0.69 0.89 1.09

Load

t(kN

/m)

Displacementt(9toftanchoragetlength)

P90

P10

P75

P50

P25

P10P 25

P 50P 75P90

Figure 7.4: Distribution of secant stiffness and ultimate capacity for all FRS bolts tested


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 25th

percentiles of both ultimate capacity and secant stiffness of an FRS with an anchorage length of 2 m.

The intersect of the 25th percentiles is at 33 kN/m and 0.32 % 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.

7.2.2 Characterization of FRS Behaviour

Figure 7.5 shows the load-displacement relationships for all FRS pull tests for which displacement was

recorded. Displacements at 0.5 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.

0

10

20

30

40

50

60

70

80

90

100

0 2 4 6 8 10 12 14 16 18 20

Load

(kN

)

Displacement (mm)

FA35

FA39

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 0.5 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 90th percentile of displacement is equal to the 10th 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 10th percentile of stiffness must end at the

10th percentile of ultimate capacity, as only 90% 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


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).

0

10

20

30

40

50

60

70

80

90

100

0 2 4 6 8 10 12 14 16 18 20

Load

(kN

)

Displacement (mm)

P90

P10

P75

P50

P25

Figure 7.6: Distribution of displacement measured during pull testing of FA35 and FA39 withanchorage lengths of 1.52 m

0

10

20

30

40

50

60

70

80

90

100

0 2 4 6 8 10 12 14 16 18 20

Load

c(kN

)

Displacementc(mm)

InitialcCapacitycRange(FromcSupplier)

P90

P10

P75

P50

P25

Figure 7.7: Conceptual displacement distribution of FRSs with anchorage lengths of 1.52 m subject toa pull test

Figure 7.6 shows that in situ FRS behaviour may be quite variable; it is possible that next to no load


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.

0

10

20

30

40

50

60

70

0.09 0.29 0.49 0.69 0.89 1.09 1.29 1.49

Load

t(kN

/m)

Displacementt(9toftanchoragetlength)

P90

P10

P75

P50

P25

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

1.524 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 60% 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


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 % PassingFA35 27 kN 98.8% 17.7 kN/m 98.8 %

54 kN 66.7% 35.4 kN/m 69.1%FA39 27 kN 100.0% 17.7 kN/m 100.0%

54 kN 53.3% 35.4 kN/m 53.3%FA46 54 kN 84.0% 35.4 kN/m 64.0%

89 kN 14.0% 58.4 kN/m 4.0%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 78.0% 35.4 kN/m 54.7%

89 kN 16.0% 58.4 kN/m 0.9%

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.


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 0 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.


7.3.1 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 P10 P25 P50 P75 P90

A 34 112.6 kN 118.6 kN 124.1 kN 127.7 kN 132.8 kNWorking capacity B 20 107.7 kN 115.7 kN 123.8 kN 124.6 kN 129.6 kN

Total 54 112.6 kN 117.5 kN 124.1 kN 127.1 kN 131.1 kN

A 23 9.1 kN/mm 15.3 kN/mm 28.9 kN/mm 43.8 kN/mm 46.9 kN/mmSecant stiffness B 21 8.1 kN/mm 15.3 kN/mm 27.9 kN/mm 35.2 kN/mm 46.6 kN/mm

Total 44 9.2 kN/mm 15.8 kN/mm 28.1 kN/mm 39.5 kN/mm 45.9 kN/mm

0

20

40

60

80

100

120

140

160

0 2 4 6 8 10 12 14 16 18 20

Load

A(kN

)

DisplacementA(mm)

RebarAA

RebarAB

P90

P10

P75P50P25

P90

P 10

P 75

P 50

P 25

Figure 7.9: Distributions of secant stiffness and working capacity for a pull test with a pre-load of 17.8kN on a rebar rock bolt

7.3.2 Characterisation of Rebar Rock Bolt Behaviour

Figure 7.10 shows all pull tests performed. Tests performed with a pre-load of 0 are shown in red, a

pre-load of 17.8 kN in grey, and 26.7 kN in black.


0

20

40

60

80

100

120

140

160

0 2 4 6 8 10 12 14 16 18 20

Load

(kN

)

Displacement (mm)

3 t (26.7 kN) Pre-load2 t (17.8 kN) Pre-load0 Pre-load

Figure 7.10: All rebar pull test load-displacement relationships

It can be seen in Figure 7.10 that large displacements at low loads may occur if the bolt is not pre-

loaded. 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 75th and 90th percentiles not

clearly showing yield behaviour.

0

20

40

60

80

100

120

140

160

0 2 4 6 8 10 12 14 16 18 20

Load

(kN

)

Displacement (mm)

P90

P10

P75

P50

P25

Figure 7.11: Distribution of displacement measured during pull testing of 20 mm rebar 1.8 m to 2.4 min 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;


0

20

40

60

80

100

120

140

160

0 2 4 6 8 10 12 14 16 18 20

Load

(kN

)

Displacement (mm)

P90

P10

P75

P50

P25

Figure 7.12: Distribution of displacement measured during pull testing of 20 mm rebar 1.8 m to 2.4 min 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 7.13. The behaviour shown in Figure

7.12 is extrapolated to P90 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.

0

20

40

60

80

100

120

140

160

0 2 4 6 8 10 12 14 16 18 20

Load

W(kN

)

DisplacementW(mm)

P90P10 P75P50P25

PercentilesWofWWorkingWCapacityP90

P10

P75P50P25

131WkN127WkN124WkN117WkN113WkN

Figure 7.13: Conceptual distribution of displacement during pull testing of 20 mm rebar without apre-load


Figure 7.13 shows that less than 2 mm of displacement may be observed when pull testing a 20 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 displace-

ment 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 % PassingMinimum thread yield strength 89 kN (Supplier A), 86 kN (Supplier B) 100%

Minimum thread tensile strength 134 kN (Supplier A), 116 kN (Supplier B) 100%

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 MCB33.

7.4.1 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 P10 P25 P50 P75 P90

Initial Stiffness 57 5.7 kN/mm 7.5 kN/mm 9.7 kN/mm 11.4 kN/mm 15.2 kN/mmPlough Load 67 26.7 kN 35.4 kN 53.4 kN 68.4 kN 91.6 kNPlough Stiffness 58 1.2 kN/mm 1.4 kN/mm 2.1 kN/mm 2.8 kN/mm 3.7 kN/mmYield Strength 48 106.8 kN 115.7 kN 125.1 kN 130.3 kN 136.7 kNSecant Stiffness 43 1.5 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.


Displacement (mm)

0

20

40

60

80

100

120

140

160

0 10 20 30 40 50 60 70 80 90 100

Load

(kN

)

Displacement (mm)

P90

P10

P75P50

P25P 90 P10

P 75 P 50 P 25

Figure 7.14: Distributions of secant stiffness and working capacity for a pull test with a pre-load of17.8 kN on an MCB33

7.4.2 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 5.4.3 is seen for

most of the pull tests. As the objective of pull testing cone bolts is to demonstrate their large displace-

ment 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.

0

20

40

60

80

100

120

140

160

0 10 20 30 40 50 60 70 80 90 100

Load

m(kN

)

Displacementm(mm)

4mtm(35.6mkN)mPre-load3mtm(26.7mkN)mPre-load2mtm(17.8mkN)mPre-load

Figure 7.15: 56 MCB33 pull tests collected from Vale’s Sudbury operations


Percentiles of displacement are presented in Figure 7.16. 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; 107 kN) by eliminating dummy displacements from the

percentile calculations at 107 kN. Above 107 kN, dummy displacements were reintroduced to maintain

the n = 32 dataset as it decreased in size at higher values of load.

0

20

40

60

80

100

120

140

160

0 10 20 30 40 50 60 70 80 90 100

Load(kN)

Displacement (mm)

P90

P10

P75

P50P25

Figure 7.16: Distribution of displacement measured during pull testing of a 2.4 m MCB33 with apre-load of 17.8 kN

Figure 7.16 clearly shows the two distinct behaviours of the cone bolt described in Section 5.4:

the 10th and 25th 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 7.16. Using the initial stiffness of the MCB33s in the

database, percentiles are extrapolated to 0 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 100 mm. Extreme cases of low plough fall below P10, 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 % PassingMinimum thread yield strength 98.5 kN 100%Typical thread yield strength 113.6 kN 81%


0

20

40

60

80

100

120

140

160

0 10 20 30 40 50 60 70 80 90 100

Load

f(kN

)

Displacementf(mm)

PercentilesfoffYieldfLoadP90

P10

P75P50P25

137fkN130fkN125fkN116fkN107fkN

P90P10 P75P50P25

Figure 7.17: Conceptual distribution of displacements for a pull test of a 2.4 m MCB33 without apre-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.

7.5.1 Characterisation of D-Bolt Performance

Figures 7.18 and 7.19 show percentiles of stiffness plotted with minimum and typical yield loads of the 20

mm and 22 mm D-Bolts according to Normet (2014). 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 20 mm and 22 mm D-Bolts

Bolt n P10 P25 P50 P75 P90

20 mm D-Bolt 12 6.5 kN/mm 8.5 kN/mm 23.3 kN/mm 35.8 kN/mm 59.5 kN/mm22 mm D-Bolt 19 3.9 kN/mm 5.3 kN/mm 11.1 kN/mm 24.3 kN/mm 26.1 kN/mm


0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

y(kN

)

Displacementy(mm)

MinimumyYieldyLoadTypicalyYieldyLoadP 90P

10

P 75P50

P25

Figure 7.18: Distribution of secant stiffness for a pull test with a pre-load of 17.8 kN on a 20 mmD-Bolt

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

y(kN

)

Displacementy(mm)

MinimumyYieldyLoad

TypicalyYieldyLoad P90P10

P 75P 50P25

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 dis-

placements and behaviours, with the 22 mm bolt acting in a less stiff manner than a 20 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 config-

urations 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


to implement the D-Bolt into a ground support standard.

7.5.2 Characterisation of D-Bolt Behaviour

Figures 7.20 and 7.21 show the pull tests collected for the 20 mm and 22 mm D-Bolts. It can be seen

that very few of the bolts yield during testing.

20 25 30

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

(kN

)

Displacement (mm)

Figure 7.20: 20 mm D-Bolt pull tests collected from Vale’s Sudbury operations

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

(kN

)

Displacement (mm)

Figure 7.21: 22 mm D-Bolt pull tests collected from Vale’s Sudbury operations

Figures 7.20 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


relatively high stiffness is observed in Figure 7.21. 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.

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load(kN)

Displacement (mm)

P90

P10

P75

P50P25

Figure 7.22: Load–displacement behaviour of a 20 mm D-Bolt with a pre-load of 17.8 kN

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load(kN)

Displacement (mm)

P90

P10

P75

P50P25

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 P75 and P90

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


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, 2014). Behaviours are extrapolated to 0 kN.

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

y(kN

)

Displacementy(mm)

MinimumyYieldyLoadTypicalyYieldyLoadP90P10 P75P50P25

Figure 7.24: Conceptual distribution of displacements for a pull test of a 20 mm D-Bolt without apre-load

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

y(kN

)

Displacementy(mm)

MinimumyYieldyLoad

TypicalyYieldyLoadP90P10 P75P50P25

Figure 7.25: Conceptual distribution of displacements for a pull test of a 22 mm D-Bolt without apre-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


capacity/yield was captured during the test, it always occurred above the minimum yield load specified.

This includes 3 20 mm D-Bolts yielding above 140 kN, and 3 22 mm D-Bolt yielding above 170 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 20 kN greater than that of the Mn24; Atlas Copco,

2012). 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, 2005).

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.

7.6.1 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 25th and 50th 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.

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

(kN

)

Displacement (mm)

P90

P10

P75P50P25/

P 75 P 50P25

Figure 7.26: Distributions of secant stiffness and working capacity for a pull test on Swellex Pm12 andMn12 without a pre-load


Table 7.9: Swellex secant stiffness percentiles

Metric Bolt n P10 P25 P50 P75 P90

Working capacity Pm12 & Mn12 17 84.5 kN 89 kN 89 kN 93.4 kN 99.6 kNSecant Stiffness Pm12 & Mn12 8 N/A 5.0 kN/mm 5.8 kN/mm 10.1 kN/mm N/A

Pm24 & Mn24 16 4.8 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 90 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.

7.6.2 Characterisation of Expandable Bolt Behaviour

Figures 7.27 and 7.28 show pull test data collected for the four configurations of Swellex.

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

(kN

)

Displacement (mm)

Figure 7.27: Swellex Pm12 and Mn12 pull tests collected from Vale’s Sudbury operations


0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

(kN

)

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 7.30. Insufficient data was available to calculate

10th and 90th 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 Mn24.

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

(kN

)

Displacement (mm)

P75P50P25

Figure 7.29: Load–displacement behaviour of Pm12 and Mn12 bolts


0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load(kN)

Displacement (mm)

P75P50P25 P90P10

Figure 7.30: Load-displacement behaviour of Pm24 and Mn24 bolts

Figures 7.29 and 7.30 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.

0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

(kN

)

Displacement (mm)

Minimum Yield Load (Pm12, Mn12)

P75P50P25

Figure 7.31: Conceptual load–displacement behaviour of Pm12 and Mn12 bolts subject to a pull test


0

50

100

150

200

250

0 5 10 15 20 25 30 35 40 45 50

Load

(kN

)

Displacement (mm)

P90P10 P75P50P25

Minimum Yield Load (Pm24)

Minimum Yield Load (Mn24)

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 7.10. As displacement data was collected in 1 ton (8.9 kN) load increments,

one of the data bins is the 89 kN - 97.9 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 90 kN. As such, bolts with a recorded working capacity of 89 kN are assumed

to have passed the 90 kN threshold specified by the supplier.

Table 7.10: Swellex Pm12 and Mn12 performance compared to supplier specifications

Metric Specification % PassingMinimum yield strength (Pm12, Mn12) 90 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 90 kN. Additionally, the supplier’s specifications for the Pm24 and Mn24 should be verified by

pulling these bolts to loads above 200 kN and 180 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, 2012). 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 0 load as a result of a lack of data.


0

50

100

150

200

250

0 20 40 60 80 100

Load

(kN

)

Displacement (mm)

(a) FA35 and FA39

0

50

100

150

200

250

0 20 40 60 80 100

Load

y(kN

)

Displacementy(mm)

MedianyWorkingyCapacity

(b) Rebar

0

50

100

150

200

250

0 20 40 60 80 100

Load

(kN

)

Displacement (mm)

(c) MCB33

0

50

100

150

200

250

0 20 40 60 80 100

Load

(kN

)

Displacement (mm)

Typical Yield Load

(d) 20 mm D-Bolt

0

50

100

150

200

250

0 20 40 60 80 100

Load

(kN

)

Displacement (mm)

(e) 22 mm D-Bolt

0

50

100

150

200

250

0 20 40 60 80 100

Load

(kN

)

Displacement (mm)

Minimum Yield Load

(f) Swellex Pm12 and Mn12

0

50

100

150

200

250

0 20 40 60 80 100

Load

(kN

)

Displacement (mm)

Minimum Yield Load (Pm24)Minimum Yield Load (Mn24)

(g) Swellex Pm24 and Mn24

Figure 7.33: Conceptual load–displacement behaviour extrapolated to 0 load for various rock boltssubject to a pull test


Figure 7.34 shows the median behaviours of all bolts discussed in this chapter extrapolated to a load

of 0 kN.

MinimumgYieldgLoad

,

5,

9,,

95,

2,,

25,

, 2, 4, 6, 8, 9,,

Lo

ad

g8kN

(

Displacementg8mm(

FA35sgFA39

MCB33

22gmmgDFBolt

SwellexgPm24

SwellexgMn24

2,gmmgDFBolt

2,gmmgRebar

SwellexgPm923Mn92

MediangWorkinggCapacityTypicalgYieldgLoad

Figure 7.34: Median load–displacement behaviour for all bolts with no pre-load


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 20 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 20 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 7.34. 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 (150 kN) for the rebar

tested by Stillborg (1993), while median rebar working capacity is 124.1 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 (108 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.

Chapter 8

Conclusions

In the final chapter, the contributions made by this thesis are outlined. Limitations are discussed,

recommendations made and the path forward laid out.

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 2003 and 2015 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

D4435-13) 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

140

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 0.15 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 40.0 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 0.3

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 123.8 kN with a standard deviation of 7.5

kN. Rebar B had an average working capacity of 120.9 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 im-

mediately after installation, with an average value of 53.2 kN and standard deviation of 22.7 kN. The

average MCB33 yield strength was 117.2 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


of many rock bolt types, and although the ASTM pull testing standard (D4435-13) 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 D4435-08

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 un-

noted 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 (2000) 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


regards to information recorded and pull test practices.

In addition to those outlined by Soni (2000), 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. If load is

recorded in tons, then short, long, or metric should be specified.

• 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 0.5 tons once load has been

built to within 20 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, un-

loading 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


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 funda-

mental 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 perfor-

mance, 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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Appendices

151

Appendix A

Pull Testing Forms

A.1 ASTM D4435-13 Sample Form

Figure A.1: Rock bolt pull test sample form (ASTM D4435, 2013)

152

Appendix A. Pull Testing Forms 153

A.2 ISRM Suggested Method for Pull Testing Data Sheet

ROCKpANCHORpTEST RESULTpSHEET TESTpq

Datepofpinstallation: Datepofptestp:

Project: Location:

Anchor: Type: Length: InstallationpTorque:

ROCK: Classification: Fracturepspacing: Strength:

BOLT: Type: Length: Untensionedplength:

HOLE: Diameter: Length: Orientationpandproughness:

Pumpppressure Boltptension Displacementpreadings RemarksReading Displacement Reading Displacement Average

TESTpRESULTS: Maximumppullpforce:

Displacementpatpmaximumppullpforce: Maxbpdisplacementpinptest:

Naturepofpfailureporpyield:

Otherpremarks:

Testedpby: Checkedpby:

Figure A.2: Rock bolt pull test data sheet (After ISRM, 1981)

Appendix A. Pull Testing Forms 154

A.3 Proposed Pull Test Information Sheets

Pull(Test(Campaign(Information Sheet

AttentionBof: PreparedBby:

MineNcompany: MineNcompany:

TestBdateB EMMNDDNYYG:

PullBtestBobjective:(e.g.wqualitywcontrol,w

productwverification,wetc.)

TestBLocation: BoltBinstallationBdateBEMMNDDNYYG:

TemperatureBduringBinstallation: oC NB oF Wet(bulb( Dry(bulb

RockwMasswParameters

RockBtype: UCSBEMPaG: EBEGPaG:

RockBmassBquality: RMR: to Q: to RQD: toGSI: to OtherBEspecifyGB : to

Notes:B______________________________________________________________________________________

BoltwandwAnchor Parameters

RockBboltBtype: BoltBlength: m N ’ ”B

GalvanizationNcoating: NominalBboltBdiameter: mm N ”

AnchorBtype: InstallationBtorque:

GroutBtype: Resin Cement Measured bolt diameter: mm N ”

GroutBdescription:B(includingwcartridgew

lengths,ww/cwratio,wetc.)

Equipment andwApparatus

DrillNinstallationBequipment: MeasuredBdrillBbitBdiameter: mm N ”

LoadingBpumpBtype: Manual Electric DataBrecording: Manual Electronic

LoadBunits: kN t((short)( t((long)(t((metric)(

LoadBintervalsBorBmeasurementBfrequency:

DisplacementBunits: mm( in( Preload: kN N t

DisplacementBmeasured: Ram(travel( Bolt(head(relative(to(stable(point(

AdditionwComments

Pull Test Data Sheet1 2 3 4 5 6 7 8 9 10

MeasuredSdiameterDrillholeSdiameter

DriveS/SInflationS/SSpinStime

PartialSembedmentSorSencapsulation?

Preload

NOTES Load Disp Load Disp Load Disp Load Disp Load Disp Load Disp Load Disp Load Disp Load Disp Load Disp

StoppingScriteria/SnatureSofSfailure

Notes

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