block caving geomechanics s

Block Caving Geomechanics

E.T. Brown

JKMRC Monograph Series in Mining and Mineral Processing 3

JULIUS KRUTISCHNITI MINERAL RESEARCH CENTRE THE UNIVERSITY OF QUEENSLAND

Published by: Julius Kruttschnitt Mineral Research Centre Isles Road, Indooroopilly, Queensland 4068, Australia

Copyright © 2002 Julius Kruttschnitt Mineral Research Centre, The University of Queensland

All Rights Reserved. No part of this publication may be reproduced, stored in a retrieval system or transmitted in any form or by any means: electronic, electrostatic, magnetic tape, mechanical, photocopying, recording or otherwise, without permission in writing from the publishers.

National Library of Australia Cataloguing-in-Publication Entry:

Block caving geomechanics ISBN 1-74112-000-4 L Caving mining. 2. Stoping (Mining). 3. Ground control (Mining). I. Brown, E.T. n. Julius Kruttschnitt Mineral Research Centre. (Series: JKMRC monograph series in mining and mineral processing, No. 3).

Printed in Australia by University of Queensland Print On Demand Centre Cover production by University of Queensland, Brisbane

This book can be ordered directly from the Publisher: e-mail: [email protected]

Phone: +61 73365 5888 Fax: +61 733655999

ii

FOREWORD

It is important that research outcomes be disseminated in a useful form to the clients of the

research, and to the community at large if appropriate. The research monograph is a traditional

mechanism for reporting substantial bodies of research which, taken together, advance the field

to a significant degree. In 1996 the JKMRC published two such monographs, on comminution

and blasting, in a series on mining and mineral processing. The present volume continues the

series.

Caving is a mining method which is of growing interest to companies concerned with the

exploitation of massive ore deposits, because of its low cost. However the body of theoretical

and practical knowledge of the factors controlling the caving process, particularly in competent

rock masses, is limited, and the economic risk in developing a caving mine can therefore be

higher than one would like. To remedy this situation, a number of major mining companies

came together in 1997 to fund the International Caving Study, a wide-ranging research and

technology transfer project. This book records in part the results of the first phase ofthis study.

The JKMRC as the lead researcher in the rcs has been fortunate in its collaborators: the nine

companies who funded and directed the work, its research partner the Itasca Consulting Group,

Dr. Denis Laubscher, and a number of other consultants. An these are acknowledged in detail

in the book. On a personal note, I would like to recognise the good work done by the JKMRC

research staff and postgraduate students under the able direction of Dr. Gideon Chitombo.

Finally, we thank our distinguished colleague Professor Ted Brown for taking on the

demanding task of bringing the book to fruition, which he has completed with his customary

skill and energy.

I hope that the mining and geomechanics communities will find the book of interest and value.

iii

T.J. Napier-Munn

Director - JJO.tlRC

ACKNOWLEDGEMENTS

This book is an outcome of the International Caving Study Stage I carried out in the period

1997-2000 by the Julius Kruttschnitt Mineral Research Centre (JKMRC), The University of

Queensland, Brisbane, Australia, and the Itasca Consulting Group, Inc, Minneapolis, USA. The

following sponsoring companies and their representatives who monitored the progress of the

Study are thanked for their support of the Study and the preparation of this book:

De Beers Consolidated Mines Limited

CODELCO-Chile

Newcrest Mining Limited

Noranda Inc.

Northparkes Mines

PT Freeport Indonesia

Rio Tinto Limited

TVX Gold Inc.

This book draws heavily on the original work carried out for the International Caving Study

Stage I by researchers from the JKMRC and Itasca, consultants to the Study and several JKMRC research students. I wish to acknowledge, in particular, the important contributions made by:

• Dr Gideon Chitombo, JKMRC, who was the guiding force behind the Study and arranged for me, and this book, to be part of it. He wrote the first drafts of Section 4.3.6 and, with

the assistance of Italo Onederra of the JKMRC, wrote Section 5.6. He made a number of valuable suggestions about the contents of several other chapters, provided a great many

pieces of information that are included in the book, prepared the initial version of the index

and managed the arrangements for the book's production;

• Dr Bob Trueman, JKMRC, who supervised and personally carried out much of the work

reported in Sections 3.1,3.2,3.3 and 5.5, and substantially wrote those sections;

v

• Dr Loren Lorig, Itasca Consulting Group, who carried out the analyses for, and wrote the

original versions of, Sections 3.5 and 3.6, Appendix B and, in conjunction with Dr Peter

Cundall of Itasca, Appendix C;

• Matt Pierce, ltasca Consulting Group, who, with Dr Bob Trueman and Ridho Wattimena,

carried out the numerical analyses reported in Section 5.5;

• Dr Geoff Lyman, JKMRC, who carried out the original analyses for, and wrote the initial

version of, Section 2.6.3;

• JKMRC PhD students Neal Harries, Clare Mawdesley, Brian Eadie and Ridho Wattimena

whose research work made significant contributions to Chapters 2, 3, 4 and 5, respectively;

• Alan Cocker, JKMRC, who developed the software for the JointStats system reported in Section 2.5.7;

• David La Rosa, JKMRC, who developed the software for the CaveRisk system reported in Section 11.5;

• Dr Dennis Laubscher whose Block Cave Manual, including the contributions made by Nick

Bell and Glen Heslop, provided an invaluable source of information, ideas and illustrations, many of which appear in the book; and

• John Summers, CGSS, Berkshire, England, who, with input from Dr Gideon Chitombo and others, developed the CaveRisk system described in Chapter 11 and wrote the report on

which that chapter is based.

In October, 200 I, copies of the first draft of the book were distributed to the sponsors of the

International Caving Study Stages I and 11 for comment. I am grateful to the representatives of the sponsors for their support in this final stage of the process. I would especially like to thank

the following individuals for having provided valuable comments on parts of the draft and/or

additional material for inclusion in the book:

• Richard Butcher, WMC;

• Joaquin Cabello, Golder Associates;

• Dr Gideon Chitombo, JKMRC;

• German Flores, Chuquicamata Division, CODELCO-Chile;

• Dr Antonio Karzulovic, A Karzulovic & Associates;

• Craig Stewart, Northparkes Mines; and

• Dr Duncan Tyler, Newcrest Mining.

vi

I also wish to thank those who gave their support and assistance to this undertaking in a number

of important ways, especially:

• Libby Hill, JKMRC, who undertook the desktop publishing with her usual skill, grace, and

efficiency;

• Vynette Holliday and Naomi Mason, JKMRC, who assisted Libby in this process by

preparing many of the figures;

• the former Manager of the Dorothy Hill Physical Sciences and Engineering Library,

University of Queensland, Gulcin Cribb, and Library staff member, Diana Guillemin, for

their assistance in providing copies of a large number of sometimes obscure references;

• Rob Morphet and the partners and staff of the Brisbane office of Golder Associates Pty Ltd

for providing me with facilities, encouragement and support during the writing of parts of the book;

• John Markham, CEO, Itasca Consulting Group, for his efficient project administration; and

• my partner, Dr Dale Spender AM, for her continuing tolerance of my interest in holes in the

ground and for understanding that "the Earth sucks".

Finally, but most importantly, I should like to acknowledge my debt of gratitude to the

foundation Director of the JKMRC, Professor Alban Lynch AO FTSE, for inviting me to

become involved with the work of the Centre when I joined the staff of the University of

Queensland in late 1987. I also wish to record my appreciation to his disciples and successors

as Directors of the Centre, Professors Don McKee and Tim Napier-Munn, who have continued to make me welcome at the JKMRC in the intervening years. Without their friendship and

support, I would not have had the opportunity, or been able, to prepare this book.

E TBrown

Brisbane

29 March 2002

vii

CONTENTS

.FOREWORD ......................................................................... iii ACKNOWLEDGEMENTS ....................................................... v

CONTENTS ......................................................................... viii

CHAPTERl INTRODUCTION

1.1 UNDERGROUND MINING METHODS ................................................................... 1 1.1.1 General Features ........................................................................................ 1

1.1.2 Classification of Underground Mining Methods ............................................... l

1.2 BLOCK AND PANEL CAVING .............................................................................. 3

1.2.1 Outline of the Method ................................................................................ .3

1.2.2 Basic Caving Mechanics .............................................................................. 8

1.2.3 History of Block Caving ............................................................................ 12

1.3 BLOCK AND PANEL CAVING OPERATIONS ...................................................... 16

1.3.1 Overview ................................................................................................ 16

1.3.2 El Teniente Mine, Chile ............................................................................. 16

1.3.3 Premier Diamond Mine, South Africa .......................................................... 20

1.3.4 Henderson Mine, Colorado, USA ................................................................ 24

lA RISK IN CAVE MINING ...................................................................................... 27

1.4.1 Risk Factors ........................................................................................... .

1.4.2 Introduction to Risk Assessment ................................................................. 29

1.5 SCOPE Al'l"D CONTENTS OF THIS BOOK ............................................................. 30

viii

CHAPTER 2

ROCK MASS CHARACTERISATION

2.1 DEFINING THE MINING ENVIRONMENT ....................................................... 32

2.2 GENERAL DATA REQUIREMENTS ................................................................ 33

2.2.1 Geology ............................................................................................. 33 2.2.2 Surface and Groundwater Hydrology ....................................................... 35

2.2.3 Topography and Environmental Constraints .............................................. 35

2.2.4 Geotechnical Studies ............................................................................ 35

2.3 CLASSIFICATION AND DESCRIPTION OF DISCONTINUITIES ........................ 36

2.3.1 Classification ...................................................................................... 36 2.3.2 Description ......................................................................................... 41

2.4 DISCONTINUITY DATA COLLECTION BY DRILLING, CORE LOGGING,

DOWN-HOLE SURVEYS, SCANLINE AND CELL MAPPING ............................ .42

2.4.1 Introduction ........................................................................................ 42 2.4.2 Geotechnical Core Logging ................................................................... 43

2.4.3 Exposure Mapping Methods .................................................................. 49 2.5 ANALYSIS AND PRESENTATION OF DISCONTINUITY DATA ........................ 55

2.5.1 Introduction ........................................................................................ 55 2.5.2 Error and Uncertainty in Discontinuity Analysis ........................................ 56

2.5.3 Discontinuity Orientation Analysis .......................................................... 58 2.5.4 Discontinuity Frequency/Spacing (Intensity) Analysis ................................ 60 2.5.5 Discontinuity Persistence (Size) Analysis ................................................. 64 2.5.6 Definition ofGeotechnical or Structural Domains ...................................... 66 2.5.7 JK Jointstats Discontinuity Data Management System ................................ 67

2.6 SIMULATION OF ROCK MASS GEOMETRy ................................................... 77 2.6.1 Introduction ........................................................................................ 77 2.6.2 Approaches to Discontinuity Modelling ................................................... 78 2.6.3 The Development of the JKMRC 3-D Discontinuity Model ......................... 85 2.6.4 The JKMRC Hierarchical Model of Discontinuity Network Geometry ........... 92

2.7 ROCK MASS CLASSIFICATION SCHEMES ................................................... 100 2.7.1 Introduction ...................................................................................... 100 2.7.2 RMR System (Bieniawski, 1974,1976) .................................................. 101 2.7.3 Q System (Barton et a/1974) ............................................................... 105

2.7.4 Modified Basic RMR or MBR System (Kendorski et a/ 1983) .................... 108 2.7.5 MRMR System (Laubscher 1990) ......................................................... 109 2.7.6 In situ Rock Mass Rating or IRMR (Laubscher and lakubec 2001) .............. III 2.7.7 Geological Strength Index (GS1) ........................................................... 114 2.7.8 Conclusions ...................................................................................... 116

2.8 THE MECHANICAL PROPERTIES OF ROCK MASSES .................................... 117 2.8.l Scope .............................................................................................. 117 2.8.2 The Hoek-Brown Empirical Strength Criterion ........................................ 117 2.8.3 Rock Mass Deformation Modulus ......................................................... 122

2.9 IN SITU STRESSES ...................................................................................... 123

ix

CHAPTER 3 CA V ABILITY ASSESSMENT

3.1 INTRODUCTION ......................................................................................... 126 3.2 LAUBSCHER'S CAVING CHART .................................................................. 127

3.2.1 Overvie'v .......................................................................................... 127

3.2.2 The Mining Rock Mass Rating ............................................................. 128 3.2.3 Delineation of Zones of Stability ........................................................... 12 9

3.2.4 Summary .......................................................................................... 130 3.3 MATHEWS' STABILITY GRAPH APPROACH ................................................ 130

3.3.1 Overview .......................................................................................... 130 3.3.2 Extension of the Method ...................................................................... 133 3.3.3 Application of Mathews' Method to the Prediction of Cavability ................. 136

3.4 NUMERICAL MODELLING APPROACHES .................................................... 138 3.5 AXISYMMETRIC CONTINUUM MODEL ....................................................... 139

3.5.1 Model Formulation ............................................................................. 139 3.5.2 Material Parameters ............................................................................ 143 3.5.3 Results ............................................................................................. 144

3.6 PFC3D DISCONTINUUM MODEL ................................................................. 147 3.6.1 Introduction ...................................................................................... 147 3.6.2 Model Description .............................................................................. 148

3.6.3 Results of Model Observations ............................................................. 151

3.6.4 Future PFC Modelling of Cavability ...................................................... 154

CHAPTER 4

FRAGME~TATION ASSESSMENT

4.1 INTRODUCTION ......................................................................................... 156

4.2 FACTORS INFLUENCING FRAGMENTATION ............................................... !57

4.3 FRAGMENTATION MEASUREMENT ............................................................ 159 4.3.1 Overview .......................................................................................... 159

4.3.2 Digital Image Processing Methods ........................................................ 161 4.3.3 Examples of DIP Systems .................................................................... 162

4.3.4 Validation Studies .............................................................................. 165

4.3.5 Application of DIP Systems to Caving ................................................... 166

4.4 IN SITU FRAGMENTATION ASSESSMENT .................................................... 169 4.5 BCF: A PROGRAl\tl TO PREDICT BLOCK CAVE FRAGMENTATION ............... 172

4.5.1 Modelling Approach ........................................................................... 172

4.5.2 Primary Fragmentation ........................................................................ 173 4.5.3 Secondary Fragmentation .................................................................... 175

4.5.4 Hangup Analysis ................................................................................ 178

4.5.5 Discussion ........................................................................................ 179

x

4.6 AN ALTERNATIVE METHOD OF ASSESSING IN SITU AND

PRIMARY FRAGMENTATION ...................................................................... 181

4.6.1 Methodology .................................................................................... 181

4.6.2 Tessellation Procedure ........................................................................ 183

4.6.3 In situ Blocks .................................................................................... 186

4.6.4 Primary Fragmentation ....................................................................... 187

4.7 CONCLUSIONS ..................................................................................... 190

CHAPTERS

CAVE INITIATION BY UNDERCUTTING

5.1 INTRODUCTION ......................................................................................... 191

5.2 UNDERCUTTING STRATEGIES ................................................................... 192

5.2.1 Purpose ............................................................................................ 192

5.2.2 Post-Undercutting .............................................................................. 192

5.2.3 Pre-Undercutting ............................................................................... 193

5.2.4 Advance Undercutting ........................................................................ 194

5.2.5 The Henderson Strategy ...................................................................... 195

5.3 UNDERCUT DESIGN AND MANAGEMENT .................................................. 196

5.3.1 Purpose ............................................................................................ 196 5.3.2 Initiation and Direction of Undercut Advance .......................................... 196

5.3.3 Shape of the Undercut Face ................................................................. 199

5.3.4 Rate of Undercut Advance ................................................................... 200 5.3.5 Undercut Height ................................................................................ 202

5.4 UNDERCUT SHAPE AND EXTRACTION METHOD ........................................ 204 5.4.1 Introduction ...................................................................................... 204 5.4.2 Fan Undercut .................................................................................... 205 5.4.3 Flat Undercut .................................................................................... 206

5.4.4 Narrow Inclined UndercuL .................................................................. 210 5.5 STRESSES INDUCED IN THE UNDERCUT AND EXTRACTION LEVELS ......... 212

5.5.1 Introduction ...................................................................................... 212

5.5.2 Modelling Strategy ............................................................................. 215 5.5.3 Extraction Level Stresses Post-Undercut Sequence ................................ 217

5.5.4 Extraction Level Stresses - Advance Undercut Sequence .......................... 217

5.5.5 Undercut Level Stresses ...................................................................... 221 5.5.6 Summary of Parametric Study Results ................................................... 223

5.5.7 Undercut Drift Support and Reinforcement.. ........................................... 224

xi

5.6 DRILLING AND BLASTING FOR UNDERCUTTING AND DRA WBELL

CONSTRUCTION ............................................................................................. 226

5.6.1 Introduction .......................................................................................... 226

5.6.2 Factors affecting Drilling and Blasting Performance ..................................... 227

5.6.3 Experienced based Design "Rules of Thumb" for Rock Breakage ControL ....... 229

5.6.4 Undercut Drilling and Blasting ................................................................. 234

5.6.5 Drawbell Blasting .................................................................................. 241

5.6.6 Drilling Equipment Selection ................................................................... 243

CHAPTER 6

EXTRACTION LEVEL DESIGN

6.1 PURPOSE ........................................................................................................ 245

6.2 FACTORS INFLUENCING EXTRACTION LEVEL DESIGN A"'ID

PERFORMANCE .............................................................................................. 246

6.3 EXTRACTION LEVEL LAyOUTS ...................................................................... 248

6.3.1 Scope ................................................................................................... 248

6.3.2 Continuous Trough or Trench Layout ........................................................ 248

6.3.3 Herringbone Layout. ............................................................................... 250

6.3.4 Offset Herringbone Layout ...................................................................... 250

6.3.5 Henderson or Z Layout ........................................................................... 252

6.3.6 El Teniente Layout ................................................................................. 252

6.3.7 Ore Crushing and Transportation ............................................................... 253

6.4 DRA WPOINT AND DRA WBELL DESIGN .......................................................... 255

6.4.1 Gravity Flow of Caved Ore ...................................................................... 255

6.4.2 Drawpoint Spacing ................................................................................. 259

6.4.3 Drawpoint Size, Shape and Orientation ...................................................... 266

6.4.4 Dra\vbell Geometry ................................................................................ 268

6.5 SUPPORT AND REINFORCEMENT ................................................................... 270

6.5.1 Terminology ......................................................................................... 270

6.5.2 Principles ............................................................................................. 272

6.5.3 Support and Reinforcement Elements ......................................................... 274

6.5.4 Stress-Strength Analyses ......................................................................... 275

6.5.5 Support and Reinforcement of Draw points .................................................. 279

6.5.6 Examples .............................................................................................. 281

xii

CHAPTER 7

DRAW CONTROL

7.1 INTRODUCTION .............................................................................................. 293

7 .2 DRAW MECHANISMS ...................................................................................... 295 7.2.1 Basic Studies ......................................................................................... 295 7.2.2 Mass Flow ............................................................................................. 296

7.2.3 Granular or Gravity Flow ......................................................................... 296

7.2.4 Void Diffusion ....................................................................................... 297

7.2.5 Practical Implications .............................................................................. 299 7.3 DRAW CONTROL DURING UNDERCUTTING AND CAVE INITIATION ............... 30 I 7.4 DRAW CONTROL DURING PRODUCTION ........................................................ 305

7.4.1 Manual Calculation of Draw Tonnages and Estimation ofDilution .................. .305

7.4.2 Draw Control Strategies and Procedures ...................................................... 309 7.5 EXAMPLES OF COMPUTERISED DRAW CONTROL SYSTEMS ........................... 311

7.5.1 PC-BC .................................................................................................. 311 7.5.2 De Beers' Linear Programming Based System .............................................. 318

CHAPTERS

GEOTECHNICAL MONITORING

8.1 THE PURPOSES OF MONITORING .................................................................... 322 8.2 GEOTECHNICAL MONITORING SySTEMS ....................................................... 324

8.2.1 General Considerations ............................................................................ 324 8.2.2 What is Monitored? ................................................................................. 324 8.2.3 How is it Monitored? ............................................................................... 326 8.2.4 Where and When is it Monitored? .............................................................. 328

8.3 MONITORING THE INITIATION AND DEVELOPMENT OF CAVING ................... 331 8.3.] Why? ................................................................................................... 331 8.3.2 What and How? ..................................................................................... .33]

8.4 EXTRACTION LEVEL AND INFRASTRUCTURE MONITORING .......................... 337 8.4.1 Why? .................................................................................................. .337 8.4.2 What and How? ..................................................................................... 337 8.4.3 Examples ............................................................................................. .338

8.5 MONITORING SUBSIDENCE AND GROUND MOVEMENT ................................. 342 8.5.1 Why? ................................................................................................... 342 8.5.2 What and How? ..................................................................................... 344 8.5.3 Examples .............................................................................................. 344

xiii

CHAPTER 9 SURFACE SUBSIDENCE

9.1 INTRODUCTION ............................................................................................. 346

9.2 TYPES AND MECHANISMS OF DISCONTINUOUS SUBSIDENCE ....................... 347

9.2.1 Types of Discontinuous Subsidence ........................................................... 347

9.2.2 Chimney Caving Mechanisms .................................................................. 349

9.3 EXAMPLES OF SURFACE SUBSIDENCE ARISING FROM BLOCK

AND PANEL CAVING ...................................................................................... 352

9.3.1 Miami Mine, Arizona, USA ..................................................................... 352

9.3.2 San Manuel Mine, Arizona, USA .............................................................. 352

9.3.3 Henderson Mine, Colorado, USA355 9.4 ANALYSIS OF CHIMNEY CAVING AND PLUG SUBSIDENCE ............................ 356

9.4.1 Limiting Equilibrium Analysis ................................. '" .............................. 356

9.4.2 Empirical Methods ................................................................................. 362 9.5 LIMITING EQUILIBRIUM ANALYSIS OF PROGRESSIVE

HANGINGW ALL CAVING ................................................................................ 364 9.6 SUBSIDENCE PREDICTION IN PRACTICE ........................................................ 369

9.6.1 General Approach .................................................................................. 369

9.6.2 Prediction of Caving Induced Subsidence at Rio Blanco and

El Teniente Mines, Chile ......................................................................... 370

CHAPTER 10

MAJOR OPERATIONAL HAZARDS

10.1 SCOPE ........................................................................................................... 375

10.2 MAJOR COLLAPSES ........................................................................................ 377

10.2.1 Terminology ......................................................................................... 377

10.2.2 Causes ................................................................................................. 377

10.2.3 Effects ................................................................................................ 380

10.2.4 Prevention and Amelioration .................................................................... 381

10.3 ROCKBURSTS ................................................................................................. 382

10.3.1 Terminology ......................................................................................... 382

10.3.2 Causes ................................................................................................. 382

10.3.3 Effects ................................................................................................. 384


10.4 MUD RUSHES ................................................................................................. 387

10.4.1 Terminology ......................................................................................... 387

10.4.2 Causes ................................................................................................. 388

10.4.3 Effects ................................................................................................. 393


xiv

10.S AIRBLASTS ................................................................................................ 396

10.S.1 Tenninology ..................................................................................... 396

10.5.2 Causes ............................................................................................. 396

10.5.3 Effects ............................................................................................. 398

IO.S.4 Prevention and Amelioration ................................................................ 398

10.6 WATER AND SLURRY INRUSHES ............................................................... 399

CHAPTER 11

RISK ANALYSIS FOR BLOCK CAVING

11.1 INTRODUCTION TO RISK ANALySIS .......................................................... 400

11.2 DEFINITIONS .............................................................................................. 40 I

11.3 PROJECT DEVELOPMENT ........................................................................... 402

11.4 RISK ANALYSIS TOOLS AND CONCEPTS .................................................... 404

11.4.1 Risk Analysis Tools ........................................................................... 404

11.4.2 Sources of Risk ................................................................................. 404

11.4.3 Uncertainty ....................................................................................... 40S

11.5 CA VERISK .................................................................................................. 406

11.5.1 Purpose ............................................................................................ 406

11.5.2 Outline of CaveRisk ........................................................................... 406

11.5.3 Topics and Focus Issues ...................................................................... 408

I1.S.4 Likelihood and Consequences .............................................................. 413

11.5.5 Risk Detennination and Risk Acceptance ............................................... 416

11.5.6 Risk Manageability ............................................................................ 417

11.5.7 Risk Presentation .................................................................... " ......... 419

11.5.8 Rules Operating in CaveRisk ............................................................... 419

11.6 CONCLUSION ............................................................................................. 421

REFERENCES ........................................................................................................ 423

APPENDIX A: GLOSSARy ..................................................................................... 463

APPENDIX B: RELATION BETWEEN CAVED COLUMN HEIGHT AND VERTICAL STRESS AT THE CAVE BASE ......................................... 471

APPENDIX C: NUMERICAL SIMULATION OF PARTICLE FLOW USING REBOP ...... 484

APPENDIX D: LIMITING EQUILIBRIUM ANALYSIS OF PROGRESSIVE HANGINGW ALL CAVING ............................................................... 501

INDEX .................................................................................................................. 509

xv

1

CHAPTER 1

INTRODUCTION 1.1 UNDERGROUND MINING METHODS

1.1.1 General Features

he underground mining of minerals involves three general sets of activities: • the development of physical access to the mineralised zone; • the extraction of the ore from the enclosing rock mass; and • the transport of the ore to processing facilities on the mine surface.

This general process requires the development of three main types of underground excavation: • permanent access and service openings or components of the mine infrastructure; • stope access and service openings, or stope development; and • ore sources or stopes through which the ore is removed from its in situ setting. The set of stopes generated during ore extraction by underground mass mining methods usually constitute the largest excavations formed during the overall mining process. This means that their zones of influence are relatively large compared to those of virtually all other mine openings (Brady and Brown 1993). The method by which the stopes are supported in order to maintain their fitness for purpose then becomes a major consideration in mining method selection and mine design. Indeed, it is usually on the basis of whether or not stopes are supported, and if so how, that underground mining methods are classified (eg Hamrin 1982). 1.1.2 Classification of Underground Mining Methods

Most systems of classifying underground mining methods are based on methods of supporting the stopes. As Rossouw and Fourie (1996) have argued, the classification of underground mining methods is not as straightforward as might be supposed. In order to overcome some of the perceived difficulties with existing systems, they proposed a three-dimensional presentation which takes into account three forms of support - natural (pillars), artificial (fill) and none

T

Chapter 1: Introduction

2

(caving). However, Roussow and Fourie’s presentation is quite complex and has not found widespread use. The essential features to be considered are the relations between the method of working, the key orebody properties defining the applicability of that method and the country rock mass properties that are essential to sustain the method (Brady and Brown 1993).

Figure 1.1 shows one version of a common approach to underground mining method classification. Not all methods of mining currently employed are shown on this diagram (eg bench stoping) but they could be added if required. The unsupported or caving methods of mining seek to induce mass failure of, and large displacements in, the country rock which will necessarily behave as a discontinuum. At the other end of the spectrum, the supported methods seek to maintain the integrity and “elastic” response of the country rock and to strictly limit its displacement.

Figure 1.1: Classification of underground mining methods (Brady and Brown 1993)

As shown in Figure 1.1, the unsupported or caving methods of mining include block (and panel) caving, sublevel caving and longwall methods. In the longwall method applied to coal mining, the mineral (coal) is extracted mechanically and the overlying strata cave under the influence of redistributed stresses and gravitational forces. The longwall methods used to mine the deep, flat dipping, tabular gold reefs in South Africa are sometimes classified as caving methods (eg Brady and Brown 1993), although the mechanism by which the overlying rock displaces to fill the void created by the extraction of the ore usually involves “elastic” displacement of the rock on the release of extremely high stresses rather than, or as well as, caving per se. In sublevel caving methods, the ore is drilled and blasted and drawn following which the surrounding waste rock caves naturally. In the block and panel caving methods with which this book is concerned, both the ore and the overlying rock cave under the influence of


3

gravity and the redistributed in situ stresses once the orebody has been undercut. In these methods, in particular, the caving and caved ore and waste rock behave as discontinuous materials. 1.2 BLOCK AND PANEL CAVING

1.2.1 Outline of the Method

Figure 1.2 illustrates the general features of the block caving method. In this method, the full orebody or an approximately equidimensional block of ore is undercut fully to initiate caving. The undercut zone is drilled and blasted progressively and some broken ore is drawn off to create a void into which initial caving of the overlying ore can take place. As more broken ore is drawn progressively following cave initiation, the cave propagates upwards through the orebody or block until the overlying rock also caves and surface subsidence occurs. The mechanisms by which caving takes place under the influence of redistributed stresses and/or gravity will be outlined in Section 1.2.2.

Figure 1.2: Example of block caving with LHD loaders, El Teniente, Chile (Hamrin 2001)


4

The broken ore is removed through the production or extraction level developed below the undercut level and connected to it by drawbells through which the ore gravitates to drawpoints on the extraction level. In most current block caving operations, the broken ore is removed from the drawpoints by Load-Haul-Dump (LHD) vehicles although some still use the more traditional gravity - based grizzly or slusher systems as discussed in Section 1.2.3. From the extraction level, the ore is transported to the haulage level and out of the mine, sometimes following underground crushing. Block caving may be used in massive orebodies which have large, regular “footprints” and either dip steeply or are of large vertical extent. It is a low cost mining method which is capable of automation to produce an underground “rock factory” (eg Tota 1997). However, it is capital intensive requiring considerable investment in infrastructure and development before production can commence. Historically, block caving was used for massive, low strength and usually low grade orebodies which produced fine fragmentation (Lewis and Clark 1964). Where mining is to be mechanised, the low strength of the rock mass can place limitations on the practicable sizes of the extraction level excavations. Furthermore, finely fragmented ore can “chimney” when drawn requiring the drawbells to be closely spaced so that undrawn “pillars” of broken ore do not form (Ward 1981). These factors place limitations on the sizes of the equipment that can be used. Accordingly, there is now a tendency for the method to be used in stronger orebodies which produce coarser fragmentation than did the traditional applications of the method. This enables more widely spaced drawpoints and larger equipment to be used. Panel caving and other variants of the general method such as inclined drawpoint caving and front caving, operate on the same principles as block caving. In panel caving, the orebody or mining block is not undercut fully initially but, rather, a panel or strip of the orebody is undercut and allowed to cave. Development, undercutting and mining of the subsequent panels then follow some distance behind the first panel as illustrated in Figure 1.3. As a result, the cave front moves across the block or orebody at a constant angle to the direction of advance of the undercut. Examples of the application of this method will be given in Section 1.3 below. Inclined drawpoint caving (Laubscher 2000, Laubscher and Esterhuizen 1994) is used when it is not possible to develop the drawpoints on one level, usually because the orebody has a well-defined inclined footwall. In this case, the drawpoints are developed at the footwall contact from the footwall on successive sublevels with the drifts being continued to serve as undercut drill drifts. In some cases such as that at the King Mine, Zimbabwe, illustrated in Figure 1.4, local geological conditions may lead to a “false footwall layout” being used in which the inclination of the plane of the drawpoints is flatter than the footwall contact (Laubscher 2000, Laubscher and Esterhuizen 1994). Front caving was developed from the overdraw system used on the two lower levels of the sublevel caving operations at the Shabanie Mine, Zimbabwe (Laubscher 2000). In recent years, front caving has been used at the Koffiefontein Mine, South Africa, and the King Mine,


5

Zimbabwe, where the method is referred to as retreating brow caving. In essence, the method involves retreating on one or more levels from an initiating slot which can be in the centre of the orebody as at Koffiefontein, or against the orebody boundary. The lower level is the production level on which so-called semi-permanent drawpoints are fully developed ahead of undercutting on the upper level. This upper level also provides initial temporary drawpoints from which the swell from each blasted ring is drawn. The undercut is retreated in stages to points above the semi-permanent drawpoints in a manner similar to that used in sublevel caving. Ideally, the method should work best with two production levels rather than one. However, this approach may be precluded on cost or other grounds, including space and layout considerations.

Figure 1.3: Mechanised panel caving, Henderson Mine, Colorado, USA (Doepken 1982)

There are many more variants of block and panel caving methods of mining than those listed above. For example, the macrotrench (or macrozanja) method developed at the El Teniente Mine, Chile, contains elements of panel, inclined drawpoint and front caving methods. Exploitation starts through a four level sublevel cave that begins from a central slot and is then retreated to both sides leaving a large trench around the initial slot. The sublevel caving is stopped in a position which leaves the upper levels and their drawpoints more advanced than the lower levels (Diaz and Tobar 2000).


6

(a)

(b)

Figure 1.4: Inclined drawpoint caving, King Mine, Zimbabwe, (a) vertical section showing extraction level layout, and (b) plan showing sublevel layout (Laubscher and

Esterhuizen 1994)


7

Caving methods of mining may be classified according to • whether or not part of the ore column is broken by blasting or other "artificial" methods; • whether or not a crown pillar is left between mining lifts; • the undercutting strategy used (see Section 1.3 and Chapter 5); and • the method of ore loading used. Figure 1.5 shows an informative classification of caving methods of mining, including sublevel caving, developed on this basis by Flores and Karzulovic (2002).

Figure 1.5: Classification of caving methods of mining

(Flores and Karzulovic 2002) Many of the larger orebodies being mined by the caving method in fact use panel caving although the more generic term block caving may sometimes be used to describe the mining method. Generally in this book, the term block caving will be used as a generic rather than as a specific term so that the discussion will usually apply to panel caving as well.


8

It will be apparent from this introductory description of block and panel caving methods, that while their capital or development costs may be relatively high, operating costs can be expected to be lower than those of other underground mining methods. It is for this reason that caving methods are attractive for the mass mining or large, lower grade orebodies. Figure 1.6 summarises the underground mining cash costs in $US per tonne at a number of block and panel caving operations in the years 1999 and 2000. (These data were compiled by Northparkes Mines, Australia, and shared with ICS sponsors).

Underground mining cash cost US$/tonne

Figure 1.6: Comparative underground mining cash costs for block and panel caving

mines in 1999 and 2000 1.2.2 Basic Caving Mechanics

It must be expected that any unsupported rock mass will cave if it is undercut to a sufficient extent. As has been noted earlier, caving occurs as a result of two major influences – gravity and the stresses induced in the crown or back of the undercut or cave. The mechanisms by which caving occurs will depend on the relationships between the induced stresses, the strength of the rock mass and the geometry and strengths of the discontinuities in the rock mass. Much accumulated experience supports the contention of Kendorski (1978) that the successful initiation and propagation of caving requires the presence of a well-developed, low-dip discontinuity set. The structure most favourable for caving has been found to be one in which a low-dip discontinuity set is augmented by two steeply dipping sets which provide conditions suitable for the vertical displacement of blocks of rock (eg Mahtab et al 1973).

Min

e


9

If the compressive tangential stresses induced in the crown of the undercut or cave are low, or tend to be tensile, blocks of rock may become free to fall under the influence of gravity or to slide on inclined discontinuities. These conditions may occur when the horizontal in situ stresses are low or where boundary slots or previous mining have relieved the stresses or redistributed them away from the block or panel being mined. Even under these circumstances, it is sometimes possible for a self-supporting arch to develop in the crown of the cave, especially if an appropriate draw control strategy is not in place. Some of the mechanisms by which caving and arching may occur under these low lateral stress conditions are illustrated by the simple and idealised distinct element simulation shown in Figure 1.7. Each pair of drawings in Figure 1.7 represent the geometric configuration of the blocks and the interblock contact force vectors at different stages in the progressive caving of the mass. Note that two apparently independent arches form where high levels of inter-block force traverse the mass. The upper arch is stronger and is sustained longer than the lower arch but both fail eventually by slip at the rigid abutments. At the other extreme, when the induced tangential stresses are high compared with the compressive and shear strengths of the rock mass and the shear strengths of the discontinuities, failure may occur at or near the boundary of the rock mass and blocks or slabs of rock may become free to fall under the influence of gravity. Under these circumstances, the dominant mechanisms of failure are brittle fracture of the intact rock and slip on discontinuities, especially those that are flat dipping (eg Heslop and Laubscher 1981). This form of caving is sometimes referred to as stress caving. Duplancic and Brady (1999) used a seismic monitoring system to study the early stages of caving of Lift 1 at Northparkes Mines’ E26 block cave, New South Wales, Australia. From the data collected and analysed, they developed the conceptual model of caving for this case shown in Figure 1.8. The model contains five regions described by Duplancic and Brady (1999) in the following terms: 1. Caved zone. This region consists of rock blocks which have fallen from the cave back.

Material in the caved zone provides support to the walls of the cave. 2. Air gap. During continuous caving, the height of the air gap formed is a function of the

extraction rate of the material from the caved zone.


10

Figure 1.7: Idealised distinct element simulation of block caving (after Voegele et al

1978) 3. Zone of discontinuous deformation. This region no longer provides support to the

overlying rock mass. Large-scale displacements of rock occur in this area, which is where disintegration of the rock mass occurs. No seismicity is recorded from within this region. The zone was estimated to extend 15 m from the boundary of the cave crown.


11

4. Seismogenic zone. An active seismic front occurs due to slip on joints and brittle failure of rock. This behaviour is due to changing stress conditions caused by the advancing undercut and progress of the cave.

5. Surrounding rock mass. Elastic deformation occurs in the rock mass ahead of the seismic

front and surrounding the cave.

Figure 1.8: Conceptual model of caving (Duplancic and Brady 1999)

Duplancic and Brady’s observations at Northparkes confirm the previous general finding that for boundary collapse to occur, a flat lying discontinuity set is required to act as a release mechanism. A third general case must be considered. If the horizontal in situ stresses and the tangential stresses induced in the crown of the undercut or cave are high enough to develop clamping forces which inhibit gravity-induced caving, but are not high compared with the compressive strength of the rock mass, caving may be inhibited and a stable arch may develop. Under these circumstances, some form of cave induction may be required to weaken the rock mass, relieve the tangential stresses or induce slip on discontinuities (eg Kendrick 1970, van As and Jeffrey 2000). A different mechanism from those discussed so far is involved in subsidence caving in which a large mass of rock subsides rapidly as a result of shear failure on the vertical or near-vertical boundaries of a block. For this to occur the normal (horizontal) stresses developed on the vertical boundaries of the block, or the shear strength of the interface, must be so low that the total shear resistance developed is unable to resist the vertical force due to the weight of the

Direction of advancing undercut

Caved zone

Air gapZone of loosening

Pseudo-continuous domain

Seismogenic zone


12

block. For such a failure to have catastrophic consequences, there would need to exist a large mined-out void into which the caving mass could fall. This circumstance would not arise in a block or panel cave if the draw control strategy used did not allow a significant air gap to develop below the cave back. Once continuous caving has been initiated, the rate of production from the block or panel will depend on the rate at which the cave propagates following draw and the creation of a small air void into which caved material may fall. In practice this rate of caving will depend on the rate of undercutting, the quality of the rock mass and the magnitude of the induced stresses. As will be discussed in Chapter 5, the direction of undercutting with respect to the in situ stress orientation is also important. Estimated caving rates for a number of mines are summarised Table 1.1. It should be emphasised that these caving rates are estimated. They are notoriously difficult to measure. Furthermore, they may vary through the life of a cave. For example, as the height of the cave and of the column of broken ore increases, the induced stresses in the cave back may change, as may the structure and rock mass strength of the orebody.

Table 1.1: Estimated caving rates

Operation Estimated Caving Rate (mm per day)

CODELCO El Teniente Sub 6 panel cave 200 to 300

CODELCO Esmeralda panel cave 170 to 200

De Beers Koffiefontein (TKB Kimberlite) 200 to 400

De Beers Premier Mine (TKB Kimberlite) 100 to 1200

De Beers Premier Mine (HYB Kimberlite) 60 to 250

Henderson Mine 270

Northparkes E26 Lift 1 block cave 110 to 380 (pre inducement)

Under steady-state production conditions, the average rate of draw will be a function of the rate of natural caving and the bulking factor of the caved ore. In currently operating block and panel caving mines, rates of draw vary up to about 700 mm/day with the mean in the range 200 to 250 mm/day (Flores and Karzulovic 2002b). Drawing of the difference between the in situ and caved volumes following each caving episode will ensure that cave propagation is controlled and an excessive air gap does not develop. Of course, for this controlled caving to occur, a small air gap must be created by drawing the caved ore. The major consequence of allowing an excessive air gap to develop is the danger of massive rock falls and the associated air blasts to be discussed in Chapter 10.


13

1.2.3 History of Block Caving

The precursor of the modern block caving method of mining was developed in the iron ore mines of the Menominee Ranges, Michigan, USA, in the late nineteenth century. The Pewabic Mine was the first to use a form of block caving from which other methods developed (Peele 1941). In the Pewabic method, blocks of ore approximately 60-75 m long, 30-40 m high and the full width of the deposit (60 m) were caved in one operation. An unusual feature of the method from a modern perspective is the fact that the ore was handled by shovelling in drifts driven and kept open within the caved mass (Peele 1941). Variations of the Pewabic method were soon developed at other iron ore mines in Michigan and, from the early part of the twentieth century, in the copper mines in western USA. Before evolving to the use of full block caving, many mines initially used combined methods involving, for example, shrinkage stoping and caving methods for the subsequent mining of the pillars between the primary stopes. Peele (1941) gives several examples of these combined methods. A good example of the early application of a full block caving method is provided by the Miami Copper Company’s mine in Arizona. Descriptions provided by Peele (1941) and Lewis and Clark (1964), and the diagram shown in Figure 1.9, are based on a paper by McLennan (1930). (Note that the original dimensions in feet have been retained in Figure 1.9. They have been approximated in metres in the text). The flat lying orebody of considerable lateral extent varied in thickness to more than 60 m and was over- and under-lain by waste. Early mining was by top-slicing but this was replaced by shrinkage stoping with sublevel caving of the pillars and, in the 1920s, by block caving. Initially, caving practice involved undercutting and caving the orebody across its entire width of 150-200 m, starting at one end and retreating along the length of the orebody. This approach was unsuccessful and later practice was to cave and draw alternate 45 m wide panels across the entire orebody. When the waste rock had settled into the original panels and compacted, the pillar panels were caved. This method was satisfactory for moderate thicknesses of ore averaging 60 m but was modified to true block caving where thicknesses were 90 m or more. The original caving blocks of the thicker ore were 45 by 90 m in plan as shown in Figure 1.9. Experience showed that 45 m square blocks gave better results and this block size became the standard. The order of mining was such that a block being mined would be entirely surrounded by either solid ore or mined-out blocks in which the capping had settled until it was quite compact (Lewis and Clark 1964). As shown in Figure 1.9, the Miami mine used a gravity system of ore transfer to the haulage level incorporating grizzlies. By the 1920s and 30s, block caving methods were being used in a wide range of mines exploiting massive, weak orebodies. During this period, the method was introduced, for example, at the King Mine which mined asbestos in Quebec, Canada, the Climax Mine mining molybdenum in Colorado, USA, and the copper mines in Chile (Peele 1941).


14

Figure 1.9: Block caving, Miami Mine, Arizona, USA (Lewis and Clark 1964)

South Africa’s first kimberlite diamond pipe was discovered at Kimberley in 1870 a few years after the discovery of alluvial diamonds at Hopetown. The early mining of this and other kimberlite pipes was by surface methods. An initial attempt at underground mining from 1884 was unsuccessful, largely because of the uncooperative attitudes of the many claim holders (Owen and Guest 1994). The consolidation of the pre-existing companies into De Beers Consolidated Mines in 1888 provided the impetus for the successful introduction of underground mining at Kimberley from 1890. A complex mining method known as chambering was used until it was replaced progressively from the late 1950s to the early 1970s (Hartley 1981). Peele (1941) describes this method as a combination of shrinkage stoping and sublevel caving. A visit to North America in the early 1950s convinced De Beers mine managers of the potential of the block caving method for mining the diamondiferous pipes (Gallagher and Loftus 1960). After an experimental block cave had been mined successfully at the Bultfontein Mine it was decided to introduce block caving on a large scale at several of the De Beers mines. The first mine to change fully from chambering to block caving was Jagersfontein where the transition was completed in 1958 (Gallagher and Loftus 1960).


15

As Owen and Guest (1994) describe, a number of different mining methods including sublevel caving, vertical crater retreat and benching have been used at the various De Beers mines but variants of block caving remain the main group of methods used currently. A feature of the evolution of the current mining methods is that not all methods introduced were successful initially and mining plans often had to be revised. Mechanised panel caving was introduced at the Premier Mine from 1990 (Bartlett 1992). Mining operations at Premier are summarised in Section 1.3.3 below. As has been indicated above, ore has been drawn throughout much of the history of block and panel cave mining by gravity or slusher methods. In 1981, Pillar (1981) listed a wide range of well-known caving mines as then using the slusher method. The availability of LHD equipment from the 1960s has provided the potential for the introduction of mechanised and trackless cave mining, especially for the more coarsely fragmenting and stronger ores in which the necessary large extraction level openings can be developed and maintained. From the 1980s, many of the major caving mines have introduced mechanised methods (see Table 1.2) although gravity systems are still used, particularly in the weaker ores. Examples of modern mechanised methods are given in Section 1.3.

Table 1.2: Examples of current caving operations

Mine Country Type Ore Type Annual

Tonnage

Northparkes

E26 Lift 1

Australia Block Copper-Gold 4 Mt/y

Freeport IOZ Indonesia Block; LHD Copper-Gold 7 Mt/y

Palabora South Africa Block; LHD Copper 10 Mt/y

CODELCO

El Teniente

Division

Chile Panel; LHD Copper 35 Mt/y

CODELCO

Andina Division

Chile Panel; LHD

Grizzly

Copper 16 Mt/y

CODELCO

Salvador

Chile Panel; LHD Copper 2.5 Mt/y

Premier Mines South Africa Panel; LHD Diamonds 3 Mt/y

Henderson USA Panel; LHD Molybdenum 6 Mt/y

Philex Philippines Grizzly; LHD Copper

Shabanie Zimbabwe Retreating brow Chrysotile

asbestos

Tongkuangyu China Block; LHD Copper 4 Mt/y

San Manuel USA Block; Grizzly Copper


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1.3 BLOCK AND PANEL CAVING OPERATIONS

1.3.1 Overview

Because they are mass mining methods having low production costs, block and panel caving methods are currently important sources of mineral production on a world scale. Laubscher (1994) estimated that, at that time, these methods accounted for a total ore production of approximately 370,000 tonnes per day. Because of the high productivity of caving methods and the potential that they offer for mechanisation and reduced labour costs, there is a current tendency in the industrially advanced nations, in particular, to apply block caving to stronger orebodies than those to which the method has been applied in the past. This brings with it particular challenges in predicting caveability, an issue to be introduced in Section 1.4.1 and explored in detail in Chapter 3. Block caving is also being considered for the underground mining of some major orebodies previously mined by large open pits. Significant examples of this development are the Bingham Canyon mine in Utah, USA (Carter and Russell 2000) and the Palabora mine in the Republic of South Africa (Calder et al 2000). Table 1.2 lists many of the world’s major block and panel caving operations in the year 2000. The list is intended to be indicative rather than exhaustive. Because of a paucity of accessible data, cave mining operations in some countries such as China and the countries of the former USSR are not well represented in Table 1.2. Three of the major mines listed in Table 1.2 – El Teniente, Premier and Henderson - are described in the subsequent sub-sections as important but differing examples of modern cave mining practice. These descriptions may use some terms that have not been defined so far and which may not be familiar to some readers. In these cases, reference should be made to that part of the book in which the term is introduced in more detail or to the glossary of terms presented at Appendix A. Recent accounts of a number of other block and panel caving operations are given by Hustrulid and Bullock (2001). 1.3.2 El Teniente Mine, Chile

El Teniente is a division of CODELCO-Chile, Chile’s Government-owned copper mining company. The mine is located 130 km southeast of Santiago in the foothills of the Andes mountains. With production of about 100,000 tonnes per day, El Teniente is the world’s largest block or panel caving mine. The large copper porphyry orebody which was discovered in 1760, reaches a depth of more than 1 km below surface and almost completely surrounds a roughly 800-1000 m diameter circular breccia pipe known as the Braden Pipe. The orebody is approximately triangular in plan and has a radial extent of between 400 and 800 m from the perimeter of the pipe. The major items of mine infrastructure are conveniently located within the breccia pipe. Figure 1.10 shows a schematic view of the several levels in the El Teniente mine. The early block caving production level, Teniente 1, is located some 650 m above the main access level, Teniente 8, which is at an elevation of 1983 m.


17

After simple and irregular open pit mining of the near-surface secondary enriched ore, underground mining began in 1906 using overhand stoping and room and pillar methods. From 1940, a traditional gravity flow block caving method was successfully introduced (Kvapil et al 1989). The secondary ore was weak, fragmented readily and was well suited to mining by block caving. However, the quantity of the secondary ore decreases with depth so that from about 1982, mining has been increasingly in the lower grade and stronger primary ore. The primary ore is an andesite which contains pockets of even stronger diorite and dacite. Mechanised panel caving using LHDs was introduced on the Teniente 4 level (at 2347 m) in 1982 (Alvial 1992).

Figure 1.10: Schematic representation of levels, El Teniente Mine (Kvapil et al 1989)

A major factor influencing the mining of the El Teniente orebody, particularly at increasing depth in the primary ore, has been the existence of extremely high lateral in situ stresses associated with the nearby subduction zone in which the Pacific plate is thrust under the edge of


18

the South American plate. Rock bursts were first experienced on the Teniente 4 level in 1976 (Alvial 1992) and have been a continuing problem since mechanised mining began in the stronger and stiffer primary ore. El Teniente now uses mainly mechanised panel caving but there is still some production from areas which use other forms of caving (Jofre et al 2000). Accounts of successive stages in the evolution of cave mining methods at El Teniente are given by Ovalle (1981), Kvapil et al (1989), Alvial (1992), Moyano and Vienne (1993), Rojas et al (2000b), Jofre et al (2000) and Rojas et al (2001). The following description of recent mining at El Teniente draws on the accounts of Moyano and Vienne (1993) and Jofre et al (2000). As has been noted, mechanised panel caving of the primary ore was introduced on the El Teniente 4 level in 1982. A conventional or post-undercutting method was used with the undercut being mined after the development of the extraction level below and the excavation of the drawpoints. (Details of the meaning and implications of post-undercutting and related terms are given in Chapter 5). In the traditional El Teniente panel caving method, the undercut level was located 18 m above the extraction level and the undercut drifts were 3.6 by 3.6 m on 30 m centres. Over time the height of the undercut has been progressively reduced from 16.6 m in 1987 to 10.6 m in 1998 (Jofre et al 2000).

The extraction level layout developed at El Teniente is illustrated in Figure 1.11 for the Teniente Sub 6 level. The conventional or post-undercutting sequence used until recently exposed the pre-constructed extraction level to high levels of abutment stress as the undercut advanced overhead. This resulted in damage to the extraction level pillars and, in some instances, in severe rock bursts associated with the high horizontal stresses and local geological features (Moyano and Vienne 1993, Rojas et al 2000a). A series of major rock bursts on the Teniente Sub 6 level starting in January 1990 soon after it came into production, caused production to be stopped temporarily and the approach to mining revised. The measures introduced from 1994 to successfully control the incidence and severity of rock bursts on the Sub 6 level included reducing the height of the column of intact rock above the undercut level, ensuring an even spatial and temporal rate of extraction, mining at a slow rate especially initially, using only remotely controlled production equipment from 1995 and using the results of seismic and other monitoring to guide production planning. In 2000, the production rate of Teniente Sub 6 was about 10,000 tonnes per day and the undercutting rate was about 12,000 m2 per year (Rojas et al 2000a).


19

Notes: 1. Haulage drift and drawpoint drifts section wide 4.00 m : high 3.60 m 2. Distance between consecutive dump points located in the same haulage drift will be that corresponding to 6 drawpoint drifts (103.92 m)

Figure 1.11: Extraction level layout, Sub 6 level, El Teniente Mine (Moyano and

Vienne 1993) From 1992 a pre-undercut panel caving method was tested at El Teniente. In 1997 the method was brought into production on the new Esmeralda section on the Teniente Sub 5 level. In this method, the undercut level is developed and then blasted in advance of the development of the extraction level and formation of the drawbells. Thus, all extraction level development and construction takes place in a de-stressed zone below the mined undercut. The undercut drifts are 3.6 by 3.6 m on 15 m centres. Initial development on the extraction level is kept 22.5 m behind the undercut, and full construction on the extraction level occurs 45 to 60 m behind the undercut (see Figure 1.12). Experience and the results of extensive monitoring of the condition of the pillars and the extraction level installations, clearly demonstrate the effectiveness of this approach (Rojas et al 2000b, Jofre et al 2000). In early 2000, the average production from the Esmeralda section was 12,500 tonnes per day with a planned peak of 15,000 tonnes per day. The pre-undercut method was also being used on the Teniente 3 Isla section.


20

Figure 1.12: Comparison of conventional panel caving and pre-undercut activity sequences, El Teniente Mine (after Rojas et al 2000b)

1.3.3 Premier Diamond Mine, South Africa

The Premier Diamond Mine is situated 45 km to the east-north-east of Pretoria, Republic of South Africa. It exploits the Premier Pipe, the largest of South Africa’s kimberlite pipes having a surface area of 32 hectares. The Premier Pipe is unique geologically in that it is intersected by a 75 m thick, shallow-dipping gabbro sill which cuts through the pipe at depths of between 380 and 510 m below surface (Bartlett 1992). Figure 1.13 shows a diagrammatic vertical section and plan of the geology and of the recent and planned mining blocks. Open pit mining commenced at Premier in 1902. With the increasing depth of mining, underground mining started in 1948, initially by open benching. Following the successful implementation of the block caving method at other De Beers mines at Kimberley and Jagersfontein as outlined in Section 1.2.3, block caving was introduced at Premier in the early 1970s. Four separate caves using a grizzly system feeding slusher drifts were operated above the gabbro sill on the western side of the mine. A total of 85 million tonnes of ore was produced from these caves (Bartlett 1992).

PRE UNDERCUT


21

Figure 1.13: Diagrammatic section and plan, Premier Diamond Mine (Bartlett and Croll 2000)

Mining below the sill started in 1979 using an open stoping method. This method was unsuccessful and was replaced by block caving. Cave mining using LHDs to transport the ore from the drawpoints to ore passes started in 1990 in the BA5 mining block on the 630 m level on the western side of the mine (see Figure 1.13). A second cave was established in 1996 in the


22

BB1E block at a depth of 732 m in weaker ore on the eastern side of the pipe. Possible future mining of the C-cut (see Figure 1.13) would involve the exploitation of 170 million tonnes of ore by caving methods with a production level 1000 m or more below surface (Bartlett and Croll 2000). The remainder of this Section will discuss recent and planned caving operations at Premier as reported by Bartlett and Croll (2000). Accounts of the earlier stages of cave mining at Premier are given by Owen (1981), Bartlett (1992) and Owen and Guest (1994).

Mining of the BA5 block was planned as a panel retreat caving operation with a post-undercut mining sequence. With this sequence, mining of the undercut would occur after the extraction level development had been completed. Undercut drifts were developed directly above the extraction drifts on 30 m centres. The block was undercut by drilling and blasting 120 m long by 30 m wide and 20 m high slots at right angles to the directions of the undercut and extraction level drifts. As undercut rings were blasted, broken ore dropped directly into the pre-developed drawbells. As the undercut area approached that required for the onset of caving, the levels of stress on both the undercut and extraction levels increased. On the undercut level, it became increasingly difficult to drill, charge and blast the long holes and the rate of undercutting slowed. On the extraction level, shotcrete linings were extensively damaged by the abutment stresses associated with the now slow moving undercut. Footwall heave was widespread, damaging concreted roadways and disrupting production. When continuous caving was established, the stress levels stabilised but an extensive program of extraction level support and rehabilitation was required to ensure safety and uninterrupted production. The rate of caving was slower than planned and a large and potentially dangerous air gap developed. The caving process was compromised by the presence of the overlying, strong gabbro sill. The difficulties continued as the panel retreated to the east. It was then decided to adopt an advance undercut mining sequence in which only the production drift and drawpoint breakaways were developed and partly supported ahead of the mining of the undercut overhead. The height of the undercut was also reduced. As the zone below the advancing undercut became de-stressed, development of the extraction level, including the concreting of roadways and the application of a shotcrete lining, was completed. The adoption of these measures allowed undercutting to proceed at the planned rate, support and rehabilitation requirements to be reduced and production targets to be met. The lessons learned from the BA5 block were applied in planning the mining of the BB1E block (see Figure 1.13) as a panel retreat cave with an advance undercut mining sequence. Table 1.3 shows a comparison of the design and operational parameters of the BA5 and BB1E blocks. As Bartlett and Croll (2000) explain, some difficulties were experienced in the BB1E block, particularly in the early stages of undercutting and cave establishment. However, once the cave was fully established and the undercut could be advanced at the planned rate, satisfactory results were achieved. Experience with the BB1E block shows that successful


23

caving at this depth requires detailed planning, the timely availability of resources, and careful implementation and control of the entire mining process. Maintaining the planned schedule to avoid the compaction of broken ore provided a particular challenge in this case. Table 1.3: Design and operational parameters, BA5 and BB1E mining blocks, Premier

Mine, South Africa (Bartlett and Croll 2000) Parameter BA5 Mining Block BBIE Mining Block Column height 80 - 140 metres 148 - 163 metres Rock mass rating 45 - 65 45 - 55 Hydraulic radius 30 25

Mining sequence Post and advance undercut Advance undercut Rate of undercutting 900 square metres per month 1100 square metres per month Tons in mining block 42 million tons 23 million tons Tons per drawpoint 50 000 to 120 000 tons 100 000 to 200 000 tons

Drawpoint spacing 15 x 15 metres 15 x 18 metres Distance across major apex 22.6 metres 23 metres Average rate of draw 180 mm per day (109 tons) 165 mm per day (120 tons)

Initial fragmentation 30 % >2 cubic metres 30 % >2 cubic metres Fragmentation after 20 % drawn

12 %>2 cubic metres 7 % >2 cubic metres

Drawpoint support Cable anchors, rockbolts, mesh tendon straps and shotcrete

Cable anchors, rockbolts, mesh tendon straps and shotcrete

Brow wear 0 to 2 metres wear after 50 000 tons drawn

1 to 3 metres wear after 50 000 tons drawn

Tunnel size 4 x 4.2 metres 4 x 4.2 metres

Lhd type Diesel and electric 5 and 7 yard Toros Diesel and electric 5 and 7 yard Toros

Tons per LHD per hour 118 tons 131 tons

LHD average tramming distance

154 metres 134 metres

Hangup frequency 30 % of drawpoints per shift 25 % of drawpoints per shift Fragmentation Initial 20% drawn Secondary blasting

30 % >2 cubic metres 12 % >2 cubic metres

40 grams per ton

30 % >2 cubic metres 10 % >2 cubic metres

30 grams per ton

The C-cut mining block is being planned with a pre-undercut mining sequence to avoid extensive and possibly unsustainable damage to the extraction level at a depth of more than 1000 m. As a consequence, the mining sequence in the existing BB1E mining block is being changed to gain operational experience with this method in which the undercut is fully developed before the extraction level. The height of the cave is planned to be between 350 and 450 m compared with a maximum of 164 m in BB1E (see Table 1.3). New mine infrastructure is being installed to support mining of the C-cut, including two new shafts from surface and a new processing plant. The mine is being designed to extract and process 9 million tonnes of ore annually by 2008 (Bartlett and Croll 2000).


24

1.3.4 Henderson Mine, Colorado, USA

The Henderson mine is located 80 km to the west of Denver, Colorado, USA, and 3170 m above sea level on the eastern side of the Continental Divide. The top of the molybdenite orebody lies more than 1000 m below the peak of Red Mountain and the lowest excavation is at a depth of 1600 m making Henderson one of the deepest caving operations in the world. The deposit is elliptical in plan with axes of 670 and 910 m. As is illustrated in Figure 1.14, the ore is transported by conveyor from a crusher complex on the 7065 level at an elevation of 2153 m to the mill site 25 km away on the western side of the Continental Divide. This discussion of the Henderson Mine is drawn from the paper by Rech et al (2000). Earlier accounts of aspects of the operation are provided by Brumleve and Maier (1981), Doepken (1982) and Rech et al (1992).

Figure 1.14: General section, Henderson Mine (after Rech et al 2000)

Henderson commenced operation in 1976 as a mechanised panel cave with rail haulage from the 7500 level at an elevation of 2286 m. Figure 1.15 illustrates the original panel caving method. Approximately 90 million tonnes of ore were produced from the 8100 level (at 2469 m) from 1976 to 1991. In 1992, the 7700 level (at 2347 m) was brought into production and by the year 2000 had produced more than 45 million tonnes of ore. A particular feature of the operation of the Henderson mine is that during its life, the world molybdenum market has experienced several phases of over- and under-supply with the result that the mine’s production rates have had to vary accordingly. This has required flexible planning and operation of the mine.


25

The next production level to come into operation will be the 7225 level located 145 m below the 7700 level at an elevation of 2202 m. As is shown in Figure 1.14, the eastern section of the orebody has not been exploited from either the 8100 or the 7700 levels. Columns of ore up to 244 m high will become accessible on the eastern side from the 7225 level (Rech et al 2000).

Figure 1.15 shows a typical isometric section of the recent configuration of the operating section of the mine on the 7700 or 2347 m level. The undercut level at 2364 m is developed with 3.7 by 3.7 m drifts on 24.4 m centres. Future undercut drift spacing will be 30.5 m centres. From the undercut drifts, rings of both short and long holes are drilled on 2 m centres to mine the undercut and the drawbells simultaneously. Panels are 8 to 12 production drifts (244 to 366 m) in width.

Figure 1.15: General isometric view, mechanised panel caving, Henderson Mine (Rech et al 2000)


26

Figure 1.16 shows a plan view of the drawpoint and production drift layout used on the extraction level located 17 m below the undercut level. The entry angles of 56o are the sharpest that can be used with the current 7 m3 LHDs. Drawpoints are concrete lined and are fitted with steel wear plates to protect the openings from degradation over their production lives. The roadways and drawpoint floors are lined with a 300 mm thickness of concrete which permits effective clean-up and reduces tyre wear.

There is a ventilation level at 2333 m through which intake and exhaust air is transported on separate horizons. The truck haulage level at 2153 m consists of 6 by 6 m haulage drifts that provide access of the 72.6 tonne side dumping trucks to centre loading chutes at the bottoms of the ore passes. The trucks transport the ore to the gyratory crusher dump on this level.

20.57 m x 30.48 m Drawpoint spacing

Figure 1.16: Extraction level plan, Henderson Mine (Rech et al 2000)

The sequence in which the development of this series of openings takes place must be well planned and coordinated. Figure 1.17 shows the general two-year sequence in the development of a panel at Henderson.


27

Figure 1.17: Development sequence, Henderson Mine (Rech et al 2000)

1.4 RISK IN CAVE MINING

1.4.1 Risk Factors

Decisions to exploit a particular orebody by block or panel caving methods, the design of caving mines, and the mining operations themselves, involve risks of a number of types. Some of these risks have been referred to in the descriptions of cave mining operations given in Section 1.3 above. Detailed analyses of some of the major risk factors are given in subsequent chapters. Accounts of the risks associated with cave mining have been given recently by Heslop (2000) and Summers (2000a & b). The following is an indicative but not exhaustive list of some of the risk factors requiring consideration at various stages in a cave mining project: • adequacy of the geological data used in making estimates of the structure, shape, size and

grade of the orebody; • adequacy of the geotechnical data available about the orebody and country rock masses

including major structures, discontinuities, rock properties, in situ stresses and groundwater hydrology. These data are used in making assessments of caveability, cave initiation and propagation, fragmentation, caving performance, excavation stability and dilution;


28

• caveability assessment usually involving a prediction of the hydraulic radius

(area/perimeter) of the undercut at which caving will initiate for a rock mass having given or estimated geotechnical characteristics;

• cave propagation which is the ability of the cave to continue to propagate once caving has

been initiated. Cave propagation depends on a number of factors including the undercut design, the rate of undercutting, the stresses induced on the boundaries and above the cave, the orebody structure and its geotechnical characteristics, and the draw control strategy employed. Because of the capital intensive, non-selective and relatively inflexible nature of caving methods of mining, the inability to initiate or sustain caving is one of the greatest risks faced in cave mining;

• the degree of fragmentation of the ore occurring as a result of the caving process. This

factor influences drawpoint spacing and design, equipment selection and performance, the occurrence of “hangups” and the need for secondary breakage in the drawpoints, the need for underground crushers and the productivity of the cave;

• caving performance reflects the achievement or otherwise of the planned rate of cave

propagation, rate of production, degree of fragmentation, ore grades and recovery; • excavation stability refers to the stability over the design life and the need for support or

reinforcement of mine excavations including undercut drifts, extraction level excavations, drawbells and items of mine infrastructure. As has been illustrated by the examples given in Section 1.3, excavation stability can depend not only on the geotechnical properties of the rock masses involved and the in situ stress field, but also on factors such as the three-dimensional mine layout, the relative timings of certain development and mining activities and the rate of undercutting;

• major operational hazards including major excavation collapses, mud rushes, rock bursts,

air blasts, and water and slurry inflows; • environmental risks broadly defined involving issues such as the mine’s influence on

surface water and groundwater, the treatment and disposal of mine wastes, influence on flora and fauna habitats, surface subsidence effects, land rights and archaeological issues and other areas of community concern; and

• risks to profitability arising from factors such as changes to cost structures, industrial

relations, variations in metal prices and currency values and local political instability.


29

1.4.2 Introduction to Risk Assessment

Techniques known as risk analysis, risk assessment and risk management are now applied to a wide range of engineering and other undertakings. In the present context, our concern is with the assessment and management of the risks associated with the adoption and operation of a particular mining method. A useful generalised definition of risk assessment is that given by the UK Engineering Council: “Risk assessment is a structured process which identifies both the likelihood and extent of adverse consequences arising from a given activity.” Engineering decisions of the type being considered here are subject to a number of uncertainties, the manifestation of which can result in the failure of a project to meet its objectives in full or in part. These uncertainties can be considered to be of two general types: what we know we don’t know, or parameter uncertainty; and what we don’t know we don’t know, or conceptual uncertainty. Parameter uncertainty is the easier of these two types of uncertainty to account for in engineering procedures. Use of the long established concept of a factor of safety is a commonly used method of attempting to account for parameter uncertainty. Probabilistic methods are also used as an alternative approach to addressing the same issue, particularly in geotechnical engineering (eg Christian et al 1994, Pine 1992). Conceptual uncertainty or uncertainty about how particular sets of conditions will develop and their eventual outcomes, is usually of greater concern and more difficult to address. A risk assessment and management approach seeks to understand the sources of risk associated with a given project or design, to evaluate their consequences and to put in place procedures to manage those risks. Implementing risk assessment and management processes is especially important in the early stages of a potential cave mining project when critical decisions about the adoption of a particular mining method, layout and excavation sequence are being made. In this book, and particularly in Chapter 11, a series of definitions associated with the assessment of risk will be used. Some of these definitions will be introduced here. Hazard - a potential occurrence or condition that could lead to injury, delay, economic loss or damage to the environment. Risk – the product of the probability of occurrence of a hazard and the magnitude of the consequences of the occurrence.


30

Risk analysis – a structured process that identifies both the likelihood and the consequences of the hazards arising from a given activity or facility. Risk evaluation – the appraisal of the significance of a given quantitative (or, when acceptable, qualitative) measure of risk. Risk assessment – comparison of the results of a risk analysis with risk acceptance criteria or other decision parameters. Risk management – the process by which decisions are made to accept known risks or the implementation of actions to reduce unacceptable risks to acceptable levels. The application of these concepts and processes to caving methods of mining will be discussed in Chapter 11. Because the emphasis of this book is on the geomechanics of cave mining, five principal forms of risk will be considered – caveability, fragmentation, caving performance, excavation stability and major operational hazards. Each of these issues has been raised in Section 1.4.1 and will be discussed in detail in subsequent chapters. 1.5 SCOPE AND CONTENTS OF THIS BOOK

This book is intended to provide a digest of the state-of-the-art of block and panel caving from a geomechanics perspective. It reflects the outcomes of the International Caving Study (ICS) Stage I, including the Block Cave Manual (Laubscher 2000) but also contains chapters on several topics that were not part of the ICS research program. Much of the information presented on current caving operations and key caving issues is drawn from papers presented at the international mass mining conference, MassMin 2000, held in Brisbane, Australia, in late 2000 (Chitombo 2000). The subsequent chapters deal with the following topics: Chapter 2 – Rock mass characterisation reviews the needs for and methods of collecting data for use in characterising rock masses for cave mine engineering purposes. The chapter discusses new methods of archiving, correcting, processing and modelling the data developed as part of the ICS. Chapter 3 – Caveability assessment addresses a topic that is vitally important in the study and design of any potential block or panel caving operation. It reviews the available empirical and numerical methods of predicting caveability with an emphasis on the extended Mathews method developed as part of the ICS.


31

Chapter 4 – Fragmentation assessment reviews the factors influencing the fragmentation produced by caving, methods of fragmentation measurement and the available methods of predicting in situ, primary and secondary fragmentation. Chapter 5 – Cave initiation by undercutting discusses the factors influencing undercut design and performance, and the undercutting strategies and undercut designs used in practice. It reports a parametric study of the influence of undercut strategy on the stresses induced in the undercut and extraction levels and the associated support and reinforcement requirements. Chapter 6 – Extraction level design is concerned with the layout and design (including the support and reinforcement) of production and drawpoint drifts, drawpoints, drawbells, pillars and ore handling facilities on extraction or production levels. Chapter 7 –Draw control discusses the importance of draw control in caving mines, the factors influencing the flow of broken ore, the numerical modelling of particle flow, and draw control practice including computer based techniques. Chapter 8 – Geotechnical monitoring considers the nature and purposes of geotechnical monitoring systems and their application to monitoring extraction level and infrastructure performance, surface subsidence and ground movements in caving mines. Chapter 9 – Surface subsidence discusses the mechanisms of caving to surface that may be associated with caving operations and presents methods of analysis of plug subsidence, chimney caving and progressive hangingwall caving. Chapter 10 – Major operational hazards associated with block and panel caving are taken to include major collapses, rock bursts, mud rushes, air blasts and water and slurry inrushes. The causes, effects and means of prevention or amelioration of the effects of these hazards are discussed. Chapter 11 – Risk assessment addresses the increasingly important topic of risk assessment and its application to cave mine engineering. Concepts are introduced, terms defined, the available tools outlined and their application to cave mining described through an account of the tool CaveRisk developed as part of the ICS.

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CHAPTER 2

ROCK MASS CHARACTERISATION 2.1 DEFINING THE MINING ENVIRONMENT

s was noted in Chapter 1, in recent years there has been renewed interest internationally in the mining of massive, often lower grade, orebodies by caving methods. A feature of block and panel caving methods of mining is that, while they

have low operating costs, they require high levels of capital investment in infrastructure and development before production can commence. A second feature of these methods is that the mine construction and development required are not readily or economically adaptable to other methods of mining if, for some reason, the chosen mining method proves to be unsuccessful. Accordingly, it is especially important that, when these methods of mining are being considered, the mining environment, especially the geotechnical environment, is understood sufficiently well to permit critical decisions to be made reliably in the pre-feasibility and feasibility study stages of a project. Not to do so invites disaster. As Laubscher (1993) has suggested, the mining environment as it is being identified here must be defined, or re-defined, in a number of circumstances: • for new mining projects on greenfield sites; • for planning the mining of new mining blocks or orebodies in current operations at

established mines; • where a change from open pit to underground mass mining is being considered; and • where difficulties encountered in operations require a review of the current mining method,

planning parameters, layout or detailed mine design. Examples of each of these circumstances have been encountered in cave mining projects and operations in recent years. In each of these circumstances, data of the types being considered here may be required for the following major purposes:

A

Chapter 2: Rock Mass Characterisation

33

• mining method selection including caveability studies; • detailed design of mining excavations including their sizes, shapes and requirements for

support and reinforcement; • fragmentation studies which influence issues such as drawpoint spacing and design and

equipment selection (including crushers); • production or extraction level layout and detailed design including support and

reinforcement requirements; • mine infrastructure location and design; • impacts of mining on the surface including the nature and extent of caving zones,

interactions with water courses or storages, impacts on surface installations, and impacts on local communities; and

• risk assessment, especially for major hazards such as mud rushes, rock bursts, major

instabilities and associated air blasts. The data required for these purposes can be considered as falling into several categories. The general requirements in each of these categories will be outlined in Section 2.2. 2.2 GENERAL DATA REQUIREMENTS

2.2.1 Geology

It may be assumed that the regional geology will have been assessed during the exploration stage of a new project or will be well established and understood on continuing projects (eg Howell and Molloy 1960). The local mine geology must be known and understood in some detail for the purposes identified above. Knowledge is required of issues such as: • orebody shape, size and the distribution of grades; • location and nature of the contact of the orebody with the country rock; • the nature of the country rocks and of any weathered or transported overburden

materials; and • structural features such as faults, shear zones, dykes, sills and folding.


34

Generally, the information available from exploration drilling is incomplete with the result that planning and production engineers may find themselves "mining blind" (Hood et al 1999). Poor detailed knowledge of the orebody geometry in underground metalliferous mines can result in dilution or incomplete ore recovery or both. Developing the ability to "see" through the rock mass in order to gain a more detailed knowledge of the ore grades and boundaries, and of the rock structure and strength, would bring immense benefit. Geophysical techniques using seismic and electromagnetic methods, for example, are considered likely to provide an effective means of supplementing the information available from drilling (Hood et al 1999). It is especially important that both major and minor faults and shear zones intersecting the orebody and the nearby country rock be identified. The classification of faults and shear zones is considered in Section 2.3.1 below. The potentially deleterious effects of faults intersecting, or in close proximity to, mining excavations have long been recognised and dealt with. The nature and magnitudes of these effects may vary with the orientation of the fault, the geomechanical properties of the adjoining rock, the nature of the fault material and fault surfaces (friable or broken material, clay or other filling, slickensiding), the size of the excavation and the presence of water. Some of the observed effects of faults on underground mining excavations include: • off-setting of the orebody; • slip on the fault leading to a re-distribution of stresses around the excavation; • fretting or chimneying of friable fault and surrounding material above the back or

hangingwall; • isolation of large blocks or wedges that become free to slide or fall into the excavation; • general sloughing of destressed or unrestrained rock leading to dilution; • inability to form a satisfactory anchorage for, or to complete the installation of,

reinforcing elements such as rock bolts and cable bolts; and • the provision of a conduit for water flows into the excavations. Sourineni et al (1999) recently carried out a study of fault-related sloughing in open stopes and gave several examples of major fault-induced failure. Heslop (2000) points to several effects of faults in block caving operations while Laubscher (2000) notes particularly the potential for faults to isolate large wedges which may "sit down" on major apices or drawpoints and inhibit uniform draw and cave propagation.


35

2.2.2 Surface and Groundwater Hydrology

Surface and groundwater management is of little concern in some caving operations but is vitally important in others. It is necessary, therefore, that issues such as the location of surface water courses and storages, rainwater drainage and groundwater hydrology (including the potential for recharge) are evaluated in the feasibility study stage. If in caving operations, the ingress of water into the caving zone can be prevented, the mining excavations will serve to drain the surrounding rock mass with the result that the mine will be dry. However, in other cases, including areas of extremely high rainfall and where there are adjacent water storages and tailings dumps, water control and mud rush problems can be of concern (eg Barber et al 2000, Butcher et al 2000). These issues will be discussed in Chapter 10. 2.2.3 Topography and Environmental Constraints

The topography in the area of the orebody will have a major influence on the locations and costs of surface infrastructure and underground accesses. The local topography will also have an influence on the hydrological issues just discussed, on the local in situ stresses (see Section 2.8) and on the way in which any caving zone eventually propagates to surface (eg Brown and Ferguson 1979). Obviously, the existence of communities and of utilities such as roads, power lines and pipe lines of various types in the zone likely to be affected by the mine must be established and taken into account. There are many examples of the positive impacts of new mining projects on local communities by providing jobs, improved services and custom for local businesses. The social and environmental impacts of mining have become of major concern in recent years. They are noted here for completeness and will be considered no further. A particular issue in some parts of the world, including Australia, is native land rights and the existence of sacred and archaeological sites in the area influenced by mining. Expert studies of all of these issues are usually required to inform the definition of the mining environment of any mass mining project. The Century Zinc Project in North West Queensland, Australia, provides an especially good example of the successful resolution of issues of this type (Williams 1999). 2.2.4 Geotechnical Studies

Most of the key issues referred to in Section 2.1 outlining the uses of the data required to define the mining environment, require geotechnical data for their resolution. Accordingly, the emphasis of the remainder of this chapter will be on the collection and assessment of geotechnical data for rock mass characterisation. As well as the general geological and hydrogeological data referred to above, the geotechnical data required for the purposes being considered here includes:


36

• discontinuity survey data obtained through core logging, downhole logging of boreholes or scanline mapping of exposed faces. The data required are the locations, orientations, nature and condition of all discontinuities encountered and, in exposures, their terminations. These data are vitally important in caveability, fragmentation and excavation stability studies;

• measurements of the physical and mechanical properties of the lithological units

making up the orebody and the immediate country rocks. These include - unit weights, - uniaxial compressive and tensile strengths, - shear strengths of discontinuities, - shear strength parameters of intact rocks, - stiffnesses and deformation moduli of discontinuities and intact samples, and - hardness, toughness, abrasivity and drillability indices.

Standard methods of measuring these properties are given by Brown (1981). • rock mass classification of all lithological units using one of the established methods

such as those due to Barton et al (1974), Bieniawski (1976) or Laubscher (1977, 1994). These values are used in caveability studies, empirical methods of stability assessment and in estimating rock mass strengths using methods such as those developed by Laubscher (1977, 1994) and Hoek and Brown (1980, 1997); and

• measurements or estimates of the regional and mine site in situ stresses. The stresses

induced around mining excavations have major influences on excavation stability and, importantly in the current context, on cave propagation (Kendrick 1970, Krstulovic 1979, van As and Jeffrey 2000).

This geotechnical data collection phase is not always carried out adequately in terms of the nature, quantity or quality of the data collected or the time at which it becomes available for use in feasibility and subsequent mine design studies. The remainder of this chapter will concentrate on geotechnical data collection and its use in rock mass characterisation. A general review of the field will be given with emphasis on those topics to which particular contributions have been made as part of the International Caving Study Stage I as reported by Harries (2001). 2.3 CLASSIFICATION AND DESCRIPTION OF DISCONTINUITIES

2.3.1 Classification

The term discontinuity refers to any mechanical discontinuity in a rock mass having zero or low tensile strength. It is a collective term for most types of joints, weak bedding planes, weak


37

schistosity planes, weakness zones and faults. It contrasts with more specific terms such as bedding plane, joint or fault which describe discontinuities formed under particular conditions and mechanisms. Rock mass discontinuities can be classified using a number of criteria. A useful geometric classification used by structural geologists is to describe the discontinuity as a penetrative or non-penetrative structure. A particular structure is said to be penetrative if, on the scale under consideration, that structure is repeated with much the same spacing and orientation pattern, from one sample of the rock mass to the next. Otherwise the structure is non-penetrative; that is, the same structure may occur within different samples in different regions of the rock mass, but its distribution, spacing and orientation are not similar from one sample to the next (Hobbs 1993). Penetrative rock mass structures that can influence the mechanical properties and hence the behaviour of a rock mass include the bedding surfaces of sedimentary rocks, flow foliation of igneous rocks and foliation of regionally deformed metamorphic rocks. Typical non-penetrative discontinuities that influence the rock mass behaviour are joints and faults. It has been argued by Hobbs (1993) that "many engineering geologists have a preoccupation with joint surfaces as potential failure planes, to the exclusion of all other structures. While such a preoccupation with fracture systems is probably safe enough in weakly deformed sedimentary and igneous rocks, it can potentially lead to disaster if a keen appreciation of the penetrative structures in metamorphic rocks is not also actively maintained". Penetrative structures such as slaty cleavage or schistosity could be important in determining mechanical responses such as caveability and fragmentation. For engineering applications the most useful geometric classification of discontinuities is by scale. Discontinuities can be divided into two classes by size (Cruden 1977): • major discontinuities such as faults, dykes, contacts and related features with a size of the

same order of magnitude as that of the site to be characterised. The position in space, physical properties and geometrical characteristics are usually established deterministically for each of these major discontinuities; and

• minor discontinuities such as joints, minor shears and bedding planes which, for practical

purposes, represent an infinite population in the area of design. As a result, their geometrical characteristics and physical properties must be estimated by measurements of a representative sampled (smaller) population.

This division of discontinuities is important as it separates those features that may be represented deterministically from those that must be represented statistically. For minor discontinuities, representative sampling, sample size and the definition of structural domains are important aspects of a rock mass characterisation program. Structural domains are zones


38

of a rock mass in which the geometrical and physical properties of the discontinuities can be treated as being statistically homogenous. Discontinuities may also be classified according to their origins. Joints (minor fractures)

Joints are the most common result of brittle fracture in the Earth’s crust. They are ubiquitous geological structures, occurring in a wide variety of rock types and tectonic environments. They also have a dominant effect on the mechanical and hydrological responses of rock masses to engineering activity. In discussing minor fractures (joints), Price and Cosgrove (1990) comment that one of the topics that bedevils fracture interpretation is nomenclature. Griggs and Handin (1960) classify discontinuities as faults if they exhibit shear displacement and as joints if they are dilatational features which exhibit no shear. In practice, this approach could lead to the misinterpretation of geological data. The shortcoming of the definition is that the scale at which the observation is made is not specified. Geologists who are unable to demonstrate a shear displacement on a fracture may be tempted to classify such a fracture as a ‘joint’ resulting only from extension.

Figure 2.1: Photomicrograph of quartz and calcite filled discontinuities (Price and Cosgrove 1990)

Figure 2.1 shows a photomicrograph of intersecting micro-fractures. The photomicrograph of the rock section under crossed polarised light quite clearly shows two quartz filled fractures (both showing undulose extinction) oriented vertically and a single horizontal calcite filled fracture showing high order birefringence. From this photomicrograph, it can be seen that the calcite-filled vein displaces the quartz filled veins (by approximately 0.025 mm). This small displacement can only be inferred because of the infill of the fractures and the use of

Calcite

Quartz


39

microphotography. The two earlier quartz veins are dilational (ie extensional) features and the later calcite vein is a hybrid extension/shear fracture. It is likely that from a field observation, the shear movement would be missed and all three fractures would be identified as joints using Griggs and Handin’s definitions. On the basis of field experience, Harries (2001) has suggested that, in order to avoid such problems, the term joint should be used to describe any minor fracture which does not exhibit a shear displacement of greater than 0.01 m (10 mm). The term fault is used where a shear displacement of greater than 0.01 m is observed along the discontinuity plane during mapping or core logging. Joints often display spatial and orientational relationships with folds, anticlines, synclines and faults formed during some period of tectonic activity (Price 1966, Price and Cosgrove 1990). An example of the jointing associated with folding is shown in Figure 2.2. Veins, or cemented joints, as illustrated in Figure 2.1, are mineral infillings of joints or fissures. They may be sheet-like or tabular or irregular. They are generally of igneous origin but may also result from sedimentary processes. They are commonly associated with metalliferous orebodies and have been found to have important influences on orebody caveability and fragmentation. They may be weaker or stronger than the wall rock of the joint or fissure and should be taken into account in rock mass characterisation schemes (eg. Laubscher and Jakubec 2000). For example, at the El Teniente mine, Chile, the stronger primary ore contains large numbers of veins or filled joints. These veins are described in order of decreasing strength as (Flores and Karzulovic 2002a): • quartz filled joints; • anhydrite-chalocpyrite filled joints; • anhydrite-gypsum filled joints; or • joints with soft fillings (eg. sericite, oxides).

Figure 2.2: Jointing associated with an asymmetrical anticline (Priest 1993 after Price 1966)


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Faults and shears

As discussed above, a fault is defined as a discontinuity dividing portions of rock that have been displaced one past the other in shear. Faulting of rock can occur along a single plane or along many planes within a zone. A fault zone is a closely spaced group of parallel or anastomising faults. A zone in which blocks of rock have been displaced but do not display visible fault structures, is termed a shear zone (Hobbs et al 1976). Although shear displacements on discontinuities can range from micrometres to several hundred kilometres, the term fault is reserved for the more extensive features that show significant displacement (>0.01 m, for example). Faults are usually classified on the basis of shear direction, with three main types of fault being defined: • normal faults in which the hangingwall has moved down with respect to the footwall; • reverse (thrust) faults in which the hangingwall has moved up relative to the footwall; and • strike-slip (transcurrent) faults in which the movement is predominantly sideways along

strike. It is assumed that a fault is induced when changing tectonic stresses produce a shear stress that exceeds the shear strength developed on a particular plane in the rock mass. The type and characteristics of the fault will be controlled by the shear strength of the rock mass and the in situ stress conditions at the time of fault formation. Minor shears and joints will often form as secondary features of the main fault (Kersten 1990). Their orientation, persistence and thickness, and the nature and strength of the infilling materials, influence the effects of faults on mining operations. For a given fault, these features may not be uniform with depth. For example, at the El Teniente and Andina mines in Chile, several examples exist of faults with thicknesses of, say, 1.5 m on the surface, having thicknesses of only 15 cm at depths of 1000 m (Flores and Karzulovic 2002a). Cleavage or schistosity

Cleavage or schistosity is predominantly a planar rock fabric (foliation) produced by preferred alignment of platy minerals (generally phyllosilicates). This alignment is not perfect, but is of a statistical nature. Cleavage imparts a special property to the rock, in that it splits preferentially in a direction parallel to the cleavage planes (Ramsay and Huber 1983). The most common type of cleavage is flow cleavage, which is caused by the recrystallisation and realignment of platy minerals during tectonic deformation. This is often associated with low grade regional metamorphism. Fracture cleavage describes incipient, cemented or welded parallel discontinuities that are independent of any parallel alignment of minerals (foliation). Fracture cleavage is another product of high deviatoric stresses developed during tectonic deformation.


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Bedding

Bedding is typically a continuous plane that divides sedimentary rocks into beds or strata. Similar structures can also occur in igneous rocks from lava flows or the deposition of pyroclastic material. They are created by changes in such factors as grain size, grain orientation, mineralogy or chemistry during deposition. Bedding does not always create discontinuities; in many cases it forms only a slight change in colour or texture in an otherwise intact rock material (Priest 1993). Bedding planes may exist as open fractures or as closed planes along which the rock may part easily (Gerrard 1988). At depth bedding planes are often closed but the release of stress in the rock immediately adjacent to a newly formed excavation may allow bedding planes to part and discontinuities to form (Beer et al 1983). Although initially horizontal and generally planar, bedding can be tilted, folded and even inverted to a complex range of orientations. Bedding features can be recognised by the fact that they are generally parallel, even when tilted or folded. Sometimes there may be minor stratification planes known as cross-bedding, oblique to the major bedding planes. 2.3.2 Description

In the ‘Suggested methods for the quantitative description of discontinuities in rock masses’ of the International Society of Rock Mechanics (ISRM 1978), ten parameters are identified as being required for the quantitative description of discontinuities and rock masses. These parameters are illustrated schematically in Figure 2.3.

Figure 2.3: Discontinuity parameters

(Hudson and Harrison 1997)


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Five of the ten parameters (orientation, spacing, persistence, number of sets and block size) may be considered to be ‘geometric’ parameters. They will define the geometry of the rock mass structure, the size and shape of rock blocks formed and the nature of intact rock bridges in the rock mass. The other five parameters (roughness, aperture, filling, wall strength and seepage) may be regarded as ‘strength’ parameters because they influence the discontinuity’s shear strength and stiffness. 2.4 DISCONTINUITY DATA COLLECTION BY DRILLING, CORE LOGGING, DOWN-HOLE

SURVEYS, SCANLINE AND CELL MAPPING

2.4.1 Introduction

The geometric and mechanical characterisation of discontinuities is a vitally important pre-cursor to engineering design in rock masses. Ideally, the complete characterisation of a rock mass would include a description of each fracture in the rock mass as well as the determination of its geometric and mechanical properties. Currently, this cannot be achieved for a number of reasons: • the visible parts of discontinuities are usually limited to discontinuity traces only; • discontinuities distant from the exposed rock surfaces and drill core cannot be observed;

and • direct and indirect (eg geophysical) measurements of discontinuities have limited

resolution and accuracy. For these reasons, the discontinuities in a rock mass are usually described as an assemblage, rather than individually. This assemblage is modelled stochastically simply because the discontinuity characteristics vary in space (Dershowitz and Einstein 1988). The methods used to collect rock mass discontinuity data can be divided into four main categories: • geotechnical core logging and borehole imaging; • face exposure (planar) mapping; • geophysical or indirect mapping; and • aerial and photogrammetric techniques. The fourth category will not be discussed here as it has little relevance to rock mass characterisation for cave mine engineering other than in the exploration stage. Aerial photography can identify major lineaments and the orientations of large-scale discontinuities. Photogrammetric methods have proved useful for rock mass characterisation for open pit mines (Tsoutrelis et al 1990).


43

2.4.2 Geotechnical Core Logging

The recovery of core by diamond drilling allows information to be obtained from volumes of a rock mass which cannot be observed directly. It is one of the most important and valuable methods of sub-surface exploration and, in some cases, is likely to provide the only direct sampling or observation of much of the rock mass that is to be mined. A large assortment of drilling rigs, core barrels and drill bits are available to provide drill core at diameters from 20 to 150 mm at varying depths and from rocks of varying strength. Hoek and Brown (1980) and Brady and Brown (1993) discuss the principal types of drill rigs, core barrels and drill bits used in diamond drilling. Most of the core drilled during the feasibility and planning stages of a block caving operation will be used to determine the rock types, ore grades, ore textures and geological structures required to develop a geological model of the orebody. Nevertheless, for engineering purposes, valuable information concerning the rock mass can be obtained by a critical examination of core, providing that the examiner is aware of those geological features that are of significance (Deere 1964). The quality of the core record that is obtained and logged is dependent on a number of factors, including: • the rock strength and behaviour of discontinuities during drilling; • the drilling equipment and core diameter used in the coring process; • the competence of the drill rig operators; and • how the core is handled and stored. These four factors have been discussed in detail by Onederra (1999) and earlier in the seminal paper by Rosengren (1970). After drilling, wherever possible, the drill core should be oriented before being logged geotechnically. Drill core orientation and the determination of discontinuity orientation are important because of the impact that discontinuity orientation has on the caveability, fragmentation and support requirements of a rock mass. During a drill run, the core will tend to rotate within the core barrel. Thus, the true orientations of discontinuities within rotated core will remain unknown unless they can be correctly reoriented to their original positions within the borehole. The techniques that can be used to assist with the reorientation of core include: • down-hole orientation instrumentation; • the use of reliable reference planes; • the use of cameras and geophysical visualisation techniques; and • the use of multiple holes and stereographic techniques. These techniques will be discussed in turn.


44

Down-hole orientation instrumentation

Down-hole orientation devices assist with orienting sections of the core into their correct positions within the drill hole. If the plunge and trend of the borehole are known, the measured orientation of the discontinuity in the borehole can be used to determine its true orientation. Sullivan et al (1992) compiled a useful summary highlighting the advantages and disadvantages of the main down-hole core orientation techniques used in Australia (see Table 2.1).

Table 2.1: Core orientation techniques (after Sullivan et al 1992)

Technique Cost* Complexity

in Use

Applicable rock

strength

Advantages Disadvantages

Low High

Christensen-

Hügel

High Moderate Yes Yes Can be used in low

strength rocks and

vertical holes

Requires

conventional

drilling and is

time consuming

Craelius Moderate Moderate Marginal

**

Yes Simple in use,

negligible drilling

delays

Can be subject

to damage. Not

accurate for high

angle defects

Spear Negligible Very low Marginal

**

Yes Simple in use,

negligible drilling

delays

Reliability of

result

Clay imprint Negligible Low Marginal

**

Yes Simple in use,

minimal drilling

delays

Requires

interpretation of

imprint

Acid etch Moderate Moderate Yes Yes Provides core

orientation and

additional borehole

survey

Time consuming

Rocha Very high High Excellent No Can be used in

extremely low

strength rocks

Expensive and

requires

specialist

equipment * Both initial outlay and delay in drilling ** Due to top of core run being disturbed during drilling


45

Sullivan et al (1992) reported that the most common and reliable of the core orientation methods used in Australia was the clay imprint technique. However, there is a trend for drilling contractors to use the down-hole spear technique. The advantages of the spear technique are the almost negligible costs and delays to the drilling operation. As most exploration drilling contractors are paid primarily by the metre drilled, any delay in drilling is seen to be a waste of money. During the geotechnical drilling campaign at Mont Porphyre, Canada, Coulson et al (1998) and Nickson et al (2000) compared five core orientation methods. They found that the Core Tech Canada diamond drill core orientation system, which uses an acid etching technique, gave reliable orientation results. The clay imprint method worked with some success but at depths greater than 1200 m difficulty was experienced in obtaining an adequate imprint. A third system used scribes three lines on the core in a known orientation. Problems were found in using this technique because of the high rock hardness and the induced rotation of the core tube in the barrel. The other methods considered were a borehole camera and the orientation to known bedding, both of which will be discussed below. Because of the depth of the Mont Porphyre orebody (1000 to 1700 m below surface), operational difficulties were experienced in orientating the drill core. The high cost of drilling at such depths and the small number of holes drilled, meant that the cost and time associated with one of the more advanced core orientation systems was acceptable at Mont Porphyre. However, this is unlikely to be the case in other operations which have a greater number of holes, shallower drilling depths and direct underground access to the orebody from exploration drifts, for example, enabling reliable discontinuity orientation data to be obtained. Utilising reference planes of known orientation

If the rock mass contains a planar fabric that is consistent throughout the site (eg bedding planes or a regionally consistent cleavage), and if the orientation of the fabric is known, the true orientation of the discontinuity can be established. As long as the reference plane is consistent and the borehole direction (trend and plunge) is known, the angle measured between the reference plane and the discontinuity can be used to determine the dip and dip direction of the discontinuity. The method described by Priest (1985) makes use of the stereographic projection technique. Ideally, another orientation method should be used to check the results obtained using this technique. If there is a major deflection of the reference fabric, perhaps due to faulting or folding, then gross errors can be produced in the estimated discontinuity orientation. Utilising instruments such as cameras or geophysical visualisation techniques

Downhole cameras or geophysical techniques that can measure discontinuity orientation can be used to examine the sides of a borehole. Discontinuities intersecting the borehole can be analysed and their orientations determined. These discontinuities can then be related to the


46

corresponding structures found in the core at the same depths. Although this is a slow process for determining discontinuity orientation, it may be of use where no other orientation methods are available or where processes such as core washing and core blockage have damaged the core to such an extent that traditional orientation methods cannot be used. Using stereographic techniques and multiple holes with different orientations

The true orientations of sets of discontinuities can be determined using the orientations of two or more non-parallel boreholes. Using the angle between the core axis and the discontinuity normal, vital information on the discontinuity orientation can be obtained. Rotation of the core during drilling and handling may mean that the true orientation of the discontinuity cannot be determined using this measurement alone. The angle represents a locus around the drilling direction of possible discontinuity orientations. However, using a number of measurements of this angle for the same discontinuity set, made at different drilling directions, the orientation of the discontinuity set can be calculated. Priest (1985) provides a description and an example of the use of this technique to determine discontinuity orientation. Reorienting the core and establishing a reference line are the next steps required in preparation for core logging. Using the core orientation information, the core is rotated into position with the reference line placed on top. Discontinuities or broken pieces of core are put into their correct positions so that the reference line can be extended over discontinuities and drilling-induced fractures. Problems can occur if a large section of broken core or a zone of washed away fault gouge means that the core pieces cannot be fitted together. Geotechnical core logging is intended primarily to provide information on the rock mass discontinuities. There are a number of other basic steps or procedures that should be carried out prior to the detailed geotechnical logging of core, namely: • the core should be examined to determine the structural boundaries and the geological

features to be measured. Markers indicating the depths of the geological horizons and start and end of drill runs should be checked (Brown 1981);

• core recovery should be established. This parameter may be determined for individual core

runs or rock types or for entire boreholes. Recovery in a rock mass of poor quality will be strongly dependent on the drilling equipment used and the skill of the drilling crew; and

• an assessment should be made of the degree of fragmentation of the drill core. This

assessment can be made quantitatively using a number of indices including the Total Core Recovery (TCR), Solid Core Recovery (SCR), Rock Quality Designation (RQD), Fracture Index (FI) and Core Loss (CL) (Windsor and Thompson 1997). The most commonly used measures are the RQD and the Fracture Index (or Fracture Frequency) which find widespread use in rock mass classification schemes and in engineering design.


47

When determining discontinuity frequency indices care must taken to distinguish between natural discontinuities and fractures caused by the drilling and handling processes. A trained engineering geologist or geotechnical engineer can usually identify natural discontinuities because of the presence of mineral coatings, smoothness, staining and/or weathering of the surrounding rock. Three methods may be used for determining the lengths of core pieces for RQD calculations - tip to tip, centre-line and fully circular (see Figure 2.4). Tip to tip measurements involve double counting at each end of a core piece, while fully circular measurement ignores core pieces that happen to have been drilled with a small subtended angle to one discontinuity in otherwise massive rock. Consequently the centre-line method is recommended (Brown 1981). A discussion of the calculation of RQD is given in Section 2.7.2 below.

Fully

circ

ular

Cen

tre li

ne

Tip

to ti

p

Cen

tre li

ne

Zero

Fu

lly c

ircul

ar

Figure 2.4: Different interpretations of the length of core pieces (after Brown 1981)

Following the definition of geotechnical domains and the calculation of core recovery and discontinuity intensity measures, the actual logging of the core can proceed. The amount of information concerning rock mass discontinuities that can be obtained from logging drill core will depend on whether or not the core has been oriented. The ten specific parameters described under the ISRM suggested methods (Brown 1981) are listed in Table 2.2. If the core is not oriented, only three of the ten parameters can be recorded reliably. The more disturbance the core has undergone during drilling and handling then generally the less the information that can be obtained on the rock mass discontinuities. For example, care needs to


48

be taken when assessing discontinuity filling. Discontinuity filling and mineralisation may be washed away by drilling fluids. In some situations drilling mud may be deposited into open discontinuities.

Table 2.2: Discontinuity parameters measurable from core record (Harries 2001)

Discontinuity Parameter

(ISRM 1978) Oriented Core Unoriented Core

1. Orientation Yes No (possible)

2. Spacing Yes No (possible)

3. Persistence No No 4. Roughness Yes Yes

5. Wall Strength Yes Yes

6. Aperture No No 7. Filling Yes Yes

8. Seepage No No

9. Number of Sets Yes No (possible) 10. Block Size No No

Even when the core has been successfully oriented, the errors associated with sampling in one direction only always need to be considered. A ‘blind zone’ will be present even when a perfect drilling operation is achieved. Only by the use of multiple drill holes at different orientations can this sampling issue be resolved. The main drawback of characterising discontinuities from drill core is the lack of information available on persistence. Because of the small volume of the rock mass contained in the core, it is unlikely that many observations on discontinuity termination will be made or much information obtained on discontinuity shape or size. Optical imaging of the drill hole walls is also used in rock mass characterisation. A variety of techniques are available including photoelectric transformers, conventional cameras operated remotely from a wireline, television cameras operated remotely using coaxial cable and digital borehole scanners. Discontinuity detection applications are sometimes limited because the imaging techniques used require clear borehole fluids, and because they commonly image the borehole at oblique angles using local illumination sources that produce shadows which interfere with discontinuity detection and interpretation (Paillet et al 1991). Furthermore, some data may not be in digital form initially, and so not amenable to sophisticated processing. The digital borehole scanner (DBS) is a logging tool which provides optical, true-colour images of a borehole wall. The Borehole Image Processing (BIP) system was developed by the Raax Company of Kitaku, Sapporo, Japan (Kamewada et al 1990). The digitally recorded data yield high resolution images, enabling detailed measurement of discontinuities, identification of


49

discontinuity mineralisation, and observation of other microscale properties not generally available with other instruments. However, the sampling issues (orientation blind zone and lack of persistence information) associated with logging oriented geotechnical drill core also apply to borehole imaging techniques. 2.4.3 Exposure Mapping Methods

Mapping of natural (outcrop) or artificial (excavated) rock mass exposures, can enable discontinuity parameters to be characterised in greater detail than is possible from drill core logging. In particular, important information about how discontinuities terminate, the sizes of discontinuities and reliable measurements of discontinuity orientation can be obtained from exposure mapping. Mapping techniques can be divided into three main classes: • spot mapping; • lineal mapping; and • areal mapping. The technique used in a given rock mass characterisation campaign depends on the degree of detail and hence the sampling effort required. All techniques will suffer from some degree of sampling bias (see Section 2.5.2), because discontinuities having three dimensional parameters (size, shape and orientation) are being sampled in a two dimensional section (plane). The mapping technique used will determine which analysis techniques can be used to correct for these sampling biases. Spot mapping

Spot mapping is a technique in which the observer selectively samples only those individual discontinuities that are considered important. An example of the application of this technique is shown in Figure 2.5. The actual traces observed in the exposure are shown on the left of the figure. The discontinuities depicted on the right of the figure are those that have been mapped because the mapping personnel have considered them important. The value of the results obtained using this technique depends on the judgement of the observer. By recording the characteristics of only those discontinuities thought to be "important", the volume of data, and thus the mapping effort required, are greatly reduced. However, because the user biases the data, the repeatability is poor. Individual users may consider different discontinuities to be important. This may be the case particularly when those performing the mapping are collecting the data for different engineering purposes (eg caveability analysis, fragmentation analysis, cable bolt design). The use of this method is advised only for preliminary ‘reconnaissance’ style data gathering exercises where there is a need to gather some initial orientation data quickly. This can provide useful input into planning more comprehensive and objective mapping exercises (eg by scanline or window mapping).


50

Original exposure plane Spot mapping

Figure 2.5: Spot mapping technique (Harries 2001) Lineal (scanline) mapping

Lineal mapping involves measuring or recording the geometric and mechanical characteristics of all discontinuities that intersect a given sampling line. Examples of lineal or scanline mapping techniques are given by Piteau (1970), Priest and Hudson (1981), Villaescusa (1991), Priest (1993), Windsor and Thompson (1997) and Harries (2001). When carried out correctly lineal mapping can provide significant amounts of data from a sample of the rock mass in a structured and objective way. This technique is illustrated schematically in Figure 2.6. The left of the figure shows the original rock mass discontinuity traces and the right of the figure shows those discontinuities selected during mapping. All discontinuities that intersect either of the horizontal or vertical sampling scanlines and are larger than a designated "cut off" limit are mapped and characterised systematically.

Original exposure plane Line mapping

Figure 2.6: Scanline mapping technique (Harries 2001)


51

A typical scanline mapping sheet is shown in Figure 2.7. The header information for the scanline survey includes the coordinates of the mapping location, scanline orientation and elevation, wall information, date and personnel involved in the scanline mapping. The scanline (tape measure) is set up on an appropriate clean wall. This may require the wall to be cleaned. If the outcrop is too weathered or the exposure too blast damaged to enable the discontinuities to be mapped, another location should be selected. After selection of an appropriate site, starting at one end of the scanline, every discontinuity that intersects the scanline tape is measured. The discontinuity parameters measured will depend on the scheme adopted. Using the Villaescusa (1991) scanline mapping format, the following information is collected for each discontinuity intersected: • distance of intersection along the tape; • number of endpoints observed in the plane (0, 1 or 2); • discontinuity type (joint, fault, vein, bedding, shear); • orientation (dip and dip direction of the discontinuity); • roughness (smooth, rough, slickensided); • planarity (planar, wavy, irregular); • trace length (length of discontinuity seen in the sample plane); and • termination types (intact rock, another joint or hidden)

Figure 2.7: Scanline mapping sheet (adapted from Villaescusa 1991)

Simple notation and abbreviations are used to optimise the detail of the observations and increase the speed of mapping. Discontinuity parameters are assessed using simple tools and observations. Discontinuity orientation is usually measured using the Brunton or Clar compass.

(>20°)


52

The trace length is measured using a tape measure and the other parameters can be assessed visually. The ‘remarks’ column can be used for any other general observations or explanation such as the nature of discontinuity mineralisation or alteration. The length of the scanline is normally extended until a prerequisite number of observations are obtained. What is a prerequisite number? Numbers varying from 40 per discontinuity set and from 150 to 350 total discontinuities per scanline have been suggested. Priest (1993) suggests that between 150 and 350 discontinuities should be recorded. The lower number would be sufficient for a ‘simple’ rock mass containing just three discontinuity sets and the higher number for a more complex rock mass containing up to six discontinuity sets. Savely (1972) found that at least 60 observations were required to stereologically define a discontinuity set along a given sample line. Villaescusa (1991) suggests that at least 40 discontinuity observations are required per set to provide a sound statistical database of the discontinuity set characteristics. Areal (window) mapping

Areal mapping involves collecting all data from within a specified area of a rock face, often referred to as a ‘window’. Pahl (1981) gives a discussion of the window mapping technique. The preliminaries and measurement techniques for window mapping are the same as those for scanline mapping, except that all the discontinuities that are above a given "cut off" size are recorded (see Figure 2.8).

Original exposure plane Window mapping

Figure 2.8: Window mapping technique (Harries 2001)

Although this method is more time consuming than scanline mapping, the amount of data collected is even greater and in some cases the amount of geometric bias may be reduced.


53

Geophysical techniques

Geophysical techniques can be used to gain a measure of the geological sequence and structure of the rock mass. These techniques measure various geophysical properties of the rock mass that can be related to its lithological and geotechnical properties. Most of these geophysical methods can measure discontinuities only indirectly. Typically, the reduced data from each method (eg seismic travel times) must be inverted to yield estimates of local rock properties (eg seismic velocities). Normally these rock properties are not discontinuity properties (eg discontinuity density). Rather, the discontinuity properties must be deduced indirectly from the rock properties. This deduction requires strong idealisations of discontinuity geometry. These discontinuity properties are not always the properties of direct interest in many engineering applications (eg discontinuity intensity or in situ block size). The required properties must be interpreted from the deduced fracture properties. This interpretation requires higher levels of subjectivity than the first (inversion) or the second (deduction) steps (National Research Council 1996). Geophysical methods naturally divide themselves into three distinct scales: • large scales associated with surface soundings; • intermediate scales associated with surface to borehole and borehole to borehole

soundings; and • small scales associated with measurements made on rock immediately adjacent to a

borehole or tunnel.

Table 2.3: Geophysical discontinuity detection methods (after National Research Council 1996)

Method Length scale

(resolution) Remarks

Elastic methods: seismic band

(10-100 Hz)

100 to 5000m zero shear modulus in fracture fluid is critical

Elastic methods: sonic band (2-20 Hz)

0.1 to 10m zero shear modulus in fracture fluid is critical

Elastic methods: ultrasonic band (200-2000 kHz)

0.1 to 10m fracture aperture is critical

Electrical methods 10 to 300m contrasting resistivity of fracture fluid is critical

Electromagnetic methods 10 to 300m contrasting resistivity of fracture fluid is critical

Radar methods 3 to 200m contrasting resistivity of fracture fluid is critical

Conventional well logs 0.1 to 10m near borehole environment

Large scale geophysical measurements such as those used in the petroleum industry can be used to identify major discontinuities. Major discontinuities such as large fracture zones and faults


54

can be characterised by seismic, electrical and electromagnetic methods. Minor discontinuities are extremely difficult to detect using these large scale geophysical methods. Intermediate scale methods involving surface to borehole and borehole to borehole soundings have many advantages over the larger scale surface geophysical methods. In many cases, overburden acts as a filter because of its attenuation properties and high contrast, requiring the use of complex correction procedures to obtain useful information. Boreholes and tunnels provide access to measurement points below the surface, allowing many of the problems arising from overburden to be avoided. Compared to surface surveys, borehole measurements sometimes require complex (and compact) sensors. Remote sensing is done with sources and receivers placed in the same borehole, or one in the borehole and the other at the ground surface (National Research Council 1996). Geophysical systems consisting of a downhole probe (or tool) attached to a multi-conductor electric cable often referred to as a ‘wireline’, can also be used to characterise rock masses. The wireline is attached at the surface to a winching assembly that controls the lowering or raising of the probe. Generally, several types of geophysical devices are combined to form one downhole logging tool. The main types of geophysical wireline logging tests include seismic velocity, seisveiwer, electrical resistivity, gamma-gamma and self potential. The applicability of these techniques in assessing discontinuity parameters is outlined in Table 2.4. None of the measurement techniques directly measure discontinuity parameters but some have a correlation between the test result and the discontinuity parameter.

Table 2.4: Wireline logging techniques (after Windsor and Thompson 1997)

Discontinuity Parameter Type of test

Orientation Spacing Type Separation Filling

Seismic

Velocity

no correlation some

correlation

no correlation no correlation no correlation

Seisviewer or

acoustic

scanner

direct

correlation

direct

correlation

no correlation direct

correlation

some

correlation

Electrical

resistivity

some

correlation

some

correlation

no correlation no correlation some

correlation

Gamma-

gamma

no correlation some

correlation

no correlation no correlation some

correlation

Self-potential no correlation no correlation no correlation no correlation no correlation

In all cases, the emphasis on using the above techniques is to determine the properties of the rock, the in situ locations, orientations and nature of the discontinuities and the hydrological conditions that exist within the rock mass. To ensure the best results, these techniques should


55

be compared with descriptions of sections of the core, and if necessary compared directly with the core. Rock material properties that may be needed in the geophysical data reduction can then be obtained. This allows the geophysical testing to be calibrated to the observed rock mass conditions. Recently, the acoustic scanner has been shown to be an excellent tool for providing oriented images of borehole walls at very fine resolution (Hatherly and Medhurst 2000). Hatherly (2001) suggests that the geophysical tools of relevance in the detection and measurement of defects are acoustic scanners (televiewers), optical scanners (for dry holes), dip meters and full waveform sonic logs. Those methods which rely on the detection of water-filled discontinuities by electrical and electromagnetic methods may not always be suitable for hard rock mining applications. 2.5 ANALYSIS AND PRESENTATION OF DISCONTINUITY DATA

2.5.1 Introduction

The objective of rock mass discontinuity analysis is to estimate the characteristics of the discontinuities and to determine how they vary over the mine site. The estimate of rock mass discontinuity properties or rock mass model, is merely one representation of the results from the collected sample data. Unfortunately, the true characteristics of the rock mass discontinuities can never be known exactly. This would require all of the discontinuities in the rock mass to be measured accurately and their engineering properties established by testing. Major discontinuities such as faults, dykes, geological contacts and unconformities may be analysed deterministically in cave mine design. For all practical design purposes, minor discontinuities represent an almost infinite population. Their characteristics will be assessed by analysis of a much smaller sample population. The objective of this statistical sampling is to infer the characteristics of a large population without including all its members in the sample. It is the statistical treatment of these minor discontinuities that will be discussed here. In order to establish the best estimate of discontinuity properties, the measurement of discontinuity characteristics should be unbiased and objective. A number of biases introduced during sampling should be accounted for either before or during the analysis process. The method of data collection is the most important factor influencing the objectivity of the discontinuity analysis; some methods are inherently more subjective than others. This subjectivity in the data collection is reflected in possible errors in the analysis. It is important to realise that it is much harder to account for error due to subjectivity than for systematic sampling errors.


56

2.5.2 Error and Uncertainty in Discontinuity Analysis

Einstein and Baecher (1982) have defined three main sources of uncertainty and error in engineering geology and rock mechanics:

• the innate, spatial variability of geological formations, where wrongly made interpretations of the geological setting may be of significant consequence;

• errors introduced in measuring and estimating engineering properties, often related to sampling and measurement techniques; and

• inaccuracies introduced in modelling physical behaviour, including the use of incorrect or inapplicable types of calculations or models.

Uncertainty and error will be associated with any rock mass characterisation study. While uncertainty can never be removed completely, it should be reduced to a minimum. An acceptable level of uncertainty can be quite difficult to establish. In an engineering study, what has been left out of the analysis cannot always be known. According to Einstein and Baecher (1982) most of the major failures of constructed facilities have been attributed to omissions. Furthermore, real rock masses will have properties and variability that can never be accounted for fully in characterisation, rock mass classification or design analyses. Uncertainties caused by inherent geological spatial variability

The geological subsurface is spatially variable in that it is composed of different materials which are stratified, truncated, and in other ways separated into more or less discrete zones. It is impossible to account for all zonation of the rock mass because much of it will be unknown, especially in cave mine engineering where a majority of the orebody will have been characterised from drill core only. This limited sampling will hinder the development of a highly detailed geological model. The other spatial variability occurs within an apparently homogenous body because material properties may vary from point to point. While this variability can be precisely characterised with sufficient observations, the numbers of observations available are usually limited. Thus, uncertainty will remain about the material properties or classifications at points that have not been observed or sampled. Error arising from variability in the geology can never be avoided. It can be reduced by the use of experienced geologists having extensive knowledge of the region concerned and by directing specific investigations towards possible key geologic structures (Palmstrom 1995). Measurement errors and sample biases

The main measurement error that occurs in discontinuity characterisation is in the measurement of discontinuity orientation. West (1979) carried out a study of measurement errors in discontinuity mapping by examining the reproducibility of frequency and orientation measurements made during scanline mapping at a quarry in the Lower Chalk in Oxfordshire,


57

England. He concluded that the orientation of a well-defined joint could be determined to within about ±6° for dip direction and ±5° for dip angle. Subsequently, Ewan and West (1981) examined the reproducibility of joint orientation measurements using six different observers performing 10 m long scanline surveys. They found that the orientation of particular joints recorded can have a mean maximum variation of ±10° for dip direction and ±5° for dip angle without there being any real difference in the joint orientation. However, joint orientation diagrams created for the collected discontinuity data identified the same major joint sets irrespective of observer. In this case, many of the measurements were taken in a full-face machine bored tunnel having smooth walls. It is difficult to obtain measurements of discontinuities from smooth walls because there may not be sufficient discontinuities exposed to enable accurate orientation measurements to be made directly. In these cases a notebook or other planar device is used to project the discontinuity plane from which a measurement can be made. There are four main types of sampling bias in discontinuity measurement: • Orientation bias – the frequency of discontinuities intersecting a particular window,

scanline or drill core depends on the orientation of the sampling geometry relative to the orientation of the discontinuity set. If a discontinuity set is oriented parallel to a window, then few discontinuities in this set will intersect the window.

• Size bias – the larger the scale of a discontinuity the more likely it is to be sampled by a

given drill core, scanline or mapping window. • Truncation bias – a truncation or size cut-off is usually used in scanline or window

mapping. For example, fractures that are less than 50 mm in length may be ignored. Although using such a small cut-off will usually have little effect on the overall discontinuity statistics, if a comprehensive, rigorous analysis is undertaken with the aim of fully describing the distribution of discontinuity sizes then the truncation size cut-off must be taken into account.

• Censoring bias – this bias is associated with the artificial boundaries that are imposed

when performing rock mass characterisation. Typically in underground mines the most limiting boundary is the height of the drives in which mapping takes place. The restriction in height of the mapping window limits the trace lengths that can be observed. Censored trace lengths provide a lower bound estimate of the true trace lengths.

For an unbiased structural analysis, the measurement process needs to be objective. Biases can be reduced by the appropriate selection of rock mass characterisation methodologies and sites. For example, if three mutually orthogonal scanlines are used, the bias due to orientation will be greatly reduced. However, it is not possible to account for all biases, which must then be removed using analytical or numerical methods.


58

Modelling errors

Models used in assessing discontinuity parameters and modelling rock mass discontinuity geometry are based mainly on observations of discontinuities and a number of assumptions made about the geometry and characteristics of the discontinuities. As they are simplifications of reality, modelling errors are introduced. An example of this kind of error is the use of a rock mass model that models all discontinuities as orthogonal planes of infinite persistence (eg Snow 1965). However, it must be recognised that a degree of simplification is required in the development of any model. 2.5.3 Discontinuity Orientation Analysis

The main aim of an orientation analysis is to establish a statistical model of the orientational arrangement of the discontinuities contained within the rock mass. The underlying premise is that geological processes have generated one or more sets (or clusters) of nearly parallel discontinuities in the rock mass. From the discontinuity orientation data collected, a statistical model that represents the discontinuity orientation characteristics of the rock mass can be constructed. The discontinuity orientation characteristics that are of most interest in rock engineering include: • the number of discontinuity sets in the rock mass; • the mean orientations (dip and dip direction) of these discontinuity sets; • the spread or dispersion of orientations around a given set’s mean orientation; and • the amount of data that lies outside defined discontinuity set limits. The first step in the analysis is to determine the number of discontinuity sets in the rock mass and to define the limits of those sets. The main technique used in the identification of discontinuity sets involves presentation of the data in a graphical format. Prior to the widespread use of computers for the plotting and contouring of orientation data, orientation analyses were conducted manually. This manual technique is described by Hoek and Bray (1974) and Priest (1985). A number of computer programs have become available to assist in the plotting and analysis of orientation data, including DIPS (Rocscience Inc. 1999), SAFEX (Windsor and Thompson 1990) and CANDO (Priest 1993). Whether a manual technique or a computer program is used, the best way of representing orientational data graphically is with respect to the surface of a reference sphere using the stereographic projection. Discontinuity orientation data are usually represented by unit vectors normal to the discontinuity planar surface. The discontinuity unit normals are recorded unambiguously by polar coordinates using two angles, the trend, α, and the plunge, β. In the development of rock mechanics and rock engineering, there has been an almost total adoption of equal-angle lower hemisphere projection (Hudson and Cosgrove 1997). An example of an equal-angle lower hemisphere projection is shown in Figure 2.9.


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Figure 2.9: Example of a stereographic projection of discontinuity unit normals (from Rocscience Inc. 1999)

The discontinuity normals (or poles) can be seen to cluster into three distinct groups which represent discontinuity sets. These independent discontinuity sets need to be identified and analysed separately. Design set boundaries for each of the discontinuity sets need to be identified. A technique that can aid in the definition of discontinuity sets and their boundaries involves contouring the data. The data shown in Figure 2.9 have been contoured using the widely used DIPS program (Rocscience Inc. 1999) with the results shown in Figure 2.10. The same discontinuity sets can be identified in the contoured data. A representative or mean orientation of the discontinuity set can then be calculated. The Fisher (1953) distribution is usually used to characterise the distribution of discontinuity orientations about some ‘true’ mean. It is relatively easy to implement and provides a measure of dispersion about the mean discontinuity orientation, called Fisher’s constant, K. It is, however, a symmetric distribution and therefore provides only an approximation for asymmetric data. Watson (1966) and Einstein and Baecher (1983) provide a number of asymmetric models, such as the Bivariate Fisher, which can provide better fits for asymmetric orientation data. A detailed discussion of the Fisher distribution is given in Section 2.6.3. A study by Dershowitz and Einstein (1988) conducted on several distributions using data from a number of sources concluded that none of the currently used distributions were statistically acceptable in all cases. The Fisher, Bivariate Fisher and Bingham distributions provided equal numbers of good fits. The added complexity of asymmetric orientation distributions does not improve the fit of real data to modelled orientation distributions.


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Figure 2.10: Contoured stereographic projection (from Rocscience Inc. 1999)

2.5.4 Discontinuity Frequency/Spacing Analysis

The discontinuity frequency is a fundamental measure of the degree of fracturing of a rock mass. Discontinuity frequency can be expressed as the number of discontinuities observed or predicted within a unit volume (volumetric discontinuity density), a unit area (areal discontinuity density) or a unit length (linear fracture frequency). Discontinuity spacing is a measure that is linked to the discontinuity frequency. At its simplest, the discontinuity spacing is the distance between one discontinuity and another, or the reciprocal of the linear fracture frequency. The volumetric discontinuity density, λv, is the most fundamental of the three measures of discontinuity intensity (Priest 1993). It is based on the assumption that discontinuities can be represented by the occurrence of a point located at the centroid of the discontinuity. The volumetric frequency, λv, is the average number of points per unit volume of the rock mass (m-3). The measure can be applied to all the discontinuities contained in the rock mass, λv, or individually for the density of each discontinuity set, λv

n, where n represents the discontinuity set number. Although λv is an attractive measure of discontinuity intensity, its direct measurement would require the rock mass to be dissected in a non-destructive manner. This is currently impractical. Accordingly, the volumetric frequency must be estimated from areal or lineal density measurements, following methods described by Baecher et al (1977), Warburton (1980) and Villaescusa (1991).

Set 1

Set 2

Set 3


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Discontinuity areal frequency, λa, is the average number of points that occur in a unit area of a section through the rock mass. This parameter can be measured from a sampling plane using a wall mapping technique. The problem of using an areal intensity measure is the need to consider the possible orientation bias of the sample plane. A simple geometric correction introduced by Terzaghi (1965) can be used to relate the true discontinuity frequency to the apparent discontinuity frequency: λ λ γas a= sin (2.1)

where λas is the apparent areal discontinuity frequency and γ is the angle between the mean discontinuity plane orientation and the vector normal to the sample plane. There are also sample biases associated with censoring and truncation effects imposed by exposures of limited extent and issues associated with the shapes of discontinuities. As a result of these complexities, areal discontinuity frequency is not a measure of discontinuity intensity that finds much practical use. The linear fracture frequency is the simplest and most commonly used measure of discontinuity frequency. It is used in the MRMR rock mass classification system (Laubscher 1990), in estimates of in situ block size (Palmstrom 1996) and is recorded in drill core logging. Linear fracture frequency is also often used to estimate the RQD parameter (needed in most classification systems, see Section 2.7) in a correlation proposed by Priest and Hudson (1976). The widespread use of linear fracture frequency as a measure of discontinuity intensity owes much to the use of scanlines and drill core as the major discontinuity characterisation techniques. The linear fracture frequency, λl, will be dependent on the orientation of the sampling line unless the discontinuity network is isotropic, which is unlikely. The linear fracture frequency of a particular discontinuity set n, λl

n, represents the linear fracture frequency of the set perpendicular to the discontinuity plane (in the direction of the mean unit normal). The apparent linear fracture frequency for a particular discontinuity set) λls

n, will be that sampled by a scanline or drill hole oriented in a particular direction (see Figure 2.11). The true and apparent linear discontinuity frequency are related by the Terzaghi (1965) correction given by Equation 2.1 with γ being the angle between the mean discontinuity plane orientation and sample line direction. Using this result and by combining the results of the number of discontinuity sets contained in the rock mass, the estimated rock mass fracture frequency (λ) for different sampling directions can be estimated. Discontinuity fracture frequency data for a rock mass, in particular the fracture frequency extrema (maximum and minimum fracture frequencies), have been investigated by Hudson and Priest (1983) who produced some useful results for obtaining loci of discontinuity frequency.


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True discontinuity set fracture frequency (λl

n) Apparent discontinuity set fracture frequency

(λlsn)

Figure 2.11: True and apparent discontinuity set linear frequency

Discontinuity spacing is a widely used measure of discontinuity frequency. It is used in all the rock mass classification schemes and a number of the techniques used to predict in situ block sizes outlined in Chapter 4. A number of different measures are used to define discontinuity spacing. The ‘Normal Set Spacing’ is the perpendicular distance between sub parallel discontinuities from the same set (the reciprocal of λl

n, see Figure 2.11). The ‘Apparent Set Spacing’ is the spacing between a pair of immediately adjacent discontinuities from a given discontinuity set, measured along a line of any specified location and orientation. The ‘Rock Mass Spacing’ is the spacing between a pair of immediately adjacent discontinuities (regardless of what discontinuity set they belong to) measured along a sampling line. Although the mean spacing of a discontinuity set or of the whole rock mass does provide a useful measure of discontinuity intensity, a greater understanding of rock discontinuity properties can be gained from investigating the full distribution of discontinuity spacing. If a sufficiently large number of individual spacing values are obtained (Hudson and Harrison (1997) suggest more than 200 individual measurements), they can be plotted in histogram form to gain an understanding of the shape of the distribution. On the basis of field measurements, Priest and Hudson (1976) concluded that the distribution of total discontinuity spacings for a variety of sedimentary rock types could be modelled by the negative exponential probability density distribution. This finding has been supported by other investigators (eg Call et al 1976, Einstein et al 1980, Baecher 1983) who worked on a variety of igneous, sedimentary and metamorphic rocks. If the occurrence of a discontinuity along a scanline or drill core is entirely random, then the location of one discontinuity has no influence upon the location of any other. In this case the discontinuity intersections are said to obey a one-dimensional Poisson process. When this


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occurs the probability density distribution of total discontinuity spacings is negative exponential with a mean spacing of 1/λl. The fact that a large number of discontinuity set spacings examined follow the negative exponential distribution suggests, but does not prove, that in a statistical sense discontinuity occurrences are random. Hudson and Harrison (1997) emphasise the fact that discontinuities are not random events, even though the Poisson process of random events can be expected to apply to field data. They suggest that spacing values converge when successive spacing distributions of any type are superimposed on the sampling line. The negative exponential distribution is expected as a result of a suite of superimposed geological events, each of which produces fracturing of a given distribution. Roleau and Gale (1985) similarly analysed an extensive discontinuity spacing data base from drill core from a granitoid rock mass. They evaluated the goodness-of-fit of three statistical models, the negative exponential, the log-normal and the Weibull distribution. The results quite clearly showed that the negative exponential distribution did not fit their data but that a log-normal distribution fitted the data very well. Other researchers (eg Bridges 1975, Barton 1977) have also fitted logarithmic spacing distributions to their observations. In reviewing the work of Rouleau and Gale (1985), Mohajerani (1998) noted how the data points which did not fit the negative exponential distribution had mean spacings values of about one metre or larger. He suggested that the distribution of spacing values may be assumed to be either negative exponential or log-normal, depending on the rock type and spacing range (between the maximum and minimum values). This tends to agree with the the superposing theory of Hudson and Harrison (1997); if there are enough geological events to create a number of discontinuity sets and a small total spacing, then the spacing distribution will follow a negative exponential distribution. Where only a few geological events have caused fracturing, or existing discontinuity sets have become healed, a larger total spacing and log-normal distribution of discontinuity spacings may result. Whether a discontinuity spacing distribution follows a negative exponential distribution or a log-normal distribution also depends on the sampling regime adopted. If a truncation level is adopted (ie discontinuities below a certain size are disregarded) then it is likely that small discontinuity spacings will be lost. As a result, a negative exponential distribution would appear as a log-normal distribution. Inaccuracy associated with a calculated mean discontinuity spacing occurs where the estimated value is consistently in error. One particular inaccuracy is that caused by small sampling scanlines. If the length of the scanline is short compared to the mean spacing then a biased result will be obtained. This type of bias can be produced in areas where only short scanlines can be established. In particular, this is a problem in vertical scanlines that are carried out in drives and tunnels. The effect of short scanlines on mean discontinuity spacing calculations is discussed by Sen and Kazi (1984). They produce graphs that illustrate the effect of scanline length on the calculation of mean discontinuity spacing values for negative exponential and log-


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normal spacing distributions. These graphs can be used to estimate population mean spacing values using the sample mean (and sample standard deviation in the log-normal case), length of scanline and type of distribution. Priest and Hudson (1981) describe a method of calculating the precision of the estimate of the population mean spacing value from the sample of size n, using standard statistical methods based on the central limit theorem. The central limit theorem states that the mean values, X , of random samples of size, n, taken from a population that follows any distribution and has some definite but unknown mean value, μx, and variance, σx

2, will tend to be normally distributed with a mean, μx, and a standard deviation (or standard error of the mean) of σ σx xn n/ ( / )2 .

This central limit theorem technique for analysing discontinuity spacing is particularly useful when applied to the negative exponential distribution because the mean and standard deviation are equal in this case (Priest 1993). It can be used in a number of ways. For example, it could be used to determine for a given sample, the degree of confidence (eg 90% probability) with which it can be said that the unknown population mean, μx, lies within some range of the sample mean, X . Another use would be to specify the given precision of the spacing population estimate that is required and to use this to calculate the sample size that will give the desired precision. Some drawbacks of using the central limit theorem in discontinuity analysis are that the technique is applicable only to discontinuity spacing values (Priest 1993) and that the analysis is still prone to biases such as those associated with small scanlines and orientation bias. Where rock mass discontinuity frequency is anisotropic, the estimate obtained of the range of mean total discontinuity spacing values using a given confidence limit, is only applicable to the orientation for which the sampling was carried out. 2.5.5 Discontinuity Persistence (Size) Analysis

Discontinuity persistence refers to the lateral extent or size of a discontinuity plane. In practice, the persistence of a discontinuity plane is almost always measured by the one-dimensional extent of its trace length on a sample plane. It is clear that no direct estimation of persistence is possible from borehole core, although an estimate of persistence can be made from geological inference (Hudson and Harrison 1997, Henry et al 1999). The distribution of trace lengths obtained from sampling a rock face depend to a great extent upon the degree to which length measurements are truncated and censored. The most commonly used measure of discontinuity persistence in engineering analyses is mean trace length. There are two main problems associated with estimating a mean trace length - precision and accuracy (Priest and Hudson 1981). In the case of precision, it is assumed that as the number of trace length samples is increased, the sample mean trace length will tend towards the true population mean as long as there is no bias in the sampling. The truncation and


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censoring biases common in sampling trace lengths require a correction to allow an accurate measure of the trace length parameter. There has been considerable debate as to whether trace lengths have a negative exponential distribution or a log-normal distribution. The results of a number of investigations carried out on discontinuity trace length distributions are summarised in Table 2.5. It is likely that some of the differences have arisen from trace length sampling bias (Hudson and Harrison 1997).

Table 2.5: Discontinuity trace length distribution and shape characteristics

Reference Trace Length Shape

Robertson (1970) exponential equidimensional

McMahon (1974) log-normal -

Bridges (1976) log-normal oblong

Call et al (1976) exponential -

Baecher et al (1977) log-normal equidimensional

Barton (1977) log-normal equidimensional

Cruden (1977) censored exp. -

Baecher & Lanney (1978) log-normal or exp. -

Herget (1982) exponential -

A number of techniques have been developed to estimate the mean discontinuity trace length. Some of these techniques account for the sampling biases and a number of them require some assumptions for their use. A number of the most relevant contributions are outlined below. Cruden (1977) proposed an original method to estimate the length of censored discontinuities as a function of the observed number of discontinuity end points and the observed discontinuity trace segment appearing above the sampling line (the semi-trace length). The advantage of the semi-trace length approach is that the uncensored distribution of trace length can be obtained without involving a point process of discontinuity trace centres. The problem is that semi-trace lengths are monotonic decreasing functions, insensitive to changes in the underlying trace length distribution (Villaescusa 1991). Warburton (1980) developed a stereological interpretation of discontinuity trace data. He was one of the first researchers to analyse discontinuities in three dimensions rather than two. The statistical model defines the analytical distribution of discontinuity diameters from observed trace length distributions. Unfortunately, the problem of discontinuity censoring is not addressed. Laslett (1982) developed a technique to estimate the parameters of the underlying trace length distribution from line sampling data collected in two dimensions. The technique corrects for bias incurred when incomplete observations form part of a data set. His work was limited to two dimensions. Villaescusa (1991) coupled Warburton’s stereological relationships to


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Laslett’s theory, to give a three dimensional determination of a maximum likelihood estimator of discontinuity size. Kulatilake (1988) developed a corrected equation to obtain unbiased estimates of spacing and intensity from finite scan lines. Procedures were also developed for orientation bias correction for finite size discontinuities intersecting finite size sampling domains, using a variety of discontinuity shapes (circular, rectangle, square and right angle triangle) and sampling geometries (Kulatilake et al 1990). Priest (1993) presented a graphical technique which is trace length distribution independent. This involves constructing a histogram of semi-trace lengths and drawing a best fit curve through the midpoints of the class intervals. Using the intercept of the graph, sample size and histogram class interval, an estimate of mean trace length can be obtained. The problem associated with this technique is that different interpretations of the histogram shape (by choosing different size class intervals) will result in different estimates of the underlying mean trace length. Mauldon (1998) provided new estimators of mean discontinuity trace length and density that correct for the effects of bias and censoring. A stereological estimator of mean trace length was developed which requires rectangular windows and parallel traces, and is similar to earlier methods which use a stereological methodology. Another estimator called the end-point estimator of mean trace length can be used in any convex window with variably oriented traces. This method is independent of the underlying trace length distribution in all cases and is independent of the trace orientation distribution when applied to circular sample windows. Zhang and Einstein (1998) developed a technique for estimating mean trace length from observations made using finite, circular sampling windows. It uses information on the number of end points observed in the circular sampling window and the sampling window dimensions. The advantage of the technique is that the trace lengths, the underlying distribution of trace lengths and discontinuity orientation measurements are not required. The disadvantage of the technique is the practical difficulty associated with sampling using a circular mapping window. This brief review shows that a great deal of progress has been made in increasing the accuracy of mean trace length estimates. However, little progress has been made in calculating the precision of these mean trace length estimates, perhaps because of the sampling biases inherent in the estimation of mean trace length. 2.5.6 Definition of Geotechnical or Structural Domains

Geotechnical domains are essentially regions of structural homogeneity. These regions are identified as areas of the rock mass that contain discontinuity characteristics that are more or less structurally and statistically homogeneous. If the heterogeneous ‘whole’ can be divided into homogenous parts then engineering analysis can be carried out for each design region.


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In the pre-feasibility stage of a project, before any comprehensive site investigation has been carried out, the amount of structural data available may be limited. Usually, the information collected will have concentrated on the lithologies present and the major structures contained in the region of interest. This information will be used to define the initial structural domains in the rock mass. Initial structural domain boundaries will usually be identified from: • geological boundaries between rock types; • major faults; • data from different levels of an underground mine; and • major changes in weathering and alteration (eg unweathered rock mass and the near surface

weathered zone). In the initial analysis, it may be apparent that the rock mass discontinuity parameters found in one domain are very similar to those found in an adjacent domain. In such a case, there is no reason for maintaining the boundary between the structural domains which can be combined. Alternatively, with additional data it may become apparent that two or more zones of different rock mass discontinuity parameters may exist in the same ‘initial’ structural domain and further subdivision of the domain will be warranted. To exhibit statistical structural homogeneity a domain should have the same number of discontinuity sets and each identified discontinuity set should display similar distributions of orientation, spacing, trace length, termination index and discontinuity conditions. Quite frequently, a comprehensive set of discontinuity data containing all these parameters is unavailable. In particular, it is the number of discontinuity sets and the orientations of these sets that are used to define structural domains. The determination of homogeneous domains is often performed by visual estimation. Although subjective, visual estimation can be a reliable technique especially where an individual is familiar with the geology of the site (Bridges 1990). If visual estimation techniques are found to be inadequate, a statistically based technique may be used to cluster orientation data. An advantage of using numerical techniques is the implicit objectivity of the approach. The main disadvantage arises from the difficulty in applying statistical techniques to orientational data, particularly because of the problems associated with sampling. In order to compare two regions to establish if they are statistically structurally homogeneous, the sampling of the two regions should be identical. Even then, because of the possible poor precision in estimates of discontinuity characteristics, it is possible that identical regions may be identified as being statistically heterogeneous. 2.5.7 JK Jointstats Discontinuity Data Management System

The JK Jointstats discontinuity data management system provides the tools required to perform three important functions necessary in providing discontinuity data for subsequent engineering analysis, namely, discontinuity data input, discontinuity selection and discontinuity set definition. Once these tasks have been completed the front end program is used to initialise the appropriate analysis module. The central discontinuity database uses a Microsoft Access


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database. The widespread industry use of Microsoft Access facilitates the transfer of data to the discontinuity database and allows the data to be exported if required. Discontinuity data input

Discontinuity data can be entered in different modes. Several modes are predefined and included in the program, including a number of mapping and core logging techniques. The minimum data entry mode involves only the intercept distance along the line and basic orientation data. A user-defined mode allows the user to select which discontinuity parameter fields will be used in mapping. Clearly, the data entry mode used will have great impact on the viability of the statistical calculations and models. For example, to use advanced statistical tools a measure of discontinuity trace length is required. If an attempt is made to use this module with limited discontinuity data that does not contain a measure of discontinuity trace length, a message is generated warning that such an option is unavailable. Discontinuity data can also be imported from Microsoft Excel via the MS Windows clipboard. An example of discontinuity data entry is shown in Figure 2.12. This example uses the JKMRC full scanline mapping method. Data recorded in the field can be either added directly by keyboard or if it exists in electronic form (as in field based hand held computers) it can be input via an MS Excel spreadsheet. It is possible for the user to develop mapping templates to satisfy project requirements as in the example shown in Figure 2.13. Firstly, when selecting the data collection method the ‘User Defined’ option is selected. Then the discontinuity parameters that are routinely recorded during logging or mapping are added to a blank template to create the active data sheet.

Figure 2.12: JKMRC full scanline data entry


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Figure 2.13: Creation of a user defined mapping method

Discontinuity selection

Discontinuity data are organised in a hierarchical manner and presented to the user in a tree view (as in MS Windows Explorer). The tree view represents an underlying database structure that stores the objects involved in discontinuity data collection in their logical relationship. These objects can be summarised as: • mine; • mining unit (eg pit, panel, block or bench); • domain (a logical grouping of discontinuity data); • plane (the face from which the data were collected); • scanline (the line along which data were collected); and • discontinuity.

Note that cores sit at the same level as planes in the structure. This structure comprises a series of one-to-many relationships (eg one mine can contain many mining units) and referential integrity is enforced (when a scanline is deleted, all the discontinuities that belong to it are automatically removed). A typical tree view with an open menu is shown in Figure 2.14.


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Figure 2.14: Data tree view

A key issue is the definition of a domain. A domain is a logical (rather than a geographic) grouping of discontinuity data. It is a region of the rock mass for which discontinuity statistics are accepted as being statistically homogenous (same orientations, size and frequency) for engineering applications. When a new mining block is created, it is automatically assigned a single new ‘hold-all’ domain. As data are collected, natural groupings will become apparent. The hold-all domain can then be split into more meaningful domains with descriptive names. Planes and cores can be cut and pasted between domains. It would be irrational to perform a statistical analysis of all the discontinuity data contained in the database, as it could contain many mines and a number of different structural domains. Discontinuities at any level from domain downward can be ‘selected’ to partition the data. Selection is achieved by clicking on the appropriate toolbar button or by toggling the selection item in the context menu. The context menu is shown after right clicking on a highlighted object and can be seen in Figure 2.14. This means that the user can toggle the selection of a given domain ‘on’ and automatically select all the discontinuities in that domain. It is possible to then manually deselect individual scanlines in the same domain to fine-tune the selection. The filter builder tool in the software package can be used to further investigate the discontinuity data for properties other than orientation (Figure 2.15). For example, it may be important to segregate the data by discontinuity type. Using such a filter the differences


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between faults, joints and veins could be assessed. Another useful exercise would be to filter the trace length parameter. Some engineering applications require only large scale discontinuities to be taken into account, so a filter could be applied to identify only those discontinuities having a discontinuity size above a set value.

Figure 2.15: Discontinuity data filter builder

Since selection is not permitted across domains, any new selection that violates this rule generates a warning message. The user can then choose to replace the old selection with the new or abort the process. The selected data can then be plotted on a stereographic projection. Polar plots of discontinuity poles can be drawn in lower or upper hemisphere projection and in equal area or equal angle space.

Discontinuity set definition

The stereographic plot is then used to define the different discontinuity sets that exist in the rock mass. This is done by clicking on a start point, holding down the mouse button, and dragging a selection mask clockwise and outwards around the discontinuity poles. Discontinuity poles that fall inside the selection mask will turn from white to black. An option to show the Fisher distribution during set definition is provided. When the selection mask crosses the plot perimeter (ie discontinuity planes cross 90°), the mask will flip over to the ‘antipole’ region and the Fisher distribution corrected accordingly. This is equivalent to merging upper and lower hemisphere information and is valid because poles have no real termination. This transition is shown in Figure 2.16.


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Figure 2.16: Set definition with context menu

Sub-horizontal sets are selected using a circular selection mask which opens from the pole. The transition from a circular to a segment selection mask is determined by the user as one of the program options. Once the set has been defined, right clicking brings up a context menu that includes an option to store the set parameters. Selecting this menu option will open the set manager form. The user is required to enter a colour for the set (so that its discontinuities can be readily distinguished in subsequent plots) and a set name. If no set name is entered, the unique ID number for that set is stored in the name field as a default. The same set manager form is also used ‘offline’ to edit or delete already established sets which are then selected from a drop-down choice box. The data set manager is shown in Figure 2.17. A variation on the data plot can be displayed along side the set manager as different sets are selected. In this mode, the plot will show only those discontinuities assigned to the current set in the stored set colour, together with the set definition mask.


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Figure 2.17: Discontinuity data set manager

Photograph gallery manager Best practice dictates that images of logged core and mapping faces are obtained during the data collection phase. The best method of storing such data is in an electronic form which allows images to be observed while the collected data are being analysed. This methodology has been incorporated into JK Jointstats by using a digital gallery manager that is linked to the discontinuity database. Images associated with the appropriate scanline or core can be easily examined using the gallery function. A number of thumbnail pictures that contain a description of the images can be viewed at once when the gallery option is selected. An example of core photographs obtained at the Newcrest Ridgeway Project is shown in Figure 2.18. The appropriate tile is simply selected to obtain a close up view of an image.


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Figure 2.18: Discontinuity photograph gallery Once discontinuities have been selected and assigned to sets, the program can run simple statistical calculations on the information. These are shown on a multi-tabbed form containing tables and histogram plots. Any plot can be copied to the clipboard, printed or dumped to file. Any table can be transferred to MS Excel via the clipboard. A number of simple statistics are available as outlined below. Rock mass spacing (average discontinuity spacing in the rock mass)

The discontinuity spacing for all the selected discontinuities, whether assigned to sets or not, is displayed on the second tab of the Basic Statistics module (see Figure 2.19). This gives an impression of the overall discontinuity spacing in the rock mass and an appreciation of the potential in situ block size. The average fracture frequency in the rock mass is also estimated.


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Figure 2.19: Rock mass discontinuity spacing

Corrected discontinuity spacing by set

Given that the orientations of the discontinuity planes are known and the mean orientation of each set has been calculated, the apparent (measured) discontinuity spacing can be corrected for relative orientation and plotted for each set. This apparent spacing can then be converted to a true or corrected spacing. The statistics of the discontinuity set spacings are calculated and the results are shown graphically as illustrated in Figure 2.20. Using the pull-down set selector near the top of the form (see Figure 2.20), it is possible to review and compare the results for the discontinuity sets selected in the analysis.


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Figure 2.20: Corrected discontinuity spacing per set Trace lengths by set and for random discontinuities

Statistics for the recorded trace lengths can be displayed and plotted for each set and for those discontinuities selected but unassigned to sets (termed ‘random’ discontinuities). An example for random discontinuities is shown in Figure 2.21. The plot takes the form of a stacked histogram, the elements of which are designed to show a qualitative assessment of ‘confidence’. This is necessary because discontinuity trace lengths can be censored by the observation window limits at either the top or bottom (shown in grey shades) or at both the top and bottom (shown in black). Clearly, the censored discontinuities will tend to occur at the greater trace lengths (depending on the window limits) so the data are less meaningful towards the right of the plot. However, this remains potentially useful information because it is known that such trace lengths fall outside the lower bins. Thus, it is important to show such data as long as some impression of confidence is included.


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Figure 2.21: Trace length statistics for ‘random’ discontinuities In the Basic Statistics selection the censoring classes provide a qualitative measure of confidence in the trace length statistics. The information on the censoring class of the discontinuities is vital when the advanced statistics option is selected as it is used in the estimation of the mean size parameter of the discontinuities. It is only by utilising the advanced option that a quantitative measure of confidence on discontinuity parameters can be obtained. 2.6 SIMULATION OF ROCK MASS GEOMETRY

2.6.1 Introduction

The simulation of a rock mass geometry involves the construction, with the aid of a computer, of a graphical representation of the rock joints within the rock mass. The graphical representation may be a 2-D section through the rock mass, effectively showing the traces of the joints on a face, or it may be a full 3-D model in which the assembly of joints may be viewed from various angles or the observer may ‘fly’ through the rock mass, inspecting the joints and their intersections. Simulations of a rock mass must be based on some conceptual model of the rock joints and the manner in which they occur in the rock mass. The conceptual model effectively provides the rules according to which the joints are placed into 3-D space. Thus the geometric features of


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the simulation and any geomechanical consequences that they may have are a function of the model and will be only as valid as the model itself. A number of models may be conceived of as being plausible (Dershowitz 1993). The extent to which any one model actually mimics nature must be established by the collection of data and comparison of the data with the predictions of the model after it has been matched to the data as well as possible. There are a number of reasons why it is useful to be able to simulate a rock mass geometry. Firstly, it is an aid to the visualisation of the joints in the rock mass. Given the belief or even reasonable evidence that there are particular joint sets in the rock mass at particular densities and with particular size and orientation distributions, the use of the computer to view the inter-relationships of the joints allows the engineer to confirm mental constructions of the joints. The interaction of joint sets in a rock mass is a complex process. The complexity derives not so much from the model as from the fact that there are many joints of many possible orientations and sizes. And, since it is generally taken that the jointing is a stochastic phenomenon, the possibilities for joint interaction are effectively infinite. The interaction of joints in the rock mass controls the formation of rock blocks. The existence of discrete, fully formed blocks is of vital interest for the design of reinforcement or for the design of block caves. The in situ fragmentation (the size distribution and volumetric concentration of formed blocks prior to the commencement of mining) is important not only for assessing the stability of the rock mass once mining commences but also for the purpose of estimating the size distribution of the ore at the draw points following secondary fragmentation. The 3-D network of joints also defines pathways for fluid flow through the rock mass. Simulation of discontinuities in a rock mass, in general, is a computational problem of varying tractability; the level of tractability depends on the rules chosen to define the way in which the discontinuities occur. The random disk model is perhaps the simplest of all models as all disks are independent. The more complex the rules, the more difficult is the program simulation and the longer the execution times to complete a simulation in a given rock mass volume. 2.6.2 Approaches to Discontinuity Modelling

Introduction

Models of rock discontinuity networks are a fairly recent development in rock engineering. The first conceptual models such as Snow’s (1965) orthogonal model were developed as relatively simple tools for hydrological modelling. More advanced methods such as the Poisson location models (Baecher et al 1977), tessellation models (Veneziano 1978), hierarchical methods (Lee et al 1990), fractal approaches (Barton and Larson 1985) and geostatistical methods (Gervais et al 1995), have allowed more realistic representations of in situ discontinuity network geometries to be created. Some of these approaches are outlined below.


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Random disk (Poisson location) discontinuity models

The fundamental feature of the random disc model (Baecher et al 1977, Warburton 1980, Villaescusa 1991, Hadjigeorgiou et al 1998) is the assumption of circular or elliptical discontinuities. The size of circular discontinuities is defined completely by a single parameter, the discontinuity radius Rj. The discontinuity radius may be defined deterministically as a constant for all discontinuities, or stochastically by a distribution of radii fr(Rj). Appropriate distributional forms for discontinuity radius include the exponential and lognormal distributions, both of which produce lognormal distributions of discontinuity trace length (Baecher 1983). Discontinuity location may be defined by a regular (deterministic) pattern or by a stochastic process. The simplest stochastic assumption is a Poisson process, in which discontinuity centres are located randomly and uniformly in space. Discontinuity orientations may be defined by any orientation distribution, or by a constant orientation. As a result of the discontinuity location, shape and size process of the Baecher model, discontinuities intersect each other and terminate in intact rock (Dershowitz and Einstein 1988). An example of this model is shown in Figure 2.22.

Figure 2.22: Random disk model (Dershowitz and Einstein 1988)

The mutual independence of discontinuities is the biggest disadvantage of the random disk model, in that the discontinuity termination often seen in rock mass exposures cannot be modelled. Interestingly though, this independence of discontinuities is also the source of the greatest strength of the random disk model. It permits the application of a number of statistical techniques in conjunction with the rock mass model which rely upon the independence of


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elements. This allows a link to be developed between two dimensional discontinuity properties measured in the field (spacing or frequency and trace length) and those three dimensional discontinuity properties modelled (intensity and size). Although the model was initially developed using circular disks, there is no reason why other shapes cannot be used to represent discontinuities. The Poisson location seeding and orientation process remains unchanged. However, the shape of the discontinuity does affect the analytical techniques developed to relate trace length properties with the discontinuity size properties but this can be overcome by using a forward modelling technique. Random coplanar polygon discontinuity models

Random coplanar polygon discontinuity models use a tessellation approach to model rock mass discontinuities. A tessellation is a division of a plane into polygons, or of space into polyhedra. When discontinuities are sufficiently connected to produce completely defined rock blocks, a tessellation approach may be appropriate. The generation of a Veneziano (1978) rock discontinuity system model requires three consecutive stochastic processes: • discontinuity planes generated as infinite Poisson planes, distributed in space by a uniform

distribution with any distribution of orientation allowed; • a Poisson line process on each discontinuity plane divides discontinuity planes into

polygonal regions; and • a portion, PA, of these polygons is randomly marked as jointed, while the remainder are

marked as intact rock where PA corresponds to persistence. Discontinuities are modelled using polygonal shapes, and discontinuity sizes are defined by the intensity of the Poisson line process and the proportion of the polygons marked as discontinuities. The Veneziano model resembles the Baecher model, except that discontinuities are represented by coplanar line segments rather than independent circles (lines in two dimensions). In the Veneziano model, an independent Poisson line process defines discontinuities in each plane. This process means that discontinuities in one plane are completely independent of discontinuities in adjacent planes, so termination at another discontinuity cannot occur except by chance (Dershowitz and Einstein 1988). One of the problems associated with the Veneziano model is in the production of in situ blocks. Discontinuities are defined as Poisson lines on previously defined Poisson planes. Intersections between discontinuities on different discontinuity planes, therefore, do not often match discontinuity edges. Rock blocks can be created with Veneziano models if the discontinuities are 100% persistent and unbounded but the resulting infinite persistent discontinuity planes provide an unrealistic assumption for the modelling a rock mass.


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Figure 2.23: Dershowitz random coplanar polygon discontinuity model (Dershowitz and Einstein 1988)

The development of the Dershowitz (1984) model overcame some of the problems of the Veneziano model. Dershowitz locates the discontinuity planes in space in a manner similar to that used in the Veneziano model but the polygons are formed by the intersections of the Poissonian flats (see Figure 2.23). A portion of the polygons are marked as broken and the remainder as intact rock. This model improves the previous one, since all the discontinuities can terminate against discontinuity edges. Although this model can define distinct blocks at any scale (block faces are either completely broken or intact), an increase in the Poissonian flat density may produce a problematically large number of small polygons (Dershowitz and Einstein 1988). Dershowitz has further advanced tessellation models to develop porosity analyses of fractured rock masses (Dershowitz 1993). An advantage of the Dershowitz (1984) model is that distinct rock blocks are defined. Another advantage is that by using a flexible orientation distribution a variety of polygonal discontinuity shapes and polyhedral block shapes can be modelled. It is possible to create discontinuities terminating in intact rock (and hence intact rock bridges) by modelling with a virtual discontinuity set which has zero persistence. Random coplanar discontinuity polygon models do not model the inherent statistical properties of discontinuities (size, shape, intensity and orientation) which is a significant disadvantage. Another disadvantage outlined by Dershowitz and Einstein (1988), is the fact that polygonal block face sizes are controlled by the intensity of intersecting discontinuity plane processes. As


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the intensity of the plane process increases, the number of intersecting lines on each plane increases, and therefore the size of the polygons defined by the line decreases. If joints are defined as a constant percentage of each plane, the increase in plane process intensity results in a larger number of smaller polygons. Fractal based discontinuity models

Fractal geometry is a way of quantifying the spatial or temporal dispersion of a quantity, and how this dispersion may change with the scale of observation. Its attraction for discontinuity pattern characterisation lies in its simplicity, the less rigorous data requirements (than geostatistical approaches) and its correspondence to certain geometric features in natural discontinuity patterns. An early application of a fractal based discontinuity model was a two dimensional model of jointing at the proposed nuclear waste repository at Yucca Mountain, USA (Barton et al 1985). One of the characteristic features of fractal models is that large discontinuities are not often found near other large discontinuities; rather, they are spaced at large intervals, so that, on average, the size of the fracture is proportional to the spacing. This autocorrelation of discontinuity trace length would be picked up using geostatistical models of trace length semi-variograms. Empirical evidence does not unequivocally support the existence of autocorrelation, so to use a model whose main strength is the ability to model autocorrelation of discontinuity properties may be pre-emptive until further study in this field is undertaken. The fractal nature of a discontinuity network is examined using discontinuity trace maps. The reliance on trace maps to characterise the fractal nature of the discontinuity network may limit their use in cave mine engineering, due to the extensive use of core logging as a source of discontinuity characterisation information. Although promising, the application of fractal geometry for engineering design purposes is largely untested. For example, no verification exists that fracture networks derived from fractals produce model input that predicts the behaviour of a flow experiment any more accurately than other methods (La Pointe 1993). Hierarchical discontinuity models

In hierarchical models the discontinuity sets are described and modelled in hierarchical order to account for dependencies among discontinuities of the same set or of different sets. The Hierarchical Fracture Trace Model of Lee et al (1990) was one of the earliest hierarchical models. The sequential generation and correlation of discontinuity sets corresponds to what happens in nature. The main features of the model are: • The spatial variation of trace density is represented through a double stochastic point

process. This class of procedure is quite versatile and is especially appropriate when the variation of trace density is due to external factors (eg state of stress, rock strength).


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• A new method, based on maximum likelihood theory, is presented for the unbiased estimation of trace length distribution using outcrop data.

• When several trace sets are present, these sets are analysed sequentially, according to a hierarchical (ideally, chronological) order. Each set is represented through a conditional stochastic process, conditioning being with regard to the lower-order sets. This is perhaps the most innovative feature of the model.

• Methods from multivariate point processes are adapted to the estimation and validation of line-segment processes in tessellation.

The model contains several interesting features that would be desirable in a rock mass model for cave mine engineering. Unfortunately, the model would have to undergo further development before becoming a useful caving rock mass characterisation tool. The advances required would include extension of the model into three dimensions and the analysis of patterns with more than two trace sets for structurally complex rock masses. A recent hierarchical model specifically designed to model real geological fracturing processes, is the MIT geologic stochastic fracture model (Meyer et al 1999). The model is hierarchical since the fractures produced are grouped into hierarchically related fracture sets. It uses tessellated Poisson planes which have been subdivided into fractured and unfractured rock, and uses the geological history of the discontinuities to recreate the geometric network seen in the field. Although applied to large scale discontinuities, namely crustal scale shears and large scale faulting, it nevertheless applies some novel techniques to obtain a good representation of fracture geometry. In the example discussed by Meyer et al (1999), four stages of faulting are modelled progressing from a simple to a compound arrangement of strike-slip fault zones. Probability is used to determine whether a new fault crosses through a previously existing fault when an intersection occurs. Geostatistical discontinuity models

In reviewing traditional stochastic discontinuity geometry models, Gervais et al (1995) found that discontinuity patterns observed in real rock masses were not adequately recreated in the models because of the variable discontinuity density and clustering of discontinuities occurring in nature. They proposed a geostatistical model of discontinuity geometry. This model features a hierarchical model of discontinuity networks, utilising a statistical characterisation and a geostatistical approach. Semi-variograms were developed to describe the spatial behaviour of the regionalised variables of the discontinuity network, and thus of the underlying random process. Semi-variograms of cumulated length of discontinuity traces per square meter were computed to study the spatial correlation of discontinuity density. For each discontinuity set, the variogram shows a structure along the direction of the mean orientation of the set. Spatial correlation of discontinuity intensity exists in this orientation. This is because well defined


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corridors of discontinuities line up with each discontinuity set. Variograms constructed perpendicularly to the direction of the mean of the set, show a pure nugget effect with no structure. To model the geometry of the discontinuity network, the first discontinuity set was generated with the classical method (random distribution of discontinuity centres, log-normal distribution of trace lengths), after which the other sets were generated successively and conditionally upon the previous ones, in order to reproduce the geometrical relations between sets and the discontinuity termination types observed in the field (Gervais et al 1995). The Gervais et al (1995) model was developed on discontinuity trace sets of a bedding plane in a limestone quarry, where extensive discontinuity trace maps could be used as the basis of discontinuity characterisation. As a result, the study is highly detailed but entirely two dimensional. Although geostatistics can be an extremely useful tool in characterising the spatial variability of discontinuity intensity, it must be applied in three dimensions to allow for its application in underground mining. However, it is even more important to realise the limits imposed by the generally poor sampling of rock masses in cave mining operations, which will rule out the possibility of developing accurate semi-variograms of discontinuity set parameters. Forward modelling techniques

Forward modelling is a method of obtaining parameters for constructing three-dimensional discrete models of discontinuity systems. In these models, each discontinuity is represented individually and has unique properties (as in the Baecher random disk model). The properties of the model are derived from iterative conditioning to observed data (La Pointe et al 1983). The forward modelling approach is illustrated in Figure 2.24. The method used to collect the structural data is simulated together with the biases described previously. This results in a simulated set of field data which can be compared directly to the field measurements. The goodness of fit between the field and simulated data can be evaluated visually, by statistical comparison, or by the use of statistical tests such as the χ 2 and Kolmogorov-Smirnov tests

(Dershowitz 1995). Based on the results derived from the comparison, the statistical description of discontinuities (orientation, size, shape, spatial distribution) can be modified and the process repeated to obtain a satisfactory match between field and simulated data. Knowing that modelled discontinuity measurements match the in situ discontinuity measurements provides greater confidence in the discontinuity model analyses. Dershowitz (1995) also illustrates other advantages of forward modelling approaches over mathematical approaches such as that used by Villaescusa (1991). The most important is the ability to directly account for known bias, censoring and truncation processes. The forward modelling approach directly simulates the methods of data collection, and therefore automatically accounts for these biases.


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CompareSimulatedand Field

Measurements

In-situMeasurements

Simulated In-situMeasurements

Assumed GeometricProperties

Accept PossibleConceptual Model

GeometricProperties

Good match

Adjust assumptions forbetter match

Generate additionalalternative conceptualmodels if appropriate

Similate process of measurement used in-situ

Figure 2.24: Forward modelling approach (adapted from Dershowitz 1995)

2.6.3 The Development of the JKMRC 3-D Discontinuity Model

This account of the JKMRC 3-D discontinuity model developed as part of the ICS Stage I, is based on the reports of Lyman (2000) and Harries (2001). The simplest model of rock joints that does not have infinitely persistent joints is the random disk model which is a Boolean Random Set (BRS) model. The term Boolean set is used because a set element is or is not present at a point in space. Set elements are placed entirely independently at Poisson points in space and are allowed to overlap or interpenetrate freely. Consider the simple example of spheres placed randomly in space. Having first decided on a volume of space in which the random set is to be constructed, the space is populated with Poisson points of a chosen spatial density, or number of points per unit volume. A set element is then placed at each one of these points. Assume that the volume is a cube 10 m on each side and the spheres have a constant radius of 1 m as shown in Figure 2.25.


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Figure 2.25: BRS of spheres in a cubical volume If the simulation is to be entirely valid within the cube, the sphere centres have to be located in a volume slightly larger than the cube to avoid edge effects. In fact, to consider sphere centres, the volume is found by extending the cube by one sphere diameter in all directions. The easiest method in this case is simply to consider a cube 12 m on edge, centred on the 10 m cube. By definition, if Poisson points have a spatial density of ρ (points per cubic meter), then the probability of finding a point in a small volume dV at an arbitrary location is dVρ . The expected number of points to be found in a finite volume V is

Vn Vρ= (2.2)

and the actual number found in V follows a Poisson distribution with expected value Vn . The

numbers of points occurring in different volumes in the space are mutually independent. If the spatial density chosen is say 2 m-3, the correct way to put the points into the 12 m cube is to first generate a Poisson random number having the expected value

2 1224

Vn = ×=

Say this random number turns out to be 29. Choosing 29 triples of random numbers that follow a uniform distribution between 0 and 12 provides the coordinates of the points. The random number generator used for this has to be a good one (see Press et al 1992).


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The centres of the spheres are placed on the points. This is now a realisation of the random set. Any number of independent realisations can be generated and the extent to which the spheres make contact can be investigated for each realisation and an average measure of the number of touching spheres determined by averaging over many realisations. The set elements used in the above example had a constant diameter; it would be quite legitimate to define the sphere diameters according to some statistical distribution. The distribution could be a discrete one with a particular average or expected proportions of, say, four different diameters. Alternatively, the distribution could be a continuous one quantified by a probability density function ( )g D such that the probability of finding a sphere of diameter D to D dD+ is ( )g D dD . To generate a realisation of spheres whose diameters follow such a distribution, the probability distribution function corresponding to ( )g D is calculated, providing the probability that a sphere has a diameter smaller than a given value. This probability is

( ) { }PrG D D D′ ′= <

( )0

D

g x dx′

= ∫ (2.3)

The function ( )G D must rise monotonically from 0 to 1. Therefore, a sphere of diameter iD can be chosen from ( )g D by choosing a random number ir uniformly distributed between 0 and 1 and using the function ( )G D as shown in Figure 2.26. To make the random set with the sphere diameter probability density ( )g D , simply choose a sphere diameter as illustrated in Figure 2.26 before placing the point in space. If disks are used as the random set elements, an additional factor comes into consideration, namely the disk orientation. The sphere is a completely symmetric object and so has no detectable orientation. A very thin disk has only one axis of symmetry and so needs two numbers to specify its orientation.


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D

( )G D

0

1

ir

iD

Figure 2.26: Illustration of a probability distribution function and the sampling of values from a specific density function

As has been noted previously, the orientation distribution most commonly used for jointing is the Fisher distribution (Fisher 1953). The Fisher distribution is the counterpart of the normal (Gaussian) distribution that is suitable for the topology of the sphere. Figure 2.27 shows a unit sphere and the two angles necessary to specify a point on its surface.

θθ

δ Nθ

Ex

y

z

Figure 2.27: Unit vector from origin to surface of unit sphere showing definition of angles

A unit vector from the origin to a point on the unit sphere can be defined by the angle θ between the North-directed y-axis, measured clockwise when looking from above and the angle δ between the z-axis and the vector. The vector is then defined as

ˆ ˆ ˆsin sin sin cos cos0

δ θ δ θ δ π θ πδ π

= + + − < ≤≤ ≤

u i j k (2.4)


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The orientation of a disk is completely defined by the direction of its unit normal vector. Equation 2.4 can be used to define the normal. In fact, because a disk also possesses a plane of symmetry, it can define either an upward- or a downward-pointing vector. For joint orientations, a convention of downward-pointing vectors is adopted and these vectors are represented on the lower hemisphere stereoplot. The Fisher distribution is defined on the entire sphere in terms of a mean vector direction μ and a dispersion value K. The probability density function for the Fisher distribution is

( ) ( )( ), , , exp , sin4 sinhFisher

KK d d K d dK

φ δ θ δ θ δ θ δ δ θπ

= ⋅μ n μ (2.5)

where ( ),δ θn is a vector in a direction defined by angles ( ),δ θ on the unit sphere and

( ),μ μδ θμ is the average direction of the vectors defined by angles ( ),μ μδ θ . Note that

( ) ( ), , cos nμ μδ θ δ θ δ⋅ =n μ (2.6)

or the cosine of the angle between the vector n and the mean direction μ . It is very important

to recognise that this distribution is defined on the entire sphere and not on just the upper or lower hemisphere. The vectors are illustrated in Figure 2.28.

Figure 2.28: Vector n, a sample from a Fisher distribution with mean direction μ

It has been shown by Lyman (2000) that the expected value and variance of cos nδ for the

Fisher distribution are given by

{ } 1cos cothnE KK

δ = − (2.7)


90

{ } 22

1cos 1 cothnVar KK

δ = − + (2.8)

For large K, coth 1K → quite rapidly so that

{ } 2

1cos nVarK

δ → (2.9)

Equation 2.9 shows how K provides a measure of spread of the distribution. The Fisher distribution is plotted in Figure 2.29.

0

1

2

3

4

5

6

7

0 30 60 90 120 150 180

Angle δ [degrees]

Den

sity

100

50

20

10

5

1

Figure 2.29: The Fisher probability density function for a series of values of the Fisher constant K

If the Fisher distribution is modified to account for a non-symmetric distribution of the normal vector directions, it can be simply multiplied by an appropriate density function in θ to arrive at, for example,

( ) ( ) ( ), , , exp cos sin2sinhn n n n

Kd d f K d dK

φ δ θ δ θ θ δ θ δ δ δ θ= (2.10)

A completely general form of the orientation density function is

( ) ( ) sin, , , ,4n n n nd d b d dδφ δ θ δ θ δ θ θ δ δ θπ

= (2.11)


91

where ( ), 1b θ δ = for a uniform distribution on the sphere.

To sample from the Fisher distribution, the distribution function for the Fisher constant is required. If r is a random number uniformly distributed on the interval 0 1r≤ ≤ , then

( )21cos ln sinh sinh 1n r K r KK

δ ⎡ ⎤= + +⎢ ⎥⎣ ⎦ (2.12)

is a Fisher deviate. The angle nθ must also be specified, and any convenient reference

direction from which to measure nθ may be chosen. A value of nθ can be generated as

( )2 1n rθ π= − (2.13)

where r is a second uniform random number. Choosing the plane defined by μ and a purely

vertical vector as a reference for nθ , the Fisher normal can be rotated back to the original

coordinates by the calculation

1

2

3

cos cos sin sin sin sin sinsin cos cos sin cos sin cos

0 sin cos cos

n n

n n

n

nnn

μ μ μ μ μ

μ μ μ μ μ

μ μ

θ δ θ δ θ δ θθ δ θ δ θ δ θ

δ δ δ

⎡ ⎤⎡ ⎤ ⎡ ⎤⎢ ⎥⎢ ⎥ ⎢ ⎥= −⎢ ⎥⎢ ⎥ ⎢ ⎥⎢ ⎥⎢ ⎥ ⎢ ⎥−⎣ ⎦ ⎣ ⎦⎣ ⎦

(2.14)

These relatively simple calculations are all that is needed to generate a set of vectors coming from a Fisher distribution. Note that scanline orientation bias is not included. These procedures are sufficient to create a BRS realisation of random disks following a Fisher distribution in a chosen volume of space. The line of intersection between a plane within the simulation volume and any one of the disks may then be determined. The intersections between a line representing the axis of a core and the set of disks generated may also be found. The basic BRS model can also be varied to use shapes other then disks for the joints if desired. Lyman (2000) has applied maximum likelihood theory to the BRS model to produce a number of important results. Lyman’s approach is statistically rigorous, allowing all geometric and censoring biases to be accounted for exactly in establishing values of parameters such as joint size and density. The likelihood distribution functions for joint parameters may be determined by using fitted parameter values and trial data sets.


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The direct use of the likelihood distribution function permits confidence intervals to be established for estimates of joint parameters. The application of this modelling approach has allowed the impact of additional scanline sampling on parameter confidence intervals to be tested and the impact of data quantity and quality on the parameters used in cave design to be assessed (Harries 2001). In principle, the approach is also capable of being used to develop an automatic, statistically based clustering method to sort joint orientation data into sets (Lyman 2000). 2.6.4 The JKMRC Hierarchical Model of Discontinuity Network Geometry

The geometrical properties of the rock fractures that have been modelled in the JKMRC 3-D rock fracture model (see Section 2.6.3) include:

• discontinuity orientation; • discontinuity shape; • discontinuity size; and • discontinuity volumetric density.

The first three parameters wholly describe the geometries of individual discontinuities. Discontinuity volumetric density relates to the density or intensity of fracturing of a given set. However, these are not the only discontinuity parameters needed to describe fully the geometry of the discontinuity network. There are also discontinuity network parameters such as termination style, degree of interconnection and hierarchy seen in the network. Figure 2.30 illustrates some of these network parameters schematically.

Case A Case B

Figure 2.30: Influence of fracture network parameters on geometry The two fracture trace maps shown in Figure 2.30 have identical individual fracture geometry statistics. The fracture sets in both cases have matching orientations and sizes, only their locations differ. However, it can be seen clearly that the fracture network geometries for the two cases are quite different. The distributions of primary fragments or blocks defined by the


93

fracture traces are very different. In Case A, the fractures have been created as independent events with no mutual interaction (as in the JKMRC 3-D model). Case B has a definite hierarchy in that the fractures do interact with each other. These interacting traces more often fully define polygons in the resulting simulation. Case B illustrates the influence of fracture termination on the resulting discontinuity network geometry. The statistical procedures used for estimating fracture parameters in the 3-D model provide the starting point for the hierarchical modelling exercise. However, given the statistical (and independent) nature of the rock discontinuity parameters, an infinite number of rock discontinuity geometries are possible. This may result in the creation of a rock mass model in which the fracture parameters are statistically correct when compared to those measured in the field but the modelled fracture network geometry may be vastly different from that seen in the field. For the construction of a more realistic model of the rock discontinuity network, additional rock mass parameters need to be included in the modelling exercise. A review of the literature suggests that one of the main determinants of the geometry of discontinuities within the rock mass is the timing and termination of discontinuities (Harries 2001). The geological history of discontinuities can be estimated from knowledge of the structural geology and, in particular, any cross cutting relationships observed in the rock mass discontinuities. A blocky pattern that appears to be defining in situ blocks can be seen in the discontinuity network shown in Figure 2.31. This is partly due to the different timings of discontinuity formation and their terminations upon one another. In Figure 2.31 many of the discontinuities terminate on other discontinuities. It is unlikely that the geometry of the rock mass could be accurately recreated without modelling the termination seen in the rock mass. A number of engineering applications rely on a representative determination and modelling of the rock mass discontinuity geometry. These include the determination of in situ block size and the hydraulic conductivity of a rock mass. This section will confine itself to the hierarchical modelling of rock mass discontinuities to better simulate the rock mass geometry. The use of this modelled discontinuity network geometry in the estimation of primary fragmentation will be discussed in Chapter 4.


94

Photograph Digitised discontinuity trace map

Figure 2.31: Discontinuity network, Kangaroo Point, Brisbane, Australia (after Harries 2001)

Discontinuity termination statistics

The International Society of Rock Mechanics (Brown 1981) has suggested that termination data be presented as a termination index, Tr, calculated as

T NN N Nr

r

r o j

=+ +100

(2.15)

where Nr is the number of fracture endpoints seen terminating in intact rock, Nj is the number of fracture endpoints seen terminating on another joint and No is the number of fracture endpoints not seen terminating (obscured). This termination index is a useful measure of discontinuity termination. A high Tr value suggests that the rock mass will contain many intact rock bridges, rather than being made up of discrete blocks. For the geometry to be accurately reproduced, a modelled rock mass must have a termination index consistent with that measured in the field. However, the Tr measure cannot account for fractures crossing other fractures (referred to as ‘X’ intersections). This factor has to be taken into account during the modelling exercise if a realistic rock mass geometry is to be recreated. Given that the intersection of fractures is a statistical event, we need to calculate a probability of fracture termination (and consequently a probability of the fracture crossing


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through the pre-existing discontinuity). The probability of fracture termination needs to be developed using the data collected on fracture termination and the number of crossing fractures seen. Discontinuity termination probability

The number of fractures terminating in intact rock, the number terminating on other fractures and statistics on fracture crossings are now routinely included in the JKMRC rock mass characterisation methodology (Harries 2001). Using this information for a given discontinuity set s, the fracture termination probability is calculated using

( )∑

∑

=

=

+= n

iii

n

ii

st

XT

T

1

1Pr (2.16)

where, Prts is the probability of termination for the given fracture set s, n is the number of

fractures in a given fracture set s, Ti is the number of observed terminations (referred to as T intersections because of the pattern they form) for each fracture in the set (0, 1 or 2) and Xi is the number of crossing X intersections seen for each fracture in the set (0, 1, 2, …, ∞). The probability of a discontinuity crossing the pre-existing discontinuity is simply Pr Prx

sts= −1 (2.17)

where, Prx

s is the probability of fracture crossing (X intersection) for the given fracture set s. Since the estimate of discontinuity termination probability is based on observations of fracture endpoints, the estimate will be biased by the size of the sampling window within which the measurements were taken. This geometrical bias is similar to the bias in the trace length distribution discussed in Section 2.5.5. However, because of the iterative nature of the hierarchical model, the initial estimate of termination probability can be modified to better fit discontinuity observations made in the field. Overview of the hierarchical model

The hierarchical rock mass model developed by Harries (2001) implements the modelling of discontinuity ordering and termination. This sequential generation and correlation of fracture sets is intended to correspond to what happens in nature. In developing the fracture termination model, four key aims were:


96

• modelling the structural history of the rock mass by simulating the ordering of discontinuities which make up the rock fracture network;

• modelling both X and T type discontinuity intersections; • the modelling method is to be generic enough to allow all structural types of rock mass to

be modelled successfully; and • to limit the change in the discontinuity parameter set statistics (number, size and

orientation) to an insignificant level, as best estimates of these parameters have been obtained from the 3-D statistical modelling.

Achieving these aims involves two main processes. Firstly, ordering the fractures to satisfy structural history considerations and secondly, modelling the fractures sequentially, checking for intersections between the fractures. Where there is an intersection, the associated probability of termination of the discontinuity should be used to assess whether to terminate the fracture or not. Model inputs

The hierarchical model requires a number of inputs for it to perform in a satisfactory manner. The computer program prompts the user for these inputs:

• the number of fracture sets in the model; • the number of different structural events to be modelled (separate geological events of

discontinuity creation); • sampling plane dimensions (height and width); • geological history ordering of the different discontinuity sets; and • probability of termination for each discontinuity set.

Initial data structure

The initial data concerning the modelled discontinuities is derived from the output of the JKMRC 3-D discontinuity model (see Section 2.6.3). An example of the visual output of the 3-D model is shown in Figure 2.32. A 20 m long and 10 m high sampling plane is depicted with the discontinuities that intersect the plane. Two different discontinuity sets are modelled in this simple example. One fracture set is steeply dipping with large discontinuities whose orientations are roughly perpendicular to the sample plane. The second fracture set modelled is oriented horizontally with a smaller mean size and a higher volumetric density than the vertical set. In reality, most rock masses have three or more fracture sets with a far greater dispersion about the mean orientation of the fractures. It often becomes difficult to develop an understanding of the rock mass by viewing such 3-D realizations when the complexity of the rock discontinuity geometry approaches reality.


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Figure 2.32: The 3-D rock mass model showing two example fracture sets

(Harries 2001) The output used in the hierarchical model is the discontinuity information contained in the 2-D sections. Figure 2.33 shows a visual example of the output plane. This plane is the same 20 by 10 m section modelled in the 3-D simulation shown in Figure 2.32. The two discontinuity sets contained in the model can be clearly identified. The flat lying horizontal line 2 m above the floor of the simulation represents the location of the scanline from which the discontinuity information was collected.

2 m

Figure 2.33: The 2-D sample plane showing the same two fracture sets as in Figure 2.32


98

There are no limitations to the number of planes that can be modelled in the 3-D model. For the sake of visual clarity only one plane is shown in Figure 2.32. Usually multiple planes are modelled with different orientations. Some plane orientations will be sub-parallel to the modelled discontinuity sets and consequently these sets will be poorly sampled. By using a large number of differently oriented planes, this geometric sampling bias can be reduced. Every discontinuity intersecting the sample plane contains information pertaining to the fracture set number, the trace length, the end point vertices, and information concerning the start and end termination codes. These termination codes are initially set to zero (representing intact rock termination) as the 3-D model does not consider termination. The fracture endpoint termination codes are used later in the termination modelling.

Ordering of discontinuities to mimic structural history

An organisational chart illustrating the hierarchical modelling process is shown in Figure 2.34. The first step involves ordering the fractures to model the structural history of the rock mass.

Model

Next Fracture

Modify Initial Fracture Sizesor Probabilityof Termination

AllFracturesModelled

YES NO

YES NO

YES

NO

Unterminated and Unordered Fractures

Structural History and Ordering of Fractures

Sequential Placement of Fractures

Does Fracture Cross Pre-Existing Fracture

Calculate Fracture Trace Length Statistics

Review Probability of Termination

Fracture Terminates

Change Endpoint Coordinates and Endpoint Codes

Conserve Tracelength Where Possible

No Termination

YES NO

Calculate Tracelength Statistics - Acceptable Fit

Calculate Termination Statistics - Acceptable Fit

Prepare for Primary Fragmentation Modelling and Calculate Engineering Indices Figure 2.34: Organisational chart of the hierarchical modelling (Harries 2001)


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The hierarchical model uses field structural geological data to designate the historical order of formation of the different discontinuity sets (see Harries 2001). Where no hierarchical structure can be identified, a simple ordering by joint length is required. The longest fractures are more likely to be the earlier fractures formed (Hudson and Cosgrove 1997) and so are modelled first. Where multiple structural events are being modelled, a more intricate sorting routine is required. A number of procedures have been devised to order the fractures firstly by structural history and then by length. Fractures are ordered by length from the longest to the shortest within the different structural history groupings. After the fractures have been sorted in this manner, the discontinuity data are ready for the termination modelling. Fracture termination model

The fracture termination model uses geometric processes to model mechanical processes that occur in nature, namely fracture intersection and termination. After the ordering of the fractures, the next step is to determine the trace length statistics. The trace length mean, standard deviation and sum are calculated. These are used later to ensure that the termination model does not significantly alter the trace length statistics. Fractures are then placed in the order designated by the structural history and length considerations. The first fracture placed has no chance of intersecting another fracture because it is the only fracture on the plane. As subsequent fractures are placed into the simulated plane they are checked to see if they intersect any previously placed discontinuities. Since the fractures are modelled as planar traces, intersections can be checked using plane equations and by checking the limits of the discontinuity traces. In Figure 2.35 the new fracture j is placed into the simulation and checked with the previous fracture i. Firstly, the coordinates of the intersection point are calculated using a vector method. When the intersection coordinates have been found, a check is performed to see if the intersection coordinates fall between the limits of both lines. Where the intersection is not within the limits of both lines (case A) no termination has occurred. Conversely, where the intersection coordinates fall between the limits of both lines (case B) then an intersection must have occurred. When an intersection occurs, a stochastic process is employed to decide whether to terminate the fracture or not. Initially the discontinuity set number of the fracture j is looked up from the data matrix. A randomly generated number is used to determine the outcome of the termination event using the probability of termination associated with the relevant discontinuity set number. When a termination occurs, a correction procedure is invoked to retain the discarded trace length. The final action is to update the endpoint coordinates and change termination codes for the affected discontinuity.


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Start

jIntersectioncoordinates

End

End

Start

i

i

jIntersectioncoordinates

Start

End

End

Start

(A) No fracture intersection (B) Fracture intersection

Figure 2.35: Check for fracture intersections

This method is used several times so that every added discontinuity in the simulation is checked with every existing discontinuity. This uses the previously ordered data list, starting with the earliest existing fracture. When all the existing fractures have been checked, the next fracture on the list is introduced into the simulation. When all fractures have been modelled, the termination part of the program is finished and a check on the discontinuity statistics is initiated to make sure that the new rock mass simulation has an acceptable statistical fit to the original modelled data. 2.7 ROCK MASS CLASSIFICATION SCHEMES

2.7.1 Introduction

Usually during the feasibility and preliminary design stages of a project, very little detailed information is available on the rock mass properties or on the stress and hydrologic regimes applying. In these cases, rock mass classification schemes may be used in an attempt to extrapolate previous experience gained in the rock mass concerned or elsewhere. These classification schemes seek to assign numerical values to those properties or features of the rock mass considered likely to influence its behaviour, and to combine these individual values into one overall rating for the rock mass. Through correlations with previous experience, rock mass classification schemes can be used to make initial estimates of support requirements and of the strength and deformation properties of the rock mass. A number of the more widely used rock mass classification schemes have been applied in engineering design analyses for caving mines. The systems that have been utilised in this way include the RMR (Bieniawski 1974, 1976), Q (Barton et al 1974) and MRMR (Laubscher 1990) rock mass classification schemes. The use of these classification systems in cave


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engineering and the rock mass characterisation data required for their application are outlined below. 2.7.2 RMR system (Bieniawski 1974, 1976)

The Geomechanics Classification or Rock Mass Rating (RMR) system introduced by Bieniawski (1973, 1974) provided, partly on a subjective basis, a measure of the quality of rock masses for the purpose of preliminary support design in tunnelling. The RMR system has since been modified a number of times by Bieniawski (1976, 1979, 1989) and extended by a number of authors ostensibly to render it more suitable for particular applications (eg Romana 1985, Laubscher 1990). To classify a rock mass using the RMR system the rock mass is separated into a number of geotechnical zones or structural units having similar rock material and discontinuity properties. The boundaries delineating these zones will most often be geological contacts, major faults, and weathering profiles. Each unit is then rated separately according to the intact material strength (uniaxial compressive strength or point load strength), the RQD (Deere 1964), the discontinuity spacing and condition, and ground water inflow. Each of these parameters is given a rating (see Table 2.6) and the summed ratings give the basic RMR of that rock mass. This basic RMR can then be modified according to discontinuity orientation with respect to the excavation. One application of the RMR system, which uses the 1976 version of the system (which will be referred to as RMR76), is in the derivation of the Hoek and Brown (1980, 1997) rock mass strength parameters, m and s. Another application of the RMR system is in the estimation of tunnel support requirements (Bieniawski 1976, Hoek and Brown 1980).

Table 2.6: The 1976 version of the RMR system (Bieniawski 1976)

Parameter Rating

Uniaxial Compressive Strength (UCS) 0 to 15

Rock Quality Designation (RQD) 3 to 20

Discontinuity Spacing 5 to 30

Discontinuity Conditions 0 to 25

Groundwater 0 to 10

Basic RMR76 8 to 100

Orientation Adjustment (for tunnels) 0 to –12

The procedure for calculating the RMR for the rock mass is to assess the spacing and condition for each of the identified discontinuity sets separately. For example, if there are three


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discontinuity sets in the rock mass, three independent spacing and discontinuity condition assessments are made. The different discontinuity sets are then assessed to determine which will have the greatest affect on rock mass instability (lowest summed spacing and condition values). Only the spacing and condition values for the controlling discontinuity set are then used to derive the rock mass rating and the other discontinuities are ignored (except when incorporated into the RQD measurement). The RMR system has been successfully utilised in many practical design situations in a wide range of applications and remains a useful way of articulation between geologists, mining engineers and geotechnical engineers. The practical utilisation of the RMR classification system will be considered by reviewing the individual components of the system. Strength of intact rock material

The RMR system assigns a value of between 0 and 15 points to represent the intact rock strength. This value is calculated from uniaxial compressive strength or point load index test results. It has been argued by Meyers et al (1993) that the behaviour of rock masses under low stress conditions, such as in surface outcrops, is influenced more by the shear strength properties of the discontinuities than by the compressive strength of the intact rock material. The only way the shear strength of discontinuities can be considered in the RMR system is as discontinuity roughness. Development work in mines using caving methods will rarely be in low stress conditions so this argument is invalid for most cases of cave design, including caveability and fragmentation studies. Rock Quality Designation(RQD)

The RQD (Deere 1964) remains the most commonly used discontinuity characterisation parameter for drill core. The RQD is defined as the percentage of core recovered in intact pieces of 100 mm or more in length in the total length of a borehole, ie

core of length Total

mm 100pieces CoreRQD ∑ >

×= 100 (2.18)

It is normally accepted that the RQD should be determined on a core of at least 50 mm in diameter which should have been drilled with double barrel diamond drilling equipment (Hoek and Brown 1980). The use of RQD as a measure of discontinuity frequency is unreliable because

• it is reliant on the ability of the geologist logging the core to discriminate between natural fractures and those caused during blasting or drilling;

• it is reliant on the shear strength of the rock material being drilled;


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• good core recovery is dependent on the drilling practice. Machine condition and the experience and care exercise by the driller are almost impossible to take into account; and

• RQD is not a good measure of the better rock mass conditions. If a rock mass has one uniformly spaced discontinuity set with a spacing of 0.l m or 5 m, the RQD will be 100 in both cases (where a 50 mm diameter core is used).

The RQD value does not discern discontinuity orientation. If the rock mass is anisotropic, the orientation of the drilling direction to the discontinuity set will be of vital importance.

Persistent

joint set

Drilling

direction

Case A - RQD 100%

Spacing ≈ 125 mm

Case B - RQD 0%

Spacing ≈ 85 mm

Figure 2.36: Example highlighting the importance of drilling direction on RQD

Figure 2.36 shows an extreme example of the influence of drilling orientation on RQD. In this example, the same rock mass may be classified as very good (Case A) or as very poor (Case B) when using the guidelines proposed by Deere (1964). Some authors (eg Laubscher 1990) have suggested methods of dealing with this problem by considering only intact cylinders of rock (see Figure 2.37).


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(Deere, 1964)

(Laubscher, 1990)

Solid cylinder of intact rock

Figure 2.37: Original and Laubscher modified RQD

However, counting only solid cylinders of intact rock is also problematic as it can also lead to orientation bias. If discontinuities are equally spaced 110 mm apart then using the traditional (centre-line) method, the RQD will always be 100%, independent of drilling direction. This is not the case if only solid cylinders of rock are counted in the RQD measure. If the drilling direction is perpendicular to the mean discontinuity direction then the RQD will again be 100%. However, if the drilling direction is at an angle of 45° to the mean discontinuity direction, the RQD will be 0% as there will only be 89 mm of solid intact core between the intersecting discontinuities (as measured during core logging). This problem becomes even worse where the core intersects discontinuities at a very flat angle because the measurement process ignores core pieces that happen to have been drilled with a small subtended angle to one discontinuity in otherwise massive rock (see Figure 2.4). Spacing of discontinuities

The rating for discontinuity spacing varies from 30 to 5 and is determined using Table 2.7.

Table 2.7: RMR discontinuity spacing parameter (Bieniawski 1976)

Spacing of discontinuities > 3 m 1 - 3 m 0.3 - 1 m 50 - 300 mm < 50 mm

Rating 30 25 20 10 5

The experience in using the RMR system for one project was that these groupings of discontinuity spacing were too insensitive in the cases studied (AMIRA, JKMRC and Rock Technology 1997). Many of the discontinuity sets were found to have spacings of between 0.3 and 1 m. The same rating was derived whether the discontinuity sets had a mean spacing of 0.35 m or 0.95 m. These present a theoretical difference of in situ block sizes of 1 m3 to 0.027 m3 in a simple orthogonal three discontinuity set case. Rock mass strength, caveability,

(Deere 1964)

(Laubscher 1990)


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fragmentation and support requirements would be expected to be different for a rock mass containing such different block sizes. Condition of discontinuities

The discontinuity condition is given a rating from 0 to 25, depending on the descriptions given in Table 2.8. These descriptions and ratings may be used to provide estimates of discontinuity shear strength.

Table 2.8: RMR discontinuity condition parameter (Bieniawski 1976)

Condition of discontinuities

Very rough surfaces

Not continuous

No separation

Hard discontinuity

wall rock

Slightly rough surfaces

Separation < 1mm Hard

discontinuity wall rock

Slightly rough

surfaces Separation

< 1mm Soft

discontinuity wall rock

Slickensided OR

Gouge <5mm thick

OR Discontinuities open 1-5mm Continuous

discontinuities

Soft gouge > 5mm thick

OR Discontinuities open > 5mm Continuous

discontinuities

Rating 25 20 12 6 0

General observations on the use of the RMR system

The greatest advantage that the RMR system provides is its ease of use in the mining environment which aids it use as a communication tool for mining engineers, geologists and geotechnical engineers. The RMR system has provided a useful tool in conjunction with the Hoek-Brown rock mass strength criterion in the estimation of rock mass strength. While the process has worked well for rock masses with a RMR greater than 25, it does not work for very poor rock masses since the minimum value which the RMR can assume is 18 (Hoek et al 1995). 2.7.3 Q system (Barton et al 1974)

The Q system of rock mass classification was developed for tunnel support in hard rock and is based on a numerical assessment of the rock mass quality using six parameters:

1. Rock quality designation RQD 2. Discontinuity set number Jn 3. Discontinuity roughness number Jr 4. Discontinuity alteration number Ja 5. Discontinuity water reduction factor Jw 6. Stress reduction factor SRF


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These parameters are grouped into three quotients which are estimates of:

• Relative block size [RQD/Jn] • Inter-block shear strength [Jr/Ja] (≈ tanφ) • Active stress [Jw/SRF]

The overall rock mass quality factor Q is equal to the product of the three quotients:

QRQD

JnJrJa

JwSRF

=⎡⎣⎢

⎤⎦⎥×⎡⎣⎢

⎤⎦⎥×⎡⎣⎢

⎤⎦⎥

(2.19)

Numerical values for the six original parameters are obtained from published tables. These tables and the resulting Q values, unlike the RMR system, have remained relatively unchanged over the years. Only one modification has been made to the SRF parameter to allow for rock bursting conditions (Grimstad and Barton 1993). When using the Q system for Hoek-Brown strength or yield estimation, or in the Mathews’ stability chart (Mathews et al 1980), the active stress component (the quotient of the joint water and stress reduction factors) is put equal to unity as these factors are treated explicitly in the strength or yield criterion and stress analysis. The resulting equation to calculate "Q prime" is

⎥⎦⎤

⎢⎣⎡×⎥⎦

⎤⎢⎣⎡=

JaJr

JnRQDQ' (2.20)

The individual components of the Q system will be reviewed to assess the practical applicability of the classification system. Block size quotient

The approximate measure of relative block size provided by the RQD/Jn suffers from potential sampling orientation error problems because of its reliance on RQD as a measure of discontinuity intensity as discussed in Section 2.7.2. This problem can be partially overcome by using an empirical relationship between the number of discontinuities per unit volume and RQD (Palmstrom 1982):

RQD ≈ 115 - 3.3 Jv (2.21) where Jv is the volumetric joint density or the number of discontinuities contained in 1 m3 of rock.


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The main consideration when using this equation is the correctness of the joints per unit volume estimate. The estimate is prone to sampling error as the three dimensional parameter is estimated from a one or two dimensional measurement of jointing intensity. A methodology to better estimate the discontinuity density without a sampling bias problem would increase the reliability of the method. It is possible that a discontinuity model simulation could be utilised to remove sampling biases and derive an accurate estimate of volumetric joint density. The discontinuity set number parameter (Jn) gives a rating of 0.5 to 20 depending on the number of discontinuity sets in the rock mass.

Table 2.9: Q discontinuity set number classification (Barton et al 1974)

Number of discontinuity sets Jn

Massive, no or few discontinuities 0.5-1.0

One discontinuity set 2 One discontinuity set plus random 3

Two discontinuity sets 4

Two discontinuity sets plus random 6 Three discontinuity sets 9

Three discontinuity sets plus random 12

Four or more discontinuity sets, heavily fractured, "sugar cube" etc.

15

Crushed rock, earthlike 20 The Jn parameter has a significant effect on the Q or Q' value. A rock mass which has many discontinuity sets forming discontinuity bounded blocks will be weaker than a rock mass with the same discontinuity intensity but not forming blocks because it has a limited number of discontinuity sets. Hence a rock mass with a greater number of discontinuity sets would be expected to cave more easily or to require more support. However, several concerns about the use of the Jn have been identified.

1. The subjective nature of choosing the number and delineating the discontinuity sets. 2. The determination of the number of discontinuity sets in a rock mass is prone to

sampling error. If a large, well planned discontinuity mapping exercise is undertaken (with three orthogonal scanlines) more discontinuity sets are likely to be identified than in a cursory examination of just one excavation face.

3. The concept of a random joint set is flawed. A joint is defined as a discontinuity

formed to release stress within rock and along which no or very little shear movement has occurred. The type and orientation of any discontinuity is governed by the relative magnitudes of the effective stresses during propagation (Hoek 1968). The principal


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stress field within a rock mass may result from external forces such as tectonic events or the weight of overburden or from internal forces such as those resulting from contraction during cooling (Rawnsley et al 1990).

4. Many rock masses have four or more discontinuity sets, yet where the discontinuity

spacing in such sets is sufficiently large, it may be incorrect to describe the rock mass as heavily fractured or "sugarcubed".

Inter-block shear strength quotient

The inter-block shear quotient [Jr/Ja] provides a good measure of the shear strength of discontinuities and can be used in estimating friction angles of the discontinuities. The very detailed treatment of discontinuity roughness and discontinuity alteration is perhaps the strongest feature of the Q-system which is not emphasised in the RMR system. The Jr and Ja parameters and the guidelines provided on their use (Loset et al 1997) are well designed and objective. 2.7.4 Modified Basic RMR or MBR system (Kendorski et al 1983)

The modified basic RMR or MBR system was developed specifically for use in establishing drift support levels in caving mines. The data from which the system was developed were collected from several block caving mines in the USA. The organisation of the MBR system is shown in Figure 2.38. It follows closely the Geomechanics Classification (RMR) system (Bieniawski 1979) and incorporates some ideas introduced by Laubscher (1981). The main differences lie in the arrangement of the initial terms and in the adjustment sequence. In the MBR system, the inputs are selected and arranged so that a rational rating is still possible using very preliminary geotechnical information obtained from drill holes. The initial ratings obtained from rock mass characterisation are shown on the left side of Figure 2.38. They include intact rock strength, RQD, discontinuity spacing, discontinuity conditions and groundwater conditions. These five parameters are the same five parameters assessed in the Geomechanics Classification (Bieniwaski 1979) and discussed in Section 2.7.2. The only difference is that the orientation parameter is not used to derive the initial MBR but is one of the mining adjustments. The MBR is an indicator of rock mass "competence", without regard to the type of opening constructed in it. The next stage is the assignment of numerical adjustments to the MBR that adapt it to the ore block development process. Input parameters relate to excavation (blasting) practice, geometry, mining depth and fracture orientation. The adjustment values are obtained from tables and charts provided by Kendorski et al (1983). The MBR is multiplied by these decimal adjustments to obtain the adjusted MBR.


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GROUNDWATERCONDITION

0-15

STRENGTH OFINTACT ROCK

0-15

DISCONTINUITYDENSITY

(RQD, SPACING)

0-40

DISCONTINUTYCONDITION

0-30

BLASTINGDAMAGE

AB0.8-1.0

INDUCEDSTRESSES

AS0.8-1.2

FRACTUREORIENTATION

AO0.7-1.0

MAJORSTRUCTURE

S0.7-1.1

DISTANCE TOCAVE LINE

DC0.8-1.2

BLOCK/PANELSIZEPS

1.0-1.3

MODIFIEDBASIC RMR

(MBR)0-100

DEVELOPMENTADJUSTMENTS

ADJUSTED MBR =MBR x AB x AS x AO

ISOLATED DRIFTOR DEVELOPMENTSUPPORT CHART

SUPPORT RECOMMENDATIONSFOR SERVICE AREAS

SUPPORT RECOMMENDATIONSFOR DRIFTS

DURING DEVELOPMENT

SUPPORT RECOMMENDATIONSFOR DRIFTS

DURING PRODUCTION

PRODUCTIONADJUSTMENTS

FINAL MBR =AMBR x DC x PS x S

PERMANENTSUPPORT

CHART

Figure 2.38: Organisation of the MBR system (Bieniawski 1984, after Kendorski et al 1983)

2.7.5 MRMR system (Laubscher 1990)

The mining rock mass rating (MRMR) system was first introduced by Laubscher in 1974 as a development of Bieniawki’s RMR system to cater for diverse mining situations. The fundamental difference was the recognition that in situ rock mass ratings (RMR) had to be adjusted according to the mining environment so that the final ratings (MRMR) could be used for mine design. Adjustments were introduced for weathering, mining-induced stresses, joint orientation and blasting effects (Laubscher 1990). For several years, the MRMR system has been the most widely used classification system for cave mine design. Laubscher and Jakubec (2001) have recently published some revisions to the MRMR system. Only the pre-2001 version of the MRMR will be considered here. The revised version will be discussed in Section 2.7.6. To derive the basic ‘Laubscher’ rock mass rating (RMRL90), the intact rock strength, discontinuity frequency and discontinuity condition must be assessed. Intact rock strength

The first parameter in the RMRL90 classification system is intact rock strength (IRS). This IRS is defined as the unconfined compressive strength of the rock between the fractures and discontinuities. A table provided by Laubscher (1990) allocates ratings of from 1 to 20 to cater for specimen strengths from 0 to greater than 185 MPa. The upper limit was selected because


110

"IRS values greater than this have little bearing on the strength of jointed rock masses" (Laubscher 1990). Discontinuity frequency parameter

In the 1990 version of Laubscher’s MRMR, two techniques are proposed for the assessment of discontinuity frequency:

• the more detailed technique is to measure the rock quality designation (RQD) and joint spacing (JS) separately, with the maximum ratings being 15 and 25 respectively;

• the other technique is to measure all the discontinuities and to record these as fracture frequency per metre (FF/m) with a maximum rating of 40, ie the 15 and 25 from the first method are added.

Both methodologies used to assess discontinuity intensity give a result of between 0 and 40. Laubscher states that when using the fracture frequency method, all discontinuities should be assessed. When using the more detailed joint spacing and RQD method, the joint spacing refers to the spacing of joints. Joints are defined as " an obvious feature that is continuous if its length is greater than the width of the excavation or if it abuts against another joint, i.e. joints define blocks of rock". Therefore, smaller discontinuous joints are ignored when calculating joint spacings. It is important, therefore, to have RQD measurements from drill core observations (that will take into account smaller or discontinuous jointing). It is also important to recognise the difference in the method of calculating RQD suggested by Laubscher as discussed in Section 2.7.2. The approach adopted by Laubscher (1990) in defining joints disregards some of the three-dimensional characteristics of rock discontinuities. It is possible that a drive in which the mapping is carried out only just cuts (samples) the edge of a large discontinuity which may play an important role in issues such as caveability or excavation stability. This is a sampling issue that is not adequately covered in the description of the classification system. If no drill core is available to obtain RQD measurements, then the RQD needs to be estimated from excavation mapping results. It is important that this mapping takes into account the smaller discontinuities that would have been included in the RQD analysis as well as the larger-scale discontinuities. Therefore, when undertaking discontinuity scanline surveys for the calculation of RMRL90 or MRMR values, orientation, trace lengths and termination of all discontinuities should be recorded. Discontinuity condition parameter

The rating for discontinuity condition starts at a base of 40 and is reduced by percentage modifiers for:


111

• large scale joint expression; • small scale joint expression; • joint wall alteration; • joint filling; and • groundwater inflow.

Although it does not provide a mechanism for estimating discontinuity friction angle, it does nevertheless provide a comprehensive method to systematically account for discontinuity conditions that will have an effect on rock mass strength. It should be noted that water is accounted for in this classification system, and not as an effective stress component as in the Hoek-Brown failure criterion (Hoek and Brown 1997). Mining adjustments

The sum of these ratings gives the Laubscher RMRL90. To derive the MRMR from the basic RMR L90 the adjustments summarised in Table 2.10 have to be applied.

Table 2.10: MRMR mining adjustments (Laubscher 1990)

Parameter Possible adjustment %

Weathering 30-100

Orientation 63-100

Stress 60-120

Blasting 80-100

Guidelines are given for the possible adjustments identified by Laubscher (1990). However, the guidelines are ill-defined for some circumstances, which can make the adjustments highly subjective and dependent on the experience of the operator (Milne et al 1998). In particular, the adjustment for stress, which may be positive or negative, has very few guidelines. A large number of factors which may effect the stress adjustment are listed but no detailed guidelines are presented on how to derive the adjustment factors. 2.7.6 Revised MRMR system (Laubscher and Jakubec 2001)

The methods introduced by Laubscher and Jakubec (2001) for establishing the In situ Rock Mass Rating or IRMR of a given rock mass and deriving the MRMR for a particular mining application, are summarised in the flow chart shown in Figure 2.39.


112

Figure 2.39: Flow chart for calculating IRMR, MRMR and DRMS (Laubscher and

Jakubec 2001)

DETAILED DESIGNS

CAVEABILITY STABILITY FRAGMENTATION SEQUENCE GEOMETRY PILLARS CAVE ANGLES SUPPORT PIT SLOPES

INPUT DATA

IRS Mpa X 80% size adj.

JOINT SPACINGRating = 0 - 35

JOINT CONDITION Rating = 0 - 40

RBS adjustment60 – 100%

Adjustment for cemented joints

70 – 100%

RBS value Rating

JOINT OVERALLRating = 0 - 75

IRMR = 0 - 100RMS = MPa

PRESENTATION COMMUNICATION

BASIC DESIGN

ADJUSTMENTSWeathering / Orientation / Induced Stress / Blasting / Water

(30 – 100%) (63 – 100%) (60 – 120%) (80 – 100%) (70 – 110%)

DRMS MPa

MRMR0 - 100

MAJOR STRUCTURES

Mpa Mpa 0 - 25


113

The essential features of each of the steps and the main changes made from Laubscher’s previous RMRL90 and MRMR system are summarised in the following points. 1. The intact rock strength (IRS) is the unconfined compressive strength derived from the

testing of rock cores. A nomogram is given for determining the "corrected" value of the IRS for the case in which the rock contains intercalations of weaker material.

2. To obtain the rock block strength (RBS) from the "corrected" IRS, two adjustments are

made. The first is a multiplication by 0.8 to allow for the size effect when applying the results of small scale laboratory tests to the field scale. The second is an adjustment for the presence of fractures and veins which will reduce the strength of the rock block. Laubscher and Jakubec (2001) give a procedure in which an adjustment is determined from the product of the inverse of the Moh’s hardness of the fractures and veins and the fracture and vein frequency per metre. Thus RBS = IRS x 0.8 x fracture/vein adjustment. A rating of 0 – 25 for the rock block strength is then read from a graph.

3. In the revised system, the joint spacing rating for open joints is reduced to a maximum of

35 and is determined on the basis of the joint spacings of one, two or three joint sets and no more. The ratings differ from those used previously. If there exists in the rock mass a set of cemented joints in which the cement strength is less than that of the rock material, a further downwards adjustment to the joint spacing rating of 70 to 100% is made depending on the number of cemented joint sets present (one or two) and the cemented joint spacing.

4. The maximum joint condition (JC) rating for single joints remains at 40 but the joint

condition adjustments have been revised. A chart is given for determining the JC rating for multiple joint sets having differing joint conditions.

5. The overall joint rating of 0–75 is the sum of the joint spacing (0–35) and the joint

condition (0–40) ratings. 6. The IRMR is calculated as the sum of the rock block strength and overall joint ratings. 7. The IRMR may be multiplied by adjustment factors for weathering (30-100%), orientation

(63-100%), induced stress (60-120%), blasting (80-100%) and water (70-100%) to give the Mining Rock Mass Rating or MRMR. Laubscher and Jakubec (2001) give guidelines and tables for use in determining these adjustment factors.

8. The following quotation from Laubscher and Jakubec (2001) is considered to be

especially important in the present context: "The adjustment procedure has been described in previous papers, where it was stated that the adjustment should not exceed two classes, but, what was not made clear is that one adjustment can supersede another and that the total adjustment is not likely to be a multiplication of all the adjustments.


114

For example, a bad blasting adjustment would apply in a low stress area but in a high stress area the damage from the stresses would exceed that of the blasting and the only adjustment would be the mining induced stress. The MRMR for a caveability assessment would not have blasting as an adjustment, nor would it have weathering as an adjustment unless the weathering effects were so rapid so as to exceed the rate of cave propagation as a result of the structural and stress effects. The joint orientation and mining induced stress adjustments tend to complement each other. The object of the adjustments is for the geologist, rock mechanics engineer and planning engineer to adjust the IRMR so that the MRMR is a realistic number reflecting the rock mass strength for that mining situation."

9. The design rock mass strength (DRMS) is the RMS reduced by the same factor as that

applied to the IRMR to produce the MRMR. Thus DRMS = RMS x MRMR / IRMR. 10. Laubscher and Jakubec (2001) do not indicate whether or not the previously published

correlations of various engineering behaviours with MRMR, most notably that given by Laubscher’s caving chart (see Figure 3.1 and Section 3.2) will change as a result of the changes made to the calculation of MRMR.

2.7.7 Geological Strength Index (GSI)

As part of the continuing development and practical application of the Hoek-Brown empirical rock mass strength criterion to be discussed in Section 2.8, Hoek (1994) and Hoek et al (1995) introduced a new rock mass classification scheme known as the Geological Strength Index (GSI). The GSI was developed to overcome some of the deficiencies that had been identified in more than a decade of experience in using Bieniawski’s Rock Mass Rating (RMR) with the rock mass strength criterion. This brief account of the GSI is based on that given by Marinos and Hoek (2000) and should be read in conjunction with Section 2.8. The GSI is an index developed specifically as a method of accounting for those properties of a discontinuous or jointed rock mass which influence its strength and deformability. The strength of a jointed rock mass depends on the properties of the intact pieces of rock and upon the freedom of those pieces to slide and rotate under a range of imposed stress conditions. This freedom is controlled by the geometrical shapes of the intact rock pieces as well as by the condition of the surfaces separating them. The GSI seeks to account for two features of the rock mass – its structure as represented by its blockiness and degree of interlocking, and the condition of the discontinuity surfaces. Using Figure 2.40, the GSI may be estimated from visual examination of exposures of the rock mass or borehole core. It will be noted that the GSI does not include an evaluation of the uniaxial compressive strength of the intact rock pieces and avoids the double counting of discontinuity spacing as in the RMR system.


115

Figure 2.40: Geological Strength Index (GSI) for jointed rock masses (Hoek 2003)

Although the origin and petrography of the rock are not represented in Figure 2.40, the rock type will usually constrain the range of GSI values that might exist for rock masses of that rock type. Marinos and Hoek (2000) present a series of indicative charts which show the most probable ranges of GSI values for rock masses of several of generic rock types. Figure 2.41 shows one of the charts for a group of rock types encountered in some block caving operations, the ultrabasic rocks or ophiolites (mainly peridotites and diabases). Marinos and Hoek (2000) note that, even when relatively unweathered, a characteristic of these rocks is that their discontinuities may be coated by weak minerals produced by alteration or dynamic metamorphosis. This translates their locations towards the right of the GSI chart compared to fresh igneous rocks, for example. The ophiolites may be transformed into serpentinites which, as indicated in Figure 2.41, can have low GSI values and produce very weak rock masses.


116

Figure 2.41: The most common GSI ranges for typical ophiolites

(Hoek 2003)

2.7.8 Conclusions

Rock mass classification systems are well utilised and will likely remain an integral part of rock mass characterisation and engineering design applications in cave mining. Most of the issues raised in reviewing the rock mass classification systems involve the treatment of rock mass discontinuities in the classification systems. A number of sampling issues associated with rock mass discontinuities have been raised and the use of RQD as a measure of discontinuity intensity has been questioned. This highlights the need to develop and use more comprehensive methods of characterising the rock mass discontinuities. With a more statistically rigorous treatment of discontinuity data it may be possible to remove some of the ambiguity surrounding some discontinuity measures. It may even be possible to develop confidence limits on rock mass classification ratings (Harries 2001).


117

2.8 THE MECHANICAL PROPERTIES OF ROCK MASSES

2.8.1 Scope

In developing rational solutions to a range of geomechanics problems encountered in the design of block and panel caving mines, it is necessary to define the mechanical properties of the rock mass, usually represented by its stress-strain behaviour. Important aspects of this behaviour are the constants relating stresses and strains in the elastic range, the stress levels at which yield, fracturing or slip occurs within the rock mass, and the post-peak stress-strain behaviour of the fractured or "failed" rock (Brady and Brown 1993). The collection of data for use in estimating some of these properties is part of the rock mass characterisation process. In some problems, it is the behaviour of the intact rock material that is of concern. This will be the case when considering the excavation of rock by drilling and blasting, or when considering the stability of excavations in good quality, brittle rock subject to rock burst conditions (see Section 10.3). In other cases, the behaviour of single discontinuities, or of small numbers of discontinuities, may be of paramount importance. Examples of this class of problem arise in the design of extraction level excavations to be considered in Chapter 6. They include the equilibrium of blocks of rock formed by the intersections of three or more discontinuities with the roof or wall of an excavation, the intersection of one or more discontinuities of critical orientation with a brow (see Section 6.4.3), and cases in which slip on a major fault must be considered. A different class of problem is that in which the rock mass must be considered as an assembly of discrete blocks as in the example shown in Figure 1. 7. In this case, the normal and shear force-displacement relations at face-to-face and corner-to-face block contacts are of importance in the analysis. Finally, it is sometimes necessary to consider the global response of a jointed rock mass in which the discontinuity spacing is small on the scale of the problem domain. Caving of jointed rock masses and subsidence to surface are obvious examples of this class of problem. Figure 2.42 illustrates the transition from intact rock to a heavily jointed rock mass with increasing sample size in a hypothetical rock mass surrounding an underground excavation such as an extraction level drift. It is beyond the scope of this book to consider in detail the full range of problems outlined above. A useful introduction to a range of aspects of rock and rock mass strength and deformability is given by Brady and Brown (1993). Here, emphasis will be placed on the overall strength, and to a lesser extent, the deformability, of jointed rock masses. In the process, some reference will be made to the strength of intact rock.

2.8.2 The Hoek-Brown Empirical Strength Criterion

The reliable determination of the global mechanical properties of large masses of in situ discontinuous rock has long been one of the most challenging problems met in the field of rock mechanics. In an attempt to provide a "first pass" method of estimating the strength of jointed


118

rock masses for use in underground excavation design, Hoek and Brown (1980) developed an empirical rock mass strength criterion based on their earlier research into the brittle fracture of rock (Hoek 1968) and the mechanical behaviour of discontinuous rock masses (Brown 1970). The criterion took the strength of the intact rock as its starting point and introduced factors to reduce the strength on the basis of the spacing and characteristics of the joints within the rock mass. Hoek and Brown (1980) used the 1976 version of Bieniawski’s Rock Mass Rating (see Section 2.7.2) as an index of the geological characteristics considered likely to influence the mechanical properties of the rock mass. Because of a lack of suitable alternatives, the Hoek-Brown empirical rock mass strength criterion was soon adopted by rock mechanics practitioners and sometimes used for purposes for which it was not originally intended and which lay outside the limits of the data and methods used in its derivation. Because of this, and as experience was acquired with its practical application, a series of changes were made and new elements introduced into the criterion. Hoek and Brown (1997) consolidated the changes made to that time into a comprehensive account of the criterion and gave a number of worked examples to illustrate its application in practice. A further update is given by Hoek et al (2002). The summary of the criterion presented here is based on those of Hoek and Brown (1997), Hoek et al (2002) and Marinos and Hoek (2000). Based on analyses of a wide range of triaxial test data on rock samples, Hoek and Brown (1980) proposed that the peak strength of the intact pieces of rock in a rock mass of a given rock type could be represented by the equation

5.0

ci

3ici31 1.0 m

⎟⎟

⎠

⎞

⎜⎜

⎝

⎛+

σ

′σσ+′σ=′σ (2.22)

where ′σ1 and ′σ3 are the major and minor principle effective stresses at peak strength, respectively; mi is a parameter obtained by the statistical analysis of a set of triaxial compression tests on carefully prepared 50 mm diameter core samples of the intact rock; and σci

is the measured uniaxial compressive strength of the intact rock. As the result of an analysis of triaxial test results on a wide range of rock types, it has been found that preliminary or approximate values of the parameter mi can be obtained from Table 2.11. The values in parentheses in Table 2.11 are estimates only. The range of values quoted for each rock group depends on the granularity and interlocking of the crystal structure; the higher values are associated with tightly interlocked and more frictional characteristics. It should be noted, however, that the values given in Table 2.11 are indicative only and that the value of mi for a given rock is established most reliably by triaxial testing. It is important that a suitable range of values of σ3 be used in carrying out the tests used to determine mi values.


119

Hoek and Brown (1980) used the range 0 < σ3 < σci in deriving the original values of σci and mi and it is recommended that this range continue to be used in practice (Hoek and Brown 1997). The generalised Hoek-Brown strength criterion for jointed rock masses is given by

31 3 ci b

ci

m sa

σσ σ σσ

′⎛ ⎞′ ′= + ⎜ + ⎟⎜ ⎟⎝ ⎠

(2.23)

where mb is the reduced value of the material constant m for the jointed rock mass, and s and a are parameters which depend on the characteristics or quality of the rock mass. The values of mb and s are related to the GSI for the rock mass by the relationships

⎟⎠⎞

⎜⎝⎛=

28100 - GSIexp m m ib (2.24)

and

⎟⎠⎞

⎜⎝⎛=

9100- GSI exp s (2.25)

In the initial 1980 version of the criterion, the parameter, a, took a constant value of 0.5. Subsequently, for GSI < 25, this value was increased to

200GSI - 0.65 a = (2.26)

which means that for rock masses of very poor quality, a ~ 0.65. Hoek et al (2002) have introduced a new expression for a which applies over the full range of GSI values:

( )20/3-GSI/15- e- e 61 0.5 a += (2.27)

Note that for GSI > 50, a ~ 0.5.


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Table 2.11: Values of the constant mi for intact rock by rock group (Hoek 2003)

Texture Rock type

Class

Group Coarse Medium Fine Very fine

Clastic

Conglomerates *

Breccias *

Sandstones 17±4

Siltstones 4±2

Greywackes (18 ±3 )

Claystones 7±2

Shales (6 ±2) Marls (7 ±2)

Carbonates Crystalline Limestone

(12 ± 3)

Sparitic Limestones

(10±2)

Micritic Limestones

(9±2)

Dolomites (9 ± 3)

Evaporites

Gypsum 8±2

Anhydrite 12±2

Non-Clastic

Organic Chalk 7±2

Non Foliated

Marble 9±3

Hornfels (19±4)

Metasandstone (19 ± 3)

Ouartzites 20 ±3

Slightly foliated

Migmatite (29 ±3)

Amphibolites 26 ±6

Gneiss 28 ±5

Foliated** Schists 12 ±3

Phyllites (7 ± 3)

Slates 7 ± 4

Light

Granite Diorite 32 ±3 25 ±5 Granodiorite (29 ± 3)

Plutonic

Dark

Gabbro Dolerite 27 ±3 (16 ±5) Norite 20 ±5

Hypabyssal Porphyries (20 ±5)

Diabase (I5±5)

Peridotite (25 ±5)

Lava

Rhyolite (25 ± 5) Andesite

25 ±5

Dacite (25 ±3) Basalt (25 ±5)

Volcanic Pyroclastic Agglomerate (19

±3) Breccia (19 ±5)

Tuff (13 ±5)

* Conglomerates and breccias may present a wide range of mi values, depending on the nature of the cementing material and the degree of cementation, so they may range from values similar to sandstone to values used for fine grained sediments (even under 10).

** These values are for intact rock specimens tested normal to bedding or foliation. The value of mi will be significantly different if failure occurs along a weakness plane.

It will be seen from the above, that in order to use the Hoek-Brown criterion to estimate the strength of a given jointed rock mass, three "properties" of the rock mass are required: • the uniaxial compressive strength of the intact rock, σci;

SE

DIM

EN

TA

RY

M

ET

AM

OR

PH

IC

IGN

EO

US


121

• the value of the constant mi for the intact rock; and • the value of the GSI for the rock mass as discussed in Section 2.7.7. The uniaxial compressive strength of the rock mass, σcm, is obtained by setting ′σ3 to zero in

Equation 2.23, giving σcm = σci.sa (2.28)

The tensile strength of the rock mass which represents the interlocking of the particles when they are not free to dilate, is given by setting σ1

/ to zero in Equation 2.23. This produces an equation which does not readily yield a simple expression for σ3

/ = σtm . However, a value of σtm may be obtained by putting σ1

/ = σ3/ = σtm in Equation 2.23 on the basis that the uniaxial and

biaxial tensile strengths of brittle rocks are approximately equal (Brown 1974, Hoek 1968). The resulting expression is

ci

b

-s mtmσσ = (2.29)

Analytical solutions and numerical analyses of a range of mining geomechanics problems use Mohr-Coulomb shear strength parameters rather than principal stress strength criteria of the form of Equations 2.22 and 2.23. Because Equations 2.22 and 2.23 are non-linear, the corresponding shear strength envelopes are also non-linear. This means that equivalent Mohr-Coulomb shear strength parameters have to be determined for a given normal stress or effective normal stress. Methods of doing this are given by Hoek and Brown (1980, 1997) and by Hoek et al (2002). An example of the use of this approach in a limiting equilibrium analysis will be given in Section 9.4.1. It is important to recognise that the Hoek-Brown strength criterion applies only to isotropic rock masses. It should not be used when failure occurs along a particular discontinuity or small number of discontinuities. The circumstances under which the criterion is, and is not, applicable are illustrated in Figure 2.42 and discussed by Hoek and Brown (1997). The major circumstances in which the criterion is likely to be of use in the analysis and design of block caving mines, is in the design of the pillars around the extraction or production level excavations, in crown pillar design and in the analysis of a number of forms of caving to surface and surface subsidence.


122

Figure 2.42: Applicability of the Hoek-Brown empirical rock mass strength criterion at

different scales

2.8.3 Rock Mass Deformation Modulus

A value of the deformation (or Young’s) modulus of the rock mass is required when carrying out numerical stress analyses to determine the stresses and displacements induced in the rock mass when stresses are redistributed as a result of excavation. This becomes especially important on and around the extraction levels of block and panel caving mines where the degree of excavation is high. Clearly, like its strength, the in situ deformability if a rock mass will depend on the properties of both the intact rock and the discontinuities present within the rock mass. Experience shows that this deformability can be highly variable and difficult to measure or predict. Based on the back analysis of dam foundation deformations, Serafim and Pereira (1983) proposed a relationship between in situ modulus and Bieniawski’s RMR:

RMR-10

40 10 E = (2.30) where E is expressed in GPa.


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This relationship was used with the early versions of the Hoek-Brown criterion. However, it was found that for many poor quality rock masses, this method gave modulus values that were too high and so a modified relationship was proposed for values of σci of less than 100 MPa (Hoek and Brown 1997):

GSI-10

40 10 100

ciE σ= ⋅ (2.31)

The effect of a reduction in the uniaxial compressive strength of the intact rock below 100 MPa in poorer quality rock masses is illustrated in Figure 2.43. Although Equation 2.31 has been found to work reasonably well for those cases to which it has been applied, it should be recognised that it is only an approximate method and should always be verified by field experience. It must also be remembered that although rock mass deformability is often anisotropic (Brady and Brown 1993), Equations 2.30 and 2.31 assume isotropic behaviour.

Figure 2.43: Rock mass deformation modulus as a function of GSI and uniaxial

compressive strength of the intact rock (Hoek 2003)


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2.9 IN SITU STRESSES

It is well established that the behaviour of underground excavations in rock is influenced by the pre-excavation state of stress (eg Hoek and Brown 1980). The stresses and displacements induced in the rock surrounding an excavation will depend on the initial state of stress (which may, itself, have been influenced by other nearby openings), the geometry of the excavation and the constitutive (stress-strain) behaviour of the rock mass. These induced stresses and displacements will influence the stability of the excavation, the need for reinforcement of the rock mass or filling the excavation, and the initiation and propagation of caving. The role played by in situ and induced stresses in caving mechanics has been discussed in Section 1.2.2. The influence of undercut strategy and design on the stresses induced on the extraction and undercut levels will be considered in Chapter 5. Except in particular geological environments, it is usually not possible to predict in situ states of stress using the principles of mechanics. This is because both the magnitudes and orientations of stresses are influenced by a wide range of factors including tectonic history, topography, erosion, differences in the elastic constants of the lithological units and the presence of faults and other discontinuities. It must also be remembered that stress is a tensor quantity which requires the quantification of six unknowns in order to define it fully at a point. It may not be assumed that the principal in situ stresses will be oriented horizontally and vertically although, in some circumstances, it may be both reasonable and convenient to do so. Among the first significant measurements of stresses in underground mines were those made in the iron ore mines of eastern France in the early 1950s using the then recently developed flat-jack method (Tincelin 1952). However, it was the pioneering work of Hast (1958) in Sweden which demonstrated that the horizontal stresses in rock could be several times the vertical or overburden stress. This result has been confirmed by a wide range of subsequent measurements and studies as reflected in compilations of measurements made in Australia and elsewhere (eg Brown and Windsor 1990, Hoek and Brown 1980, World Stress Map Project 2003, Zoback 1992). Since the 1950s, a number of approaches have been developed for the measurement of in situ stresses. As well as the original flat jack method which suffers from a number of inherent disadvantages, most emphasis has been placed on a variety of overcoring methods and on the hydraulic fracturing method which, like the flat-jack method, makes a number of sometimes limiting assumptions (Brown and Windsor 1990). For the last two decades, the state-of-the-art method has been the CSIRO hollow inclusion stress cell developed by Worotnicki and Walton (1976). Some ingenious larger scale methods of measuring local stress fields have also been used (eg Brady et al 1976). Other approaches used to estimate in situ stresses include the resolution of earthquake focal mechanisms, the interpretation of stress conditions associated with young geological features including faults, and the back analysis of wellbore breakouts


125

and excavation behaviour. Overviews of these various methods of measuring and estimating in situ stresses are given by Brown and Windsor (1990) and Amadei and Stephansson (1997). Despite more than 40 years of accumulated experience in measuring and estimating stresses at mine sites, the undertaking can still be fraught with difficulty. It has been found that stress magnitudes and orientations can vary markedly with the presence of discontinuities and with changes in rock properties. Even though these effects are often observed, it is still considered essential that every effort be made to measure (from exploration openings) or otherwise estimate the mine- or orebody-scale stresses before decisions are made about the adoption of a particular mass mining method or layout. In the absence of any other source of information, recourse can be made to the several excellent databases and stress maps available on mining district, regional and world scales (eg Amadei and Stephansson 1997, Hillis et al 1999, World Stress Map Project 2003). These stress maps usually show the directions of the principal horizontal stresses in the region of interest. When relying on the data presented in these databases and stress maps, it should be remembered that not all of the results may be assumed to be accurate, particularly at shallow depths where the strains and stress components being measured are low.

126

CHAPTER 3

CAVEABILITY ASSESSMENT 3.1 INTRODUCTION

As discussed in Chapter 1, block and panel caving methods of mining require more development before the start of production than most other methods. This means that they have comparatively high initial capital costs and are relatively inflexible. Initiating and sustaining the cave govern the early productivity and economics of the operation. The ability to predict with a reasonable degree of accuracy the undercut dimensions at which caving will initiate and propagate is fundamental to the success of the mining method in most orebodies. This issue is becoming increasingly important given the considerable interest now being shown in applying caving methods to stronger rock masses. If the caveability of the orebody is not assessed with reasonable accuracy, expensive and time consuming measures may be required subsequently to initiate or sustain caving (eg Kendrick 1979, van As and Jeffrey 2000). The mechanics of, and the factors influencing, caving were discussed in Section 1.2. The major factors likely to influence caveability are discontinuity geometry and strength, rock mass strength, orebody geometry, undercut dimensions, the stresses induced in the crown of the undercut or cave and the presence of any boundary weakening. Although the importance of these factors had been recognised for some time and several attempts had been made to codify or quantify their influences (eg Coates 1981, McMahon and Kendrick 1969, Mahtab et al 1973), it was not until the development of Laubscher’s caving chart approach in the 1980s that a method of achieving this became widely available. Although not used in some caving mines, Laubscher’s caving chart (Diering and Laubscher 1987, Laubscher 1990, 1994, 2001) is the general industry standard method of assessing caveability. Numerical modelling holds the possibility of providing a more fundamental and rigorous assessment of cave initiation and propagation than empirical methods. This approach may have advantages in cases for which current experience is lacking or not well developed. The application of a numerical modelling approach to establishing caveability is discussed later in this Chapter. The use of an extended version of another empirical approach, the Mathews stability graph method, is also discussed.

Chapter 3: Caveability Assessment

127

3.2 LAUBSCHER’S CAVING CHART

3.2.1 Overview

The challenge faced in attempting to develop empirical methods of predicting caveability is to find a means of combining measures of rock mass quality, undercut geometry and induced stresses into one simple and robust tool. As shown in Figure 3.1, Laubscher has done this by plotting the value of his Mining Rock Mass Rating, MRMR introduced in Section 2.7, against the hydraulic radius, S (area/perimeter), of the undercut which is a measure of the undercut size and shape. If the orebody is elongated in one direction and has a limited width, the question can arise as to whether the minimum dimension of the undercut can influence caveability, irrespective of the value of the hydraulic radius. Data collected as part of the study to be discussed in Section 3.3 suggest that the hydraulic radius (or shape factor) is a satisfactory predictor of stability or caveability for aspect ratios of less than about three.

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50 60 70 80

Hydraulic Radius, (m)

Mod

ified

Roc

k M

ass

Rat

ing,

MR

MR

Stable

Transition

Caving

Not Specified

TransitionalZone (defined by Laubscher, 1994)

Caving Zone

Stable Zone

L4 Stopes

BI 48

L4 Stopes

FredaDurnacoal

King

Andina 2nd PanelTen Sub-6

Renco

Shangani

Cassair

King

Carlsbad Cave

Rosh Pinah

La Verna Cavern

Urad

BI 31

BI 6

B4B3

BA5 (gabbro - final caving into pit)

BA5 (gabbro sill)

BA5 (gabbro)

BA5 (TKB)

BA5 (HB)

B2B1 BI 16 Shabanie

BI 16 Shabanie

Transitional Zone

Figure 3.1: Laubscher’s caving chart (from Bartlett 1998)

Cassiar


128

Over a period of time, starting with the chrysotile asbestos mines of Zimbabwe, Laubscher collected data from a range of caving mines and plotted on the chart a series of points representing particular mines, blocks or panels. The points plotted could represent stable (non-caving), transitional (major collapses or partially caving) or caving conditions of the blocks or panels concerned. On the basis of this information, Laubscher drew in boundary lines dividing the chart into stable, transitional and caving zones. As more data were gathered and analysed, and the method of calculating values of the MRMR evolved, the positions of the boundary lines became subject to change. When a possible new caving operation was being studied, the MRMR would be estimated and the value of the hydraulic radius required to initiate caving read from the chart. Laubscher’s chart has found widespread use in caving operations internationally. It has been especially successful when applied to the weaker orebodies for which it was developed initially. 3.2.2 The Mining Rock Mass Rating (MRMR)

Laubscher’s Mining Rock Mass Rating (MRMR) was discussed in Sections 2.7.5 and 2.7.6. The MRMR is based on a geomechanical classification system for rock masses termed the Rock Mass Rating (RMR) which is a modification of the well-known classification system developed by Bieniawski (1974, 1976) and discussed in Section 2.7.2. It will be recalled that Bieniawski’s original RMR involves the summation of numerical measures of five factors influencing the mechanical response of a rock mass:

• strength of the intact rock material;

• Rock Quality Designation (RQD);

• spacing of the joints;

• condition of the joints; and

• groundwater conditions.

Laubscher’s basic RMR (Laubscher 1977, 1984) uses the first four of these factors but measures some of them in different ways from Bieniawski and allows for the effects of water in the joint condition term. Following a series of modifications, Laubscher’s version of RMR (eg Laubscher 1994) has now diverged considerably from Bieniawski’s original and can produce significantly different ratings for given rock masses.

Laubscher (1990) developed a range of mining adjustments to his RMR to account for the effects of weathering, orientation of jointing, induced stresses and blasting as shown in Table 2.10. The adjusted value of RMR becomes the MRMR. Of these factors, only the joint orientations and the induced stresses are considered likely to influence caveability. Although limiting ranges of adjustment were specified for each factor, few guidelines were given for


129

making estimates of the appropriate adjustment factors to apply within the given ranges. The decision is left to individual engineering judgment. The effect of the adjustments can be large; the MRMR can range from 50% (the minimum recommended) to 120% of the RMR. Recently, Laubscher and Jakubec (2001) revised the methods used for estimating the MRMR and the associated adjustments (see Section 2.7.6). This revised system has not been tested in a wide range of practical applications as yet and it is not clear whether or not it will require the amendment of the boundaries of the zones of stability shown in Figure 3.1.

3.2.3 Delineation of Stability Zones

As illustrated in Figure 3.1, Laubscher’s caving chart is divided into three zones - stable, transitional and caving. A total of 29 case history data points are shown on Figure 3.1 - three stable, four transitional, 17 caving and five unspecified but assumed to be stable. Stewart and Forsyth (1995) have noted that, as with all empirical methods, a stability graph approach relies heavily on the database from which it is derived. Its predictive capability is likely to improve with the availability of more reliable data from a wider range of conditions. For example, the original Mathews’ stability graph for open stope design (Mathews et al 1980) was based on only 50 case studies. The collection of more data by Potvin et al (1989), Stewart and Forsyth (1995) and Trueman et al (2000) has increased the database to about 500 cases resulting in changes to the stability zones. It is possible that the addition of more data to Laubscher’s caving chart may lead to similar changes in the current positions of the zone boundaries. Lorig et al (1995) reported some observations of Karzulovic who proposed an additional “marginal caving region” for Laubscher’s caving chart based on experience at CODELCO-Chile’s El Teniente and Andina mines. There is now some suggestion that the current caving zoning may under-estimate the hydraulic radius required to ensure continuous caving in some circumstances, especially in rock masses with values of MRMR of greater that 50, although the predictions have been found to be reliable in some others. For example, Lift 1 of the Northparkes E26 block cave was predicted to cave continuously at a hydraulic radius of 25 (and an area of approximately 1 hectare) but the cave was not self-propagating (van As and Jeffrey 2000). Of course, a number of geotechnical and operational factors could have combined to produce this outcome. The suggestion also arises from the Northparkes case that, in its pre-2000 form, Laubscher’s method may not be as applicable to isolated, fully constrained, stronger orebodies having small footprints as it is to larger, weaker orebodies in which adjacent mined blocks may provide some stress release. In the most recent version of the stability graph known to the writer, Laubscher (2001) has recognised the possible influence of the plan shape of the block or the orebody on caveability. When the shape is approximately rectangular with an axial ratio of more than 1.5, the boundary between the caving and transitional zones is in a similar position to that shown in Figure 3.1.


130

When the shape is more closely circular with an axial ratio up to about 1.3, a second curve is used with values of hydraulic radius at a given value of MRMR being 1.2 times those in the first case. 3.2.4 Summary

Since the 1980s, Laubscher’s caving chart has been the major method used internationally to predict caveability in block and panel caving mines. It has been particularly successful when applied to the weaker and larger orebodies for which it was first developed. However, the perception has grown in recent years that it may not always provide satisfactory results for stronger, smaller and isolated or constrained blocks or orebodies. There may be insufficient case studies available, especially for rock masses having MRMR values of more than 50, to enable the three zones of stability to be delineated with a reasonable degree of accuracy over a wide range of conditions. This is not an unusual happening when an attempt is made to extend an empirical method outside the limits of the experience for which it was first developed. A practical difficulty in the application of the method is that there may be insufficient guidelines available for the inexperienced user seeking to establish values of the adjustment factors to be applied to the RMR. 3.3 MATHEWS’ STABILITY GRAPH APPROACH

3.3.1 Overview

The Mathews’ stability graph (Mathews et al 1980) is very similar in concept to Laubscher’s caving chart and predates it in the open literature by several years. Mathews et al (1980) first developed the method for open stope design for mining in hard rock at depths below 1,000 m. As was noted in Section 3.2.3, the initial stability graph was based on a relatively small amount of data. A number of authors have since collected significant amounts of new data from a variety of mining depths and for a wider range of rock mass conditions to test the wider applicability of the method and have proposed modifications and extensions (Potvin et al 1989, Stewart and Forsyth 1995, Trueman et al 2000). The modifications have related largely to the delineation of the stability zones. Potvin et al (1989) made some changes to the way in which some adjustment factors were calculated but these changes have been shown to make no appreciable difference to the predictive capability of the technique (Stewart and Forsyth 1995, Trueman et al 2000). Therefore, only the original Mathews method of determining adjustment factors will be described here. The design procedure is based upon the calculation of two factors - the stability number, N, which represents the ability of the rock mass to stand up under a given stress condition, and the shape factor or hydraulic radius, S, which accounts for the geometry of the surface. The stability number is analagous to Laubscher’s MRMR, while the shape factor is identical to the


131

hydraulic radius used in Laubscher’s caving chart. For given stope surfaces, these factors are plotted on the stability graph which is divided into zones of predicted stability and instability. Stability number

The stability number, N, is defined as N= Q' .A .B .C Q' is calculated from the results of structural mapping or geotechnical core logging of the rock mass using the NGI Q classification system (Barton et al 1974) discussed in Section 2.7.3 but assuming the joint water reduction parameter and stress reduction factor to be both equal to one. The rock stress factor, A, is determined from the ratio of the unconfined compressive strength of the intact rock to the compressive stress induced at the centre-line of the stope face. The induced stress is found using an elastic stress analysis package or estimated from published stress distributions (eg Hoek and Brown 1980). Figure 3.2 provides a graph which shows the relationship between the strength to stress ratio and the rock stress factor, A. The joint orientation adjustment factor, B, is a measure of the relative difference between the dips of the stope surface and the critical joint set (see Figure 3.2). The gravity adjustment factor, C, reflects the fact that the orientation of the stope surface influences its stability and is determined from Figure 3.2. To the inexperienced user, the determination of the stability number may give an inaccurate impression of the engineering rigour of this design technique. However, it must be recognised that it is difficult to separate and evaluate empirically each of the several factors that influence the stability of stope surfaces. The adjustment factors appear to have been determined largely without reference to a database of recorded observations. Nevertheless, the detailed guidelines provided for the determination of adjustment factors represent some advance over other methods. Stability zones

As shown in Figure 3.3, the original Mathews stability graph contained three zones separated by transitions - a stable zone, a potentially unstable zone and a potentially caving zone. In the modified stability graph developed by Potvin et al (1989), these three zones were reduced to a stable zone and a caved zone separated by a transition. The choice of the word “caved” to represent what is essentially an unstable zone was commented on by Stewart and Forsyth (1995) who noted that the term has a particular meaning in mining which appears not to have been adhered to in the modification proposed by Potvin et al (1989).


132

Factor A

Rock stress factor

Factor C Design surface orientation factor

Factor B Joint orientation adjustment

Figure 3.2: Evaluation of adjustment factors in the Mathews stability graph method

(after Mathews et al 1980)

Stewart and Forsyth (1995) updated Mathews’ stability graph proposing four zones separated by three transitions - potentially stable, potentially unstable, potential major failure and potentially caving. The potentially caving zone was determined using Laubscher’s (1990) caving chart as a guide and requires validation. Nevertheless, it was intended to represent true caving. In this sense, caving is defined as occurring when the rock mass fails and collapses until all the available void space is filled with broken rock and then continues to fail when broken rock is removed from contact with the stope surface. A major failure is defined as occurring when either greater than 30% dilution occurs or when the depth of collapse into the stope walls is greater than 50% of the smaller dimension of the opening (Stewart and Forsyth 1995). Wide transition zones between the unstable to major failure and the major failure to caving zones were assumed. Mawdesley et al (2001) give a more detailed description of the development of the Mathews stability graph method.


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Figure 3.3: The original Mathews stability graph (after Mathews et al 1980)

3.3.2 Extension of the Method

Trueman et al (2000) collected significantly more stable, minor failure and major failure case history data enabling the Mathews method to be extended to cover a much wider range of open stope sizes and rock mass characteristics. These data were combined with existing cases to produce a database of about 500 entries. The stability numbers for all these case studies were determined using the guidelines originally proposed by Mathews et al (1980). As shown in Figure 3.4, the S-N data points were plotted on log-log axes rather than the log-linear axes normally used for the Mathews method. With the increased database, a logistical regression analysis was used to define stable, minor failure and major failure boundaries. Mawdesley et al (2001) provide a detailed account of the statistical technique used.


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Figure 3.4: Extended Mathews stability graph for open stopes based on logistic

regression (Mawdesley et al 2001)

The statistical distinction between failure and major failure was not as definitive as the boundaries between stable and failure and major failure and caving. This is hardly surprising given the rather arbitrary definition of major failure. The possibility of misclassifying the two types of failure is much greater than the possibility of misclassifying a stable or a caving data point. Accordingly, Trueman et al (2000) found that it was not possible to separate failure from major failure statistically for the extended database. Although zones of stability can be defined statistically, a number of data points apparently report to the wrong stability zones. This is to be expected given the inherent variability of rock masses, data that can be somewhat subjective and the fact that the design technique is non-rigorous. Diederichs and Kaiser (1996) proposed drawing iso-probability contours to account for the uncertainty inherent in design limits. Mawdesley (2002) and Mawdesley et al (2001) calculated iso-probability contours for the stable, failure, major failure, and combined failure and major failure cases in the extended database (see Figures 3.5 and 3.6).


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Figure 3.5: Isoprobability contours for stable cases (Mawdesley et al 2001)

Figure 3.6: Isoprobability contours for combined failure and major failure cases (Mawdesley 2002)

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100


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The risks associated with using the technique for open stoping can now be quantified. For example, the current boundary separating stable from failed cases (see Figure 3.4) indicates a 60% probability of the stope surface being stable, a 40% probability of either a failure or a major failure occurring and a 0% probability that continuous caving will occur (Mawdesley et al 2001). 3.3.3 Application of Mathews’ Method to the Prediction of Caveability

Purpose

There are good reasons to seek to extend the Mathews stability graph approach to the prediction of caveability. As has been noted, Mathews’ approach provides detailed guidelines for the determination of the adjustment factors used. This reduces the subjectivity and degree of personal experience involved in determining the required factors in comparison with Laubscher’s caving chart method. Despite this, it must be acknowledged that a significant degree of subjectivity will always be present in applying techniques based on rock mass classifications. Nevertheless, Mathews’ approach does appear to offer some benefits in this area. Furthermore, a very large database on stope surface stability using Mathews’ approach has now been assembled. This has the potential to aid in the rational delineation of a caving zone. The importance of a large case history database for these empirical design techniques cannot be over emphasised. Caveability data collection

As part of the International Caving Study Stage I, raw data were collected from blocks and panels at the Andina, El Teniente, Salvador, Henderson and Northparkes E26 mines (Mawdesley 2002). Data collected by African Consolidated Mines staff for MRMRs of less than 40 at the Shabanie and King mines were converted to Q values using a local correlation between Q and Laubscher’s RMR or RMRL. A zero stress adjustment and the presence of a flat lying discontinuity set were assumed in order to determine adjustment factors for the weak rock masses in question. Cases in which continuous caving had not initiated were designated as major failures.

A number of case histories for blocky (no continuous bedding planes), strong coal mine roofs having semi-stable spans at hydraulic radii of up to 55 m were collected as part of the study. Although they were not from block or panel caving mines, these case histories should have some relevance to caveability assessment given their large undercut dimensions (Mawdesley 2002). Delineation of a caving zone

The data points for the case histories collected were plotted on the extended Mathews’ stability graph. A logistical regression analysis was used to delineate a caving zone as shown in


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Figure 3.7 (Mawdesley 2002). The line separating major failures from continuous caving on the extended Mathews stability graph does not represent a 100% probability of caving. Insufficient data are available to permit iso-probability contours to be defined accurately. The availability of additional data from well-documented case histories would allow the uncertainty in the design limits to be quantified, giving greater confidence in the use of this technique for predicting caveability.

Figure 3.7: Extended Mathews stability graph based on logistic regression showing the stable and caving lines (Mawdesley 2002)

The influence of in situ and mining induced stresses on the caveability of an orebody has been discussed in Chapter 1. The effects of stress were readily apparent in the back-analyses of caving case histories carried out using the extended Mathews stability graph method. For example, some cases such as the Esmeralda section at El Teniente and Northparkes E26 Lift 1 had similar unadjusted rock mass classification ratings but exhibited very different caving responses. The Esmeralda section caved continuously at a hydraulic radius of 27 m (1.1 hectares of undercut area), while Northparkes E26 Lift 1 did not cave continuously at a hydraulic radius of 44 m (3.2 hectares of undercut area). However, the in situ stresses and therefore the induced stresses in the cave back for Esmeralda were much higher than those for Northparkes E26 Lift 1. Therefore, the stability numbers (analogous to MRMR in Laubscher’s method) in the two cases were significantly different despite the fact that the rock masses were perceived to have similar strengths. The difference in the stability numbers was reflected in the different caving responses.

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Summary

A review of the Mathews stability graph method suggested that it might provide an alternative method of predicting caveability, even though no caving case histories were documented. Caving case histories were collected and analysed and a caving zone delineated for an extended version of the Mathews stability graph (Mawdesley 2002). Nevertheless, more caving and transitional caving data are required to increase confidence in the use of this technique for predicting caveability. With additional data it should be possible to quantify the current uncertainty in the design limits. 3.4 NUMERICAL MODELLING APPROACHES

As noted previously, numerical modelling holds the possibility of providing a more rigorous assessment of caveability than the empirical methods just described. Numerical methods are used widely to solve stress-deformation boundary value problems in mining geomechanics for which analytical solutions cannot be obtained. This may occur when the boundary conditions, including the problem geometry, cannot be described by simple and tractable mathematical functions, the governing partial differential equations are non-linear, the problem domain is inhomogeneous, or the constitutive equations of the rock masses concerned are non-linear or insufficiently simple mathematically (Brown 1987). Clearly, most of these conditions apply in the analysis of caving which, by definition, involves non-linear and discontinuous rock mass behaviour. Numerical approaches have the advantage over empirical methods of assessing caveability of being able to treat the complex mechanics of the problem more completely and accurately. Numerical methods are also able to allow for the presence of a number of geomechanical units and for inhomogeneous responses within the problem domain. They can model major faults explicitly and represent undercut shape more completely than does the use of the hydraulic radius, or shape factor, as in the empirical methods discussed in the preceding sections. However, as Brekke and Howard (1972) have noted, rock masses are so variable that the chance of ever finding common sets of parameters and of constitutive equations that are valid for all rock masses is quite remote. Although great advances have been made in numerical modelling capabilities in the intervening years and systematic approaches to numerical modelling of rock mechanics problems have been developed (eg Starfield and Cundall 1988), Brekke and Howard’s comments remain valid. For this reason, numerical models must be shown to accurately reproduce observed caving behaviour before being used to predict caveability in practice. Most numerical models treat the rock mass as a continuum or an equivalent continuum allowing it to be assumed that the material response may be described by the equations of the


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theories of elasticity or plasticity. Special methods of solution are required for those rock mechanics problems involving the interaction of discrete blocks of rock in which the ratio of the block size to the size of the problem domain is such that equivalent continuum behaviour may not be assumed. The most powerful and versatile method available for simulating such discontinuum behaviour is the distinct element method developed by Cundall (1971, 2001). Because of the inherently discontinuous nature of the caving process, discontinuum or distinct element approaches are intuitively attractive for use in the assessment of caveability. They also have obvious application in the modelling of particle flow as will be discussed in Chapter 7. Even if their complexity and computationally intensive nature preclude them from use in solving industrial scale problems, these methods should provide an important aid to our understanding of the caving process as illustrated by the simple example shown in Figure 1.7. An illustration of the application of the state-of-the-art three dimensional particle flow code PFC3D (Itasca 1998a) for caveability analysis, carried out by Dr Loren Lorig of the Itasca Consulting Group Inc as part of the International Caving Study Stage I, is summarised in Section 3.6 below. But before that, the application of a less numerically intensive axisymmetric continuum model to caveability prediction will be presented. Here again, the modelling described was carried out by Dr Loren Lorig as part of the International Caving Study Stage I. 3.5 AXISYMMETRIC CONTINUUM MODEL

3.5.1 Model Formulation

The conceptual caving model shown in Figure 1.8 illustrates the main behavioural regions of a propagating cave. The characteristics of each of these regions were described in Section 1.2.2. It is important to note that the boundaries between these regions are diffuse rather than sharp and that they may be in different locations in different cases. Clearly, the rock mass undergoes a gradual reduction in strength from its in situ state to its caved state. In particular, the cohesion of the in situ rock mass appears to reduce during the caving process until it reaches the caved state in which the rock mass is essentially cohesionless. This observation suggests that the rock mass can be represented as a strain-softening material in which the rock mass cohesion diminishes from an initial in situ value to zero.

The numerical model of caving presented here attempts to capture many of the important features of the conceptual caving model. One important assumption in the formulation of the numerical model is that the rock mass can be treated as a continuum in the sense discussed in Section 3.4. It is further assumed that the rock mass shear strength is limited by a simple Mohr-Coulomb failure criterion. In the discussion presented here, the two-dimensional explicit finite-difference code FLAC (Itasca 1998b) is used to model the caving process. It is possible to extend the analysis to three-dimensions using an appropriate numerical model such as FLAC3D (Itasca 1997b).


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The axisymmetric continuum model used to study the caving process is shown in Figure 3.8. The model considers an undercut of cylindrical shape of radius R and height ho excavated at a depth, H, below the ground surface. The initial state of stress is assumed to be lithostatic. The vertical stress is given by σv = γ z (where γ is the unit weight of the material and z is the depth), and the horizontal stress is given by σh = Ko σv (where Ko, representing the ratio σh /σv, is assumed constant). Stress boundary conditions are imposed at the undercut level. The ‘support’ pressure, p(t), is reduced monotonically to simulate the extraction of material and study the resulting extension of the plastic region that develops on top of this level (Figure 3.8b). The plastic region is then interpreted as a volume of caved material available for extraction.

The material on top of the undercut area is considered to be elasto-plastic, with a stress-strain response characterised by strength softening, as shown in Figure 3.9a. Two linear Mohr-Coulomb yield surfaces define the ‘peak’ and ‘residual” stages, respectively (Figure 3.9b). The elastic behaviour is characterised by two elastic constants, the shear modulus, G, and the bulk modulus, K. The plastic behaviour is characterised by six plastic parameters: the peak friction angle, φ; the peak cohesion, c; the residual friction angle, φr; the residual cohesion, cr; a parameter s

critε , which defines the rate of softening (ie the accumulated plastic strain for which the residual stage is achieved); and the peak dilation angle, Ψ, which corresponds to a non-associated flow rule.

(a) (b)

Figure 3.8: Schematic representation of(a) the axisymmetric model, and (b) the evolution of the undercut pressure and height


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(a) (b)

Figure 3.9: (a) Constitutive behaviour for the material, with σ3 assumed constant, and

(b) yield conditions for intact and caved materials

In order to implement a material extraction scheme in the continuum model, the undercut area is considered to be filled with a fictitious material. The elastic parameters of this material are assumed to be a fraction, fS, of the values specified for the rest of the rock mass so that the shear and bulk modulii for the undercut region are GF = fS G and KF = fS K, respectively. As noted previously, the extraction process at the undercut level is simulated by monotonically reducing the vertical pressure, p(t), at this level. Figure 3.10a shows how the support reduction is performed in steps. (The continuous type of reduction represented in Figure 3.10b has only descriptive purposes.) It is important to note that the time variable, t, included in the diagrams does not correlate directly with the actual physical time; this variable appears naturally from the explicit time formulation used by FLAC. In order to extrapolate the results obtained from the model, the number of steps (rather than the variable t) must be associated with the extraction rate at the site. Results given by the model for specified undercut geometries and mechanical properties can be used to analyse the caveability conditions above the undercut. For example, if the evolution of the cave height with continuous reduction of the undercut support pressure is as in Curve A in Figure 3.11, then the cave is stable - ie the cave will stop even as extraction of material is continued at the undercut level. On the other hand, if the evolution of the cave height is as in Curve B in Figure 3.11, the cave is unstable - ie the cave will continue to grow as material is removed from the undercut.

scrε


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(a) (b)

Figure 3.10: (a) Stepwise reduction of the undercut pressure, and

(b) details of the pressure evolution within a reduction step

Figure 3.11: Evolution of caving height for stable and unstable cases


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3.5.2 Material Parameters

Using the Mohr-Coulomb model for rock, the only realistic assumption for the caved mass is zero cohesion (cr = 0), zero dilation (Ψ = 0) and a non-zero friction angle (φr ≠ 0). The material, perhaps containing some large blocks, ‘flows’ towards the drawpoints and is thereby in a state that resembles the critical state of soil, characterised by a constant apparent strength and zero volume change. Because the material is broken, the mass tensile strength is zero (σt = 0), which implies a cohesion intercept of zero. It is also likely that the apparent tangent modulus of the caved material is considerably less than that of the “intact” material. The “uncaved” material (termed ‘intact’ for brevity), although containing fractures, must have a certain cohesion (c ≠ 0), because it is observed to form stable arches that exist above voids in the caved material. The primary mode of failure of the intact material is by extension fracturing parallel to the free surface, unless it is very weak (leading to deeper shear failures). Good material models for the spalling mode of failure that occurs during extension fracturing are not presently available. However, it appears (Cundall 1995, Cundall et al 1996) that the Mohr-Coulomb model can approximate this mode of failure given a dilation angle, ψ, greater than the friction angle, φ (eg ψ = 60°). In this case, active yielding is confined to the surface of the material, because the dilation causes a build-up of isotropic stress in the interior elements, thereby inhibiting failure. Although the yielded region appears to be a shear band when using the Mohr-Coulomb model, the real mechanism is surface spalling not shearing.

The ‘transition’ from the intact to the caved material state requires some amount of strain. This transition can be captured by making the cohesion, c, dilation, ψ, and tensile strength, σt, decrease as functions of accumulated plastic shear strain. In the absence of any experimental data, the function is taken to be linear, and the parameter that determines the gradient is the intercept, s

critε , of the softening slope with the strain axis — ie according to Figure 3.9a, the strain necessary for the strength and dilation rate to decrease to zero. The tangent bulk and shear moduli are also assumed to decrease according to a linear relation.

For a simulation in which material softening is used and the response involves shear localisation, the results will depend on the element sizes. However, it is possible to compensate for this form of mesh-dependence. Consider a displacement applied to the boundary of a body. If the strain localises inside the body, the applied displacement appears as a jump across the localised band. The thickness of the band contracts until it is equal to the minimum allowed by the grid - ie a fixed number of element widths. Thus, the strain in the band is znu Δ= /ε , where n is a fixed number, u is the displacement jump, and zΔ is the element width. If the softening slope is linear (ie the change in a property value, pΔ , is proportional to strain), the change in property value with displacement is pΔ / uΔ = s / n zΔ , where s is the input softening slope. In order to obtain mesh-independent results, a scaled softening slope can be input such that zss Δ′= , where s′ is constant. In this case, pΔ / uΔ is independent of zΔ .


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If the softening slope is defined by the critical strain (see the previous discussion), then zs

crit Δ∝ /1ε . For example, if the zone size is doubled, the critical strain must be halved for comparable results. For purposes of illustration, simulations were carried out using the parameters listed in Table 3.1.

Table 3.1: Parameters used in axisymmetric continuum analysis

Property Value Comments

Unit weight (γ ) 25 kN/m3 Density = 2585 kg/m3

g = 9.81 m/s2

Young’s modulus (E) 20 GPa

Poisson’s ratio (ν) 0.22

Peak friction angle (φ ) 35° Same as residual friction angle

Unconfined compressive strength (peak) 15 MPa Zero residual strength

Peak dilation (ψ) 60° Zero residual dilation ( 0oψ = )

Critical plastic strain (scritε ) 0.01 Varied (see Figure 3.9a)

Vertical stress at extraction level 12.13 MPa

Horizontal stress at extraction level 18.19 MPa

fp 0.1 Extraction factor

fs 20 Ratio of intact to broken modulus

ft 12 UCS/tensile strength

(residual ft = 0)

3.5.3 Results

Two mesh sizes were used: coarse (with 38 x 40 zones), and fine (with 76 x 80 zones). In both cases, the same physical dimensions of the grid and undercut area were used. All results show a similar pattern of deformation and failure. Consider, for example, Figure 3.12 which shows the extent of plastic yield for the fine model with s

critε = 0.01 and 2000 equilibrium cycles per excavation step. There is a strong stress discontinuity (see Figure 3.13, which shows the final state) between the caved region and the intact region, and this discontinuity migrates upwards with continued extraction. However, at some point, the discontinuity stops growing (see the history of cave height shown in Figure 3.14), and a stable roof of intact material forms. The material under this roof is generally in its residual state (ie with zero cohesion and dilation), carrying stresses that are of the order of gravitational stresses in a pile of granular material. The material flows with each extraction step (see Figure 3.15 which shows displacement vectors after 360 excavation steps). The physical interpretation is that a steady-state solution has been obtained, with the cave stabilised at a specific height and the material below the stable roof able to flow out freely whenever extraction is performed. Because FLAC operates (in this case) in small-strain mode, the material cannot actually be moved out of the caved region, but the


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simulation provides the evidence (ie constant pressure, zero strength, and gravitational stresses) that the caved mass is in steady-state flow. For practical application, it would be desirable to devise a way of reproducing the evolving geometry as well.

FLAC (Version 3.40)

LEGEND

13-Oct-98 13:57 step 720000 -7.333E+01 <x< 5.733E+02 -8.083E+01 <y< 5.658E+02

Boundary plot

0 1E 2

Plasticity Indicator* at yield in shear or vol.X elastic, at yield in pasto at yield in tension

0.000

1.000

2.000

3.000

4.000

5.000

(*10^2)

0.000 1.000 2.000 3.000 4.000 5.000(*10^2)

JOB TITLE :

Itasca Consulting Group, Inc.Minneapolis, Minnesota USA

Figure 3.12: Plasticity indicators (note that tension is indicated at the top of the cave;

the undercut width is 100 m)

Figure 3.13: Stress tensors (note zero stresses under roof of intact material)


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Figure 3.14: History of caving height (note that the cave stabilises at 200 m)

Figure 3.15: Displacement vectors (note that 12 m of movement has occurred)


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The stable cave height is strongly dependent on the softening modulus or the brittleness of the rock. As the material is made more brittle, the cave height, H, increases. However, for each level of brittleness, the results appear to be nearly independent of the mesh size, provided that the scaling outlined above is used. Table 3.2 lists the results of the six simulations performed. Note that each pair of ‘coarse-fine’ rows should give comparable results, since the critical strain is scaled in the same ratio as the zone size.

Table 3.2: Cave height as a function of brittleness, increasing downwards (note that the results are comparable within each pair of ‘coarse-fine’ rows)

Grid Critical strain, scritε Cavern height, H [m]

Coarse 0.01 160

Fine 0.02 150

Coarse 0.005 200

Fine 0.01 205

Coarse 0.0025 224

Fine 0.005 250 Although the cave height, H, seems to be reasonably predictable, the ‘shapes’ of the caved regions are not believed to represent necessarily the shapes that occur in reality. There are likely to be interlocking blocks of rock in the roof of the cave that resist bending with restraining moments. Such effects are not present in the FLAC model using a classical continuum formulation, so the sharp corners observed numerically would probably be smoothed out in reality if the interlocking effects were taken into account. It is possible to allow moments (generated by relative rotations of blocks) to occur by using the Cosserat continuum formulation, as implemented in a special version of FLAC (Dawson 1995). Dawson also demonstrated that the Cosserat formulation removes the mesh-dependence associated with softening material. Although this approach would eliminate the need to scale the softening slope, much bigger meshes would be necessary, because localised bands are spread over many more zones. 3.6 PFC3D DISCONTINUUM MODEL

3.6.1 Modelling Approach

As discussed above, current axisymmetric continuum models show a strong sensitivity to critical strain. It is anticipated that Particle Flow Code (PFC) models can be used to eliminate this sensitivity and the accompanying scaling process required by continuum models. PFC models have an apparent critical strain ‘built in’ as a macroscopic property which can be changed by assuming different microscopic properties and contact models. It is also anticipated


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that PFC models will provide an improved predicted shape of the caved region. Ultimately, it may be possible for information gained from PFC analysis of the caving process to be incorporated into a continuum model. The advantage of continuum models is that they are easier to set up, run and interpret than PFC models. In order to minimise the initial computational effort but still be able to evaluate the PFC approach, quarter-symmetry was used in the simulations reported here. This is part of a ‘start small and simple — then add size and complexity’ approach. Because a particle model represents a heterogeneous material, the use of symmetry is, strictly speaking, not valid. However, the errors introduced should not be large. 3.6.2 Model Description

In a PFC3D (Itasca 1998a) model, the rock mass is represented by an assembly of densely packed spherical particles bonded at their contact points. The attractiveness of the PFC approach is that a simple constitutive relation on a micro (ie particle) scale produces a rich response on the macro scale in terms of linear and non-linear material behaviour (eg Potyondy et al 1996). Thus, laboratory-determined stiffness, strength, and post peak response (eg softening rate) are ‘automatically’ expressed by the PFC model by adjusting the micro properties. In this context, the properties needed in PFC3D are particle-contact normal and shear stiffnesses (kn and ks), normal- and shear-contact bond strengths (FC

n and FCs), and the

friction angle between particles. The micro-mechanical properties of this assembly were calibrated to the basic elastic and strength properties of the rock mass given in Table 3.3. These properties were selected to represent a relatively weak rock mass, in order to ensure caving and test the application of the model.

Table 3.3: Rock mass elastic and strength properties

Young’s modulus 13 GPa

Poisson’s ratio 0.25

Unconfined compressive strength (UCS) 6.5 MPa

Tensile strength 1.4 MPa

Bulk density 2500 kg/m3

The initial stress was assumed to be gravitational, producing a vertical stress of about 12 MPa at the undercut level. Particles located along the symmetry planes were free to move in all directions except normal to the plane.


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Two models were analysed for demonstration purposes. The major difference between the two models is the level of discretization used. In the ‘coarse’ model (containing about 5,300 balls), the mean particle radius in the caving region is 4 m; in the ‘fine’ model (containing about 35,000 balls), a mean radius of 2 m is used. The coarse model is small enough to provide a relatively quick evaluation of the concept of a PFC3D model for caving analyses, while the fine model provides a measure of the effects of particle size (in terms of caving mechanics as well as computational aspects related to model size). Both models have a lateral extent of 600 x 600 m. The depth to the undercut is assumed to be 500 m. To save additional computational effort in the fine model, the vertical extent was made 250 m, with the additional 250 m being represented simply as an overburden pressure. The height of the coarse model was 500 m.

Undercuts in both models are made sequentially to radial distances of 30 m, 60 m and 90 m. The height of the undercut is made comparable to the mean diameter of particles within these radii. The effect of the undercut is simulated by gradually reducing the particle radius of all particles within the undercut and deleting these particles upon reaching some minimum radius. As a result, bonds are broken between particles in the assembly of particles above the undercut, simulating the breaking/fracturing of the cave back. Any caving particle that moves into the undercut will subsequently have its radius gradually reduced, and it eventually will be deleted. This approach minimises inertial effects of suddenly removing material in the model; it is believed to be a reasonable approximation to the creation of an undercut and the subsequent drawing of broken ore. Figure 3.16 illustrates the fine and coarse models. Only particles on and behind the diagonal vertical plane are shown. The significance of the dark and light colour bands is only to illustrate the kinematics of the caving. The radial gradation of particle size is necessary to minimise the computational effort (ie model size). Figure 3.17 shows a close-up view of the regions in the vicinity of the undercut. The numbered circles above the undercut indicate the locations of spherical volumes within which the histories of the average stress and the strain tensors are determined. This information may be useful when comparing results from PFC3D models with results from continuum models.


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(a)

(b)

Figure 3.16: Illustration of coarse and fine discretizations used in the PFC3D caving model


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Figure 3.17: Close-up of caving region showing spherical volumes for

measurement of average stress and strain

3.6.3 Results of Model Observations

Upon initiating the undercut, bonds are broken between particles representing the rock mass. If conditions become stable, no further breaking of bonds will occur. Thus, a history of accumulated bond breakage can indicate whether the cave has stabilised. In addition, particles with a certain downwards vertical velocity are indicated as being unstable, and, thus, represent ore that potentially could be mined. Because each particle represents a specific volume of ore, recording a history of the accumulation of such particles gives an estimate of (1) stable/unstable conditions, and (2) the amount of ore that can be extracted from the ore column. The coarse model was run sufficiently to evaluate the effects of 30 m and 60 m undercut radii. At 30 m, the cave quickly stabilised, with the cave back ending about 30 m above the undercut. A subsequent increase in the undercut radius to 60 m precipitated further caving. At one stage, the draw rate was set to zero (ie the particle radii were kept constant in the undercut). This resulted in stable conditions, with the cave back ending about 90 m above the undercut. With continued drawing, the 60 m undercut radius appears to have affected the stability of the rock mass up to a height of about 200 m above the undercut. The rate of bond breakage is diminishing as shown in Figure 3.18, suggesting that the cave may stabilise. Figure 3.19, showing the history of the caved volume (in m3 along the ordinate), indicates that the cave is approaching stable conditions.


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Figure 3.18: Number of broken bonds (y-axis) versus computation steps (x-axis)

60 m Undercut Radius


Draw rate set to zero

Draw rate restarted

Figure 3.19: Volume of caved rock, m3 (along y-axis) versus computational steps

(along x-axis)


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In the fine model, stable conditions also developed for an undercut radius of 30 m. Figure 3.20 shows a close-up of the caving region. Only particles on the diagonal vertical plane are shown. The force distribution for particles on the plane is also illustrated. Note the arching of these forces around the cave. At this point, it appears that the rock mass is unstable to a height of about 100 m above the undercut (determined from an examination of velocities). However, it is too early to tell whether the cave also will stabilise for a 60 m undercut radius. An example of the history of average vertical and horizontal stresses predicted in the fine model is shown in Figure 3.21 for the spherical volume labelled 1 in Figure 3.17.

Figure 3.20: Close-up of caving region during caving showing forces arching around the unstable rock mass (only particles on the diagonal vertical plane are shown)

Stress in Pascal (compression negative) is given along the ordinate and computational cycles along the abscissa. As the vertical stress decreases because of the 30 m undercut, the two horizontal stresses increase slightly. Eventually, all stresses become constant as a result of the stable conditions. Upon increasing the undercut radius to 60 m, the stresses decrease dramatically, indicating that this material is no longer providing support for the overlying rock mass but is caving.


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Figure 3.21: History of the average vertical stress (σzz) and horizontal stresses (σxx,

σyy) for measurement sphere No. 1 (stress in Pa along the y-axis; computational steps along the x-axis)

3.6.4 Future PFC Modelling of Caveability

The model discussed here represents a starting point in understanding the usefulness of PFC3D to study caveability. To optimise the rate of understanding, the models have been kept relatively small, taking advantage of planes of symmetry. Larger, fully 3D models would have to be analysed to study full scale problems and the sensitivity of these models to conditions and material properties would need to be understood. Nevertheless, the results presented here illustrate the value of the discontinuum approach in studying the mechanics of caving. As noted earlier, PFC models have an apparent critical strain ‘built in’ as a macroscopic property that can be changed by assuming different microscopic properties and contact models. The value of the critical strain should be measured and recorded as a PFC model property in future PFC analyses, even if its definition is not exactly the same as the continuum definition. Characterising the rock mass as an assembly of bonded particles may result in a material in which the tensile strength is too high relative to the compressive strength. One way of overcoming this problem is to cluster particles. Weak bonds may be assigned between clusters, while the bond strength internal to the cluster remains relatively higher. Currently, the initial in



Stable Conditions

σzz

σyσx


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situ stress is taken to be hydrostatic. However, the model must be capable of evaluating caveability for a range of in situ stresses. Functions are now available that allow the assignment of different initial in situ stresses in the PFC3D model. The effect of particle size on caveability also needs to be evaluated. The coarse model used in the analyses reported here may be too coarse to serve as a reasonable benchmark in this respect.

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CHAPTER 4

FRAGMENTATION ASSESSMENT 4.1 INTRODUCTION

he overall success and profitability of a block caving operation will depend to a significant extent on the fragmentation produced in the orebody during the caving process. The design and operating parameters influenced by fragmentation include (eg

Laubscher 1994, 2000): • drawpoint size and spacing; • equipment selection; • draw control procedures; • production rates; • dilution entry into the draw column; • hangups and the need for secondary breakage/ blasting with associated costs and damage; • staffing levels; and • subsequent comminution processes and costs. The prediction of rock fragmentation during block caving requires understandings of the natural fragmentation of the rock mass and of the fragmentation processes that take place in the draw column. It is generally accepted that there are three levels of fragmentation, commencing with the in situ blocks and then progressing to primary and to secondary fragmentation. In situ fragmentation is represented by the blocks that are naturally present within the rock mass before any mining activity takes place. They are defined by the pre-existing discontinuities. As the undercut is mined and caving is initiated, the blocks in the vicinity of the cave back that separate from the cave back define the primary fragmentation. The fragmentation that occurs subsequently as the blocks move down through the draw column to the drawpoints in known as secondary fragmentation.

It is desirable that fragmentation models be developed to provide reliable estimates of fragmentation for use in mine planning. The basic requirement of any such model is to provide a measure of the range and distribution of the sizes of the rock blocks expected to be produced

T

Chapter 4: Fragmentation Assessment

157

at the various stages of fragmentation and, in particular, those finally reporting to the drawpoints. Since the production equipment and the drawpoint layouts must accommodate the resulting blocks, knowledge of their shapes as well as their sizes will be of value. It is generally accepted that because of the limited understanding of the mechanisms involved and, for good practical reasons, the lack of availability of sufficient data, the development of a complete, mechanistically based fragmentation model is not currently plausible. In this chapter, the factors influencing fragmentation, methods of fragmentation measurement and the existing methods of predicting fragmentation are reviewed. The most widely used model for predicting fragmentation, Block Cave Fragmentation or BCF, is discussed in Section 4.5. This method depends on a large number of experiential relationships and interpretations. As part of the International Caving Study Stage I, an attempt was made to address some of the perceived weaknesses of the BCF model. The revised model was found to still give results that are considered to be not entirely satisfactory in some respects. Accordingly, a new model for identifying in situ blocks and predicting primary fragmentation known as JKFrag was developed from first principles as part of the International Caving Study Stage I (Eadie 2002) and is described in Section 4.6. This model acknowledges the interdependence of the network of discontinuities in the rock mass and uses a tessellation approach to define primary blocks. Finally, a way forward in the development of a complete fragmentation model, including secondary fragmentation, will be suggested in Section 4.7. 4.2 FACTORS INFLUENCING FRAGMENTATION

As noted in Section 4.1, three levels of fragmentation are recognised, in situ, primary and secondary. There is a natural progression through each stage during the caving process with each level forming the starting point for the next. At each stage there is a compounding of the effects of the influencing factors and failure mechanisms involved to produce the progressive break-down of the in situ blocks. By definition, in situ blocks are completely determined by the network of discontinuities pre-existing in the rock mass. More precisely, the sizes and shapes of these blocks are a direct result of the geometry of the open discontinuities present within the rock mass. Incipient or healed discontinuities having finite shear and tensile strengths do not provide faces of in situ blocks, but rather represent planes of weakness within the rock mass on which separation may occur in the primary and secondary stages of fragmentation. The orientation, size, spacing, condition and termination are the main parameters used to describe the overall network of discontinuities. The collection, processing and presentation of discontinuity data were discussed in detail in Chapter 2. The accuracy of the prediction of block shapes and sizes, at the in situ and subsequent stages of fragmentation, is clearly subject to the ability to model these parameters adequately.


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Once caving has been initiated, primary fragmentation will result from the loading conditions imposed on the rock mass in the vicinity of the cave back. Most failures at this stage can be expected to occur on the existing planes of weakness, but under high stress or stress caving conditions, fracture of intact rock may also occur. The extent of these failures will depend on the strengths of both the discontinuities and the incipient rock blocks relative to the magnitudes and orientations of the imposed stresses. The primary fragmentation size distribution produced in this case is likely to be finer than in the case of subsidence caving in which gravity rather than induced stresses causes the detachment of blocks from the cave back.

To a large extent, the network of pre-existing discontinuities will govern the formation of blocks during primary fragmentation. As with in situ blocks, not all of the discontinuities existing in the rock mass will define primary blocks. A larger subset of the discontinuities will be involved in secondary fragmentation as more of the closed and healed discontinuities fail under the imposed stresses. New fractures will also be produced through the failure of intact rock bridges defined by the existing discontinuities. The geometry of the network of existing discontinuities together with the induced fractures will ultimately define the sizes and shapes of the primary blocks. Figure 2.30 shows two cross sections of a rock mass that are statistically equivalent in terms of joint orientation, spacing and size. Inspection of the blocks expected from these diagrams illustrates the role of network geometry in the defining the block size distribution. This highlights the importance of the collection and analysis of discontinuity data including the termination properties discussed in Chapter 2. The conditions of the discontinuities and their shear strengths must also be considered. As has been indicated, secondary fragmentation will occur as the caved ore resides in, and passes through, the draw column. The nature and degree of secondary fragmentation can be expected to vary with the stress regime within the caved mass, the composition and mechanical properties of the orebody, the rate of draw, the height through which the material is drawn and the residence time in the draw column. In general, the mechanisms of secondary fragmentation can be expected to include some or all of the following: • extension of pre-existing discontinuitues; • opening of filled of healed discontinuities; • opening along bedding or schistocity planes; • crushing under superimposed weight; • compressive (shear) failure of blocks under the influence of arching stresses within the

cave; • failure of individual blocks by induced tension produced by point or line loading at inter-

block contacts within the caved mass; • bending failure of elongated blocks; and • abrasion or “autogenous grinding” of block corners and edges to reduce block sizes and

produce fines.


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4.3 FRAGMENTATION MEASUREMENT

4.3.1 Overview

If reliable methods of predicting the fragmentation produced during caving and draw are to be developed, methods of measuring fragmentation are required to validate them. The accurate measurement of fragmentation in caving mines is difficult to achieve. Most of the techniques to be introduced here had their origins in the measurement of the fragmentation arising from production blasting. However, in some cases they may be adapted for use in caving mines. Because of their potential for future development, the major emphasis here will be placed on digital image processing methods. In general, sieving, physical measurements, production rate analysis and digital methods may by used to measure or estimate the size distribution of the rock fragments produced by caving or blasting. While sieving and physical measurements may be the most accurate methods, they unfortunately cause disruption to production, are costly and therefore impractical for use for other than the most special purposes. Accordingly, alternative methods have been investigated. Grant and Dutton (1983) and Bhandari and Tawnar (1993) adopted the boulder counting technique. This technique consists of measuring only the "oversize" fragments. It may provide a statistically representative measure of the important top size distribution but obviously does not establish the full size distribution. Another alternative technique is to record the explosive consumption from secondary blasting activities. By monitoring the amount of explosives used for secondary blasting, trends in the number of oversize boulders being produced may be recorded. This technique is obviously closely related to boulder counts and is therefore biased toward representation of the coarse end of the size distribution. Production analysis or statistics may be used in cases where fragmentation has a direct and significant influence on the productivity of equipment. In this method, the production statistics that are influenced by fragmentation are monitored. However, extraneous factors such as equipment condition and operator performance make overall correlation difficult. Grant and Dutton (1983) reported that cycle times and loading rates showed little dependence on fragmentation unless large boulders were encountered in the muck pile. By considering only a section of the production cycle, it may be possible to eliminate some of these factors. The operational importance of the larger block sizes should not be underestimated. It is common practice to identify the percentage of material reporting as blocks of 2 m3 or larger. (A block of this size can be handled by a 6 yd3 LHD. For an 8 yd3 LHD, the maximum block size becomes 3 m3.) Table 4.1 shows a classification of rock fragmentation sizes and their potential operational effects suggested by N J W Bell based on experience in the chrysotile asbestos mines in Zimbabwe (Laubscher 2000).


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Table 4.1: Rock fragmentation sizes and their potential effects (Laubscher 2000)

Rock

Fragmentation

Size

Potential

Effects

Length

Range

Metres

Mean

Length

L

metres

Mean Volume

L x L/2 x L/2

m3

Maximum

Volume

m3 A

B

C

D

E

F

G

100% through 1.5 m x 0.3 m

grizzly

100% into LHD bucket

Hang-up in drawpoint throat

High hang-up

Drawbell blocker

Double drawbell blocker

< 0.5

0.5 to 1.0

1.0 to 2.0

2.0 to 4.0

4.0 to 8.0

8.0 to 16

>16

0.25

0.75

1.5

3

6

12

24

0.004

0.11

0.8

7

54

432

3456

0.031

0.25

2

16

128

1024

Infinite

It will be apparent from the preceding discussion that the fragmentation produced by block caving is extremely difficult to measure reliably and routinely. Some of these difficulties will be discussed more fully below. The primary fragmentation distribution cannot be measured directly although it may be approximated by measurements made on the broken ore drawn from the drawpoints in the early stages of caving. The finer fragment sizes produced following secondary fragmentation are particularly difficult to measure by methods other than sieving. Figure 4.1 shows a typical example of the fragmentation distribution measured at the Premier Mine, South Africa. It will be noted that the distribution is somewhat irregular and covers a relatively limited range of fragment sizes from 10 m3 down to about 0.25 m3. The distribution of sizes of less than about 0.25 m3 is absent from the distribution, presumably because they weren't measured.

Figure 4.1: An example of the fragmentation distribution measured at the

Premier Mine, South Africa (after Butcher 2002a)


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4.3.2 Digital Image Processing Methods

The only practical method of large-scale fragmentation measurement currently available is digital image processing (DIP) in which photographs or video images are analysed using computer-based image processing techniques. Image analysis is the process by which the size distribution of particles in the material of concern is identified in the image and corrected by stereological methods (Hunter et al 1990). Digital image analysis methods have a number of advantages over the other methods outlined above in terms of: • speed of sampling; • non-disruptive method of sampling; • ability to analyse many samples at a smaller expense; and • ease of practical application.

There are three stages in the image analysis process - sampling, image acquisition and image analysis. Sampling is the process of obtaining “representative” images of the fragmented material being analysed. Image acquisition involves taking images of sufficient quality and resolution for successful analysis. As has been noted previously, image analysis itself is the process by which the size distribution of fragments is identified from the images and corrected by stereological methods. There are potentially several sources of significant error in all vision based granulometry systems - sampling errors, poor edge net fidelity, scaling errors, processing errors, and missing fines. Sampling errors result from systematic bias in the process of taking an image of the fragmentation. They have the potential to be the most serious of all errors. They occur if the camera is pointed at a place in the muck pile where coarse blocks or zones of fines dominate. This topic has been explored by Maerz (1996). Sampling errors are a function of the type, scale and number of images collected. In this context, the type of image collected refers to the location and state of the material being sampled, and to the quality of the images produced. It is essential that a set of representative images of the material be captured. A number of precautions should be taken when collecting images in the field: • when selecting an area of broken rock, the particle boundaries should be clearly visible for

good particle delineation. Sometimes the coarse fragments are partly obscured by fines and are not interpreted as separate fragments;


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• shadows along the fragment edges should be minimized; and • several samples should be taken around each whole area and several zoomed images

should be taken to account for fines. A sufficiently large number of images should be taken to ensure adequate statistical sampling of the muck pile. The number of images required cannot be easily defined. Kemeny (1994) suggests that the number is usually between 8 and 20, while Palangio and Franklin (1996) recommend that at least 8 to 12 images be acquired. Smaller fragment size fractions require less material to be sampled for accuracy than do larger size fractions. Poor delineation of individual fragments also produces erroneous results. Poor delineation arises from a combination of two sources: • poor images (eg contrast too low or too high, image too grainy, lighting inadequate or

uneven), or the size of the fragments in the image is too small; and • highly textured rock in which shadows and/or colouring on the surface of the fragments are

as prominent as the shadows between rock fragments. Where the smallest fragments in a distribution are not delineated on the image, either because they are too small relative to the image to be resolved or they have fallen in and behind larger fragments, there is clearly a bias towards over-representing the coarse end of the size distribution. Where the distribution has a relatively narrow size range (well sorted or poorly graded), problems of this type do not normally arise. However, where the distribution has a relatively wider size range (poorly sorted or well graded), typically with size differences of more than one order of magnitude, missing fines can affect the measurement results. 4.3.3 Examples of DIP Systems

A number of digital image processing (DIP) systems and associated analytical techniques are available for use in mining applications. These methods may employ either manual or automatic image input. Manual input of images involves the manual digitisation of the particle outlines from a photograph and as a result is very slow and time consuming. An advantage of this method is that it allows for the human interpretation of indistinct particle edges that may otherwise cause errors. A disadvantage is that sampling errors are likely to increase because fewer images can be processed in a given time. Automatic image input allows for more rapid processing of images because the computer identifies particle outlines. This method requires good contrast between the particles and the background in the image. This presents a particular difficulty in some underground mining applications.


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One of the disadvantages of photographic methods is that they can process only the surface fragment distribution and must assume that the surface size distribution is representative of the fragment distribution of the entire pile. This disadvantage certainly applies to images of muck piles taken at caving drawpoints. It may be minimised when the ore is spread evenly to shallow depths on conveyor belts, for example. For this reason, most of the currently available systems are better suited to making fragmentation measurements on conveyor belts than in caving drawpoints. Three of the currently available DIP systems will be introduced below for purposes of illustration. FRAGSCAN

FRAGSCAN is an automatic image processing system developed by Schleifer and Tessier (1996) for assessing fragmentation distributions using images of the visible parts of muck piles. This system differs from other image processing systems in that it does not use edge detection to identify the particles, but uses an algorithm to separate them into a series of size classes, simulating sieving through the corresponding mesh sizes. Surface areas are converted to volumes or weights by assuming a spherical model and using experiments on small-scale rock piles to compensate for overlap. Schleifer and Tessier (1996) initially applied FRAGSCAN to three different problems. The first application was to compute the size distribution of material carried in haul trucks to optimise blasting parameters for crushing operations in a quarry. In this application it was found that the system is sensitive to fragmentation variation. The second application used FRAGSCAN on a conveyor belt to check the proportion of large blocks on the belt. The third application involved the use of still photographs for quality control of rock blocks prior to export. WipFrag

WipFrag is an automated image-based granulometry system that uses digital image analysis of photographs and videotape images to determine the size distribution. WipFrag analyses images from sources such as roving camcorders, fixed photographs or digital files. The system comes with an on-board video amplifier and has manual or automatic gain and offset adjustments to account for lighting effects. In the original WipFrag system, images were most frequently acquired through the use of roving camcorders (Maerz et al 1996). Identification of blocks is done in a two-stage process. The initial stage uses thresholding and gradient operators to detect faint shadows between adjacent blocks. The second stage uses reconstruction techniques to delineate the blocks that are only partly outlined during the first stage (Maerz et al 1996).


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Industry experience with the original WipFrag and other systems has shown that sampling errors may be significant with manual camera operation. As a result, Maerz and Palangio (1999) concluded that a preferred approach was to measure fragmentation on-line either on a conveyor belt or as the material falls off the belt. This led to the development of WipFrag System II for automated on-line digital image processing. This system has a variety of options for communication with process control equipment. It may not be ideally suited for underground application, although on-line time-lapse video photography has been used by INCO to characterise the fragmentation at drawpoints (Maerz and Palangio 1999, Preston and Likdea 1996). Split

Split is an image processing program designed to compute the size distribution of rock fragments from grey scale images at various stages of rock breaking in mining and mineral processing. The source of these images can be a muck pile, haul truck, leach pile, drawpoint, waste dump, stock pile or conveyor belt. The Split imaging program was developed at the University of Arizona, USA, (Kemeny et al 1993) and subsequently advanced and applied in conjunction with the Julius Kruttschnitt Mineral Research Centre (eg La Rosa et al 2001). It operated initially on Macintosh computers but a PC version is now available. The program is based on Image, a public-domain image processing program developed by the National Institutes of Health (NIH), USA (Kemeny et al 1993). Since its original introduction, Split has undergone further development and improvement, particularly in the area of fines recognition and in its software features (Kemeny et al 1999, La Rosa et al 2001). Kemeny et al (1999) describe the basic steps in the Split system in the following terms: 1. Acquire digital images, either automatically or manually. 2. Pre-process the images to correct for lighting problems and to screen for unacceptable

images. 3. Delineate the individual fragments in each image using digital image processing

algorithms (Wu and Kemeny 1992). 4. Apply statistical algorithms to the 2-D particle areas in each image to determine 3-D

particle volumes. 5. Statistically correct 3-D volumes for overlap and shape and determine histogram of

particle volumes. 6. Correct particle volume histogram for fines.


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7. Process multiple images together to get an average distribution (including images taken at different scales).

8. Data output to the screen, hard disk and network. Split is available in two formats. One is a user-controlled version which processes a group of images and the other is fully automated and operates continuously on images obtained from an on-line digital video camera. Both formats use the same algorithms for delineating particles and computing size distributions. The user-controlled version, called Split Desktop, is normally used for processing images from blast muck piles and drawpoints. The fully automated system is called Split Online, and can be used to process images taken from moving conveyor belts or tipping dumpers. 4.3.4 Validation Studies

A range of validation and calibration studies have been carried out on the Split and other digital imaging systems in recent years (eg Kemeny et al 1999, Kojovic et al 1998, Liu and Tran 1996, Maerz and Zhou 1999). Schleifer et al (1999) discussed the difficulties, including sampling, involved in using what are usually small-scale measurements to assess the fragmentation distributions of large and largely unseen masses of broken rock. They point out that this problem is generically similar to that encountered in sampling three-dimensional rock masses for geomechanics purposes discussed in some detail in Chapter 2. For validation studies of Split Desktop which is of particular interest here, images are taken of rock particles in field situations and processed by Split and the rock particles are also screened using traditional methods. Validation studies of Split Online involve stopping the conveyor belt, removing and screening the belt material, and comparing the screening results with the image processing results. In general, belt validation studies are easier to carry out than muck pile validations since the screening of much less material is involved. Kemeny et al (1999) report validation studies using the improved version of Split in which accuracies of less than 10% were achieved for sizes down to less than 1 mm. In 1995, the Noranda Technology Centre, Canada, conducted a series of validation tests. The versions of three systems, Fragscan, WipFrag, and Split, then available were used to measure the size distribution of a backfill muck pile from the Holloway Joint Venture, and the results were compared with those obtained from sieving. A pile of fragmented material was divided into four parts. One part was sieved and the other three were spread out, imaged, and analysed with the three programs. The results of these tests indicated that Split and WipFrag produced results that were similar and closer to the sieving results than FragScan (Liu and Tran 1996). Other results of interest were:


166

• all three systems under-estimated fine material, with FragScan under-estimating fine sizes more than the other two systems;

• for coarser materials, the difference between the FragScan results and results from the other two systems decreased;

• the Split system used lower resolution images and showed improvements when higher resolution images were used; and

• FragScan and Split can be applied to fully automated operations but WipFrag demonstrated more flexibility in the manual mode.

4.3.5 Application of DIP Systems to Caving

In block and panel caving, a proper assessment of drawpoint production and performance history requires the following data: • fragmentation size distribution (from random surveys); • secondary breakage activity (frequency and unit costs); • tonnes drawn between secondary breakage activity; • hangups (type and frequency); • tonnes drawn between hangups; and • drawpoint damage, repair and availability. Most of these factors will be discussed in Chapters 6 and 7. Knowledge of the size distribution of the broken ore reporting to drawpoints is required for the range of purposes outlined in Section 4.1. Digital imaging systems such as those discussed above can be used for estimating the fragmentation size distribution. However, it should be noted that most image processing systems are more suited to open pit environments and to environments having sufficient and even lighting. Provided good quality images of underground drawpoints can be obtained, current digital image analysis systems can produce acceptable results. As an illustration, an example will be given of a digital imaging exercise carried out at a sublevel caving operation. The objective was to assess the fragmentation resulting from different blast designs. This example illustrates the application of the Split system to a case in which the fragment size was relatively small in comparison to the fragmentation typically arising from the natural caving of the stronger orebodies now being mined by block and panel caving methods. In this example, photographs were taken using a Nikon Coolpix digital camera. Initially, a 315 mm long torch was used for scale. It was later replaced by a 910 mm cardboard tube placed horizontally at the bottom of the muck pile. Scaling of the images is essential for subsequent analysis. Photographs were taken 5 m back from the foot of the muck pile. An example of one of the photographs is shown in Figure 4.2a.


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Figure 4.2: (a) Drawpoint image, and (b) segmented image at a sublevel caving drawpoint

As has been noted, the Split system has an automated data analysis module. Nevertheless, a user can edit the results manually. Figure 4.2b is an example of the output binary file showing the delineated particles. Grey areas indicate parts of the photograph that have been edited out of the sizing process and black areas denote parts of the photograph estimated to be fines. From this image, the drawpoint fragmentation distribution shown in Figure 4.3 was calculated using Split. As was noted in Section 4.3.2, several images of the muck pile are required in order to obtain a reliable estimate of the fragmentation distribution.

Figure 4.3: Fragmentation analysis output

Particle Size (mm)


168

The JKMRC applied the Split system at the Northparkes E26 Lift 1 block cave to estimate the drawpoint fragmentation during the early development of the cave. The exercise was terminated following the arrest of caving and the introduction of hydraulic fracturing to induce caving (van As and Jeffrey 2000). In general, provided the lighting is adequate, digital image analysis techniques should give better results for block and panel caving than for sublevel caving because of the larger sizes of the fragments produced by the natural caving of rock masses such as those encountered at Northparkes. As an example, Figure 4.4 is a photograph of a drawpoint in the El Teniente Esmeralda block cave. The automatic segmentation of this image produced by the Split system is shown in Figure 4.5. On the basis of the experience outlined here, it is concluded that the use of image analysis techniques for drawpoint fragmentation assessment in caving mines is feasible provided there is sufficient lighting and minimum air borne dust.

Figure 4.4: Drawpoint photograph, Esmeralda section, El Teniente mine, Chile


169

Figure 4.5: Automatic segmentation of the image shown in Figure 4.4 4.4 IN SITU FRAGMENTATION ASSESSMENT

Over the last few decades, a number of methods of describing and quantifying the in situ fragmentation of rock masses have been developed in mainstream rock mechanics and rock engineering. A major need for knowledge of the sizes of blocks within a rock mass, arises in the design of rock reinforcement systems. As discussed in Section 2.7, Barton et al (1974) defined block size using the ratio RQD/Jn where the parameter Jn is dependent on the number of joint sets believed to be present in the rock mass. Although this ratio has been used in practice for its intended purpose, it provides no information about the range and distribution of block sizes. The ISRM (1978) suggested that a first estimate of block size could be achieved by visual examination. Block size was defined as the average “diameter” of a typical block in the rock mass. The method is subjective and implies that a three-dimensional view can be developed by inspecting an open face. Another early approach was to assume that the rock mass contained several (usually three) sets of persistent discontinuities having known mean spacings from which an average or typical block shape and size could be determined. Sen and Eissa (1991) used this approach with three orthogonal joint sets in developing relationships between volumetric joint count, RQD and joint spacing or frequency. For the model adopted by Sen and Eissa (1991), the volumes, of prismatic, plate or bar shaped blocks in a rock mass may be calculated from:


170

⎟⎟⎠

⎞⎜⎜⎝

⎛++=

233121

1111λλ

λλ

λλJ

Vv

where Jv is the volumetric joint count or number of joints in a unit volume, and λi is the linear frequency of the ith joint set. Here again, no information is provided about the range and distribution of block sizes and so this approach is of little interest in the present context.

Some useful advances have now been made using the rock mass modelling approaches discussed in Section 2.6. Most of the available models attempt to simulate and predict the sizes and shapes of in situ blocks and their distributions. They are based on the rigorous collection, correction and analysis of discontinuity data and have moved progressively in the direction of utilising more of the available information. Models of this type are considered to provide the preferred starting point for the development of more complete fragmentation models. The collection and analysis of discontinuity data was discussed in detail in Sections 2.4 and 2.5. The techniques described provide essential input to the simulation and modelling approaches discussed in Section 2.6. Four particular rock mass models will be reviewed briefly here to illustrate their use in the prediction of in situ fragmentation distributions. Joints is a computer program developed by Villaescusa (1991) for analysing joint set characteristics such as size, location and orientation and for simulating three dimensional rock mass geometry. One of its applications was as a first stage in the prediction of fragmentation, mainly from blasting. The program made important advances in the then state-of-the-art through the use of geometrical probability, statistical theory and mathematical stereology. Joints includes a method of predicting in situ block size distributions through the use of a system of random spheres. A random point is placed within the modelled rock mass and several rays are radiated from this point. The average of the distances along each ray to the first point of intersection with a joint is used as the radius of the sphere representing the size of a block. Although this method implicitly respects the interrelated structure of the jointing in the rock mass model, it has some weaknesses. Figure 4.6 shows a small two-dimensional joint model with a random point and its estimated random sphere. After the application of stress, the outer four traces in the diagram may connect to form a block. The centre trace may not extend enough to divide this larger block. In this situation, the established block would contain an internal discontinuity. This diagram illustrates how the application of the method of random spheres can produce finer fragments than those that would be expected in practice. No information is provided on internal jointing, and because spheres are used, no information can be provided about the expected shapes of blocks.


171

Figure 4.6: Joint model showing the sphere generated from a random point

Blocks is a deterministic model developed by Maerz and Germain (1995) which requires as input the position of any joint along a scanline and the orientation of that joint. The program generates a three dimensional volume around the mapped scanline and each joint is extrapolated into the simulation volume. Each joint divides the volume and every subsequent joint further divides the volume into blocks of varying size. Blocks is biased towards the generation of smaller blocks in the vicinity of the scanline and larger blocks as the distance from the scanline increases. This model assumes infinite persistence and so does not consider the sizes of joints. Fragmentation distributions estimated using Blocks are likely to be highly sensitive to the volumes used for the simulation. All blocks generated in this model can be considered to be fully formed. The program produces no information about internal jointing or the extent of intact rock bridges for consideration in a secondary fragmentation module. Stereoblock is a computer model developed by Hadjigeorgiou et al (1995, 1998). In common with Villaescusa’s Joints program, the model assumes that joints can be represented by circular planes (Baecher et al 1977) and makes use of stereological relationships. This model also makes use of information on the orientations, locations and trace lengths of joints obtained from scanline surveys. From the orientation data, joint sets are identified by the visual inspection of a stereonet plot. Statistical procedures are used to determine Fisher's constant, the mean trace length, mean normal spacing and the relevant standard deviations for each joint set. An arbitrary volume is chosen and joints chosen randomly from the generated joint set data are placed randomly within the simulated rock mass. Several virtual scanlines are run through the simulated rock mass and a stereonet prepared for the simulated model. If the simulated and actual stereonets are in agreement, it is assumed that the simulation is valid and representative. From the simulation, the characteristic in situ block size distributions for the rock mass are determined. The simulations may also be used in reinforcement design (Grenon and Hadjigeorgiou 2000).


172

The JKMRC hierarchical model developed by Harries (2001) has a similar basis to Joints and Stereoblock but uses improved methods of accounting for bias in data collection and applies the concept of discontinuity termination in the simulation process. It probably represents the most detailed and rigorous of the simulation approaches developed so far. Accordingly, it has been used as the basis of an original approach to identifying in situ blocks and modelling primary fragmentation developed by Eadie (2002) as part of the International Caving Study Stage I. This model, known as JKFrag, will be described in Section 4.6. Block Cave Fragmentation (BCF) is a program developed to estimate the sizes of rock fragments reporting to a drawpoint during block caving. The initial BCF concept was developed by Dr D H Laubscher in collaboration with A R Guest and P J Bartlett for application at the Premier Mine, South Africa. The programming was carried out by Dr G S Esterhuizen. Subsequently, the program was developed further and applied in the feasibility study for the Palabora block cave. The program used a simplified technique for determining in situ block sizes and empirical rules to predict how the blocks would reduce in size in a draw column (Esterhuizen 1994). The program was improved during 1998 and 1999 as part of the International Caving Study Stage I (Esterhuizen 1999). The improvements included developing a new in situ block generation module, modifying many of the empirical rules for fragmentation, introducing a hangup prediction module developed on the Palabora project and improving the user interface.

BCF is currently the most widely used method of assessing in situ, primary and secondary fragmentation in block and panel caving. Indeed, no other comparable programs are known to exist. Because of its industrial importance, a detailed account of BCF will be given in Section 4.5. 4.5 BCF: A PROGRAM TO PREDICT BLOCK CAVE FRAGMENTATION

4.5.1 Modelling Approach

BCF is an expert system program incorporating analytical and empirical rules describing the processes and factors that play a role in caving fragmentation. Many of the empirical rules used are based on the experience of Dr Dennis Laubscher.

The program consists of three main modules: • The first module calculates primary fragmentation based on the rock strength, joint

geometry and spacing statistics and the field stresses.


173

• The second module calculates secondary fragmentation taking into consideration the aspect ratios of blocks, the block strength, the cave pressure, the stresses induced by arching in the draw column, the draw rate and the height of draw.

• The third module considers the potential for hangups to occur in drawbells using the

secondary fragmentation blocks as input.

The output of BCF is a set of graphs depicting the size distributions and other statistics of the rock blocks at any stage of draw. Summaries are given of the types and numbers of hangups that can be expected in the drawpoints.

4.5.2 Primary Fragmentation

The primary fragmentation module makes use of joint set data, induced stresses and cave face orientation to simulate the way in which the in situ blocks part from the surrounding rock mass. The primary fragmentation module requires the following input data: Jointing: The mean dip and dip direction of each joint set, the range of dips and dip directions, the mean, maximum and minimum joint spacings, the distribution type of the joint spacing and the joint condition are required. If available, the user may also enter the joint trace lengths of each set. Rock mass strength: The strength of the rock mass is required to calculate potential stress spalling during primary fragmentation and to determine whether blocks will split when arches are formed in the draw column. The intact rock strength, rock mass classification and parameters for the Hoek-Brown empirical failure criterion (Hoek and Brown 1980, 1997) are required. Cave face orientation: The orientation of the cave face determines the orientations of the stresses that are active during caving. The dip and dip direction of the cave face are required. Stress: The stresses in the cave face are required for the program to decide whether clamping, shear or tensile separation will occur along the joint surfaces that form a block. The stresses are also used to consider the formation of stress fractures in the rock mass. The magnitudes of the stresses in the solid rock, just inside the rock face, are required. These stress levels may be determined by numerical modelling or in situ measurements.


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Formation of primary rock blocks

The primary rock blocks are assumed to be bounded by up to six joint surfaces selected from the defined joint sets. The program can accommodate up to six joint sets. If stress fractures are formed, they are considered to constitute an additional joint set, parallel to the caving face. Primary rock block definition starts with the selection of three joint planes to form a block corner, based on the frequency of occurrence of the joints. Once a block corner has been defined, the remaining block faces are determined by the spacing of the joints, the trace length (if known), the shear strength along the joint surfaces and the tensile strength across the joint surfaces. Depending on the stress field, a joint may either shear and form a block surface or may be clamped. If the joint surfaces are clamped, the tensile strength across the joint may hold the two surfaces together and form a larger block, called a combined block. The process is illustrated in Figure 4.7.

B lo c k c o rn e r

C o m b in e db lo c kfo rm e d

In i t ia l b lo c kfo r m e d

C h e c k s h e a r &s e p a r a t io n a lo n gb lo c k fa c e s

Figure 4.7: Formation of primary blocks in BCF


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Example of primary fragmentation results

An example of the block sizes generated by the primary fragmentation module is presented in Figure 4.8. The input data were for a kimberlite rock mass having a rock mass rating (RMRL) of 65 and an intact rock strength of 120 MPa. Three joint sets were present, with average spacings of between 0.56 m and 1.06 m. The maximum stress in the cave face was 27 MPa, insufficient to cause stress spalling of the rock blocks. Approximately 1000 primary blocks were generated. The results show that combined blocks with volumes of up to 30 m3 could be formed; all blocks larger than about 1 m3 may be assumed to be combined blocks. The primary fragmentation module allows the user to evaluate the effects of changing the cave face orientation and the stress magnitudes, and the sensitivity of the results to jointing input data. This improves understanding of the primary fragmentation process and allows rapid assessments to be made of mine design options.

0

5

10

15

20

25

30

35

0 100 200 300 400 500 600 700 800 900 1000Number of blocks

Blo

ck v

olum

e (m

3 )

Figure 4.8: Example of primary fragmentation block sizes calculated by BCF

4.5.3 Secondary Fragmentation

The basic principle used in the secondary fragmentation module is that blocks with high aspect ratios have a higher probability of splitting into smaller fragments than do blocks having low aspect ratios. The program considers the movement of rock blocks down the draw column in units of distance called cycles. The cycle height is defined as the vertical distance through which a block with an aspect ratio of 10:1 must be drawn in order to ensure that it will split. This distance is directly related to the rock block strength. The cycle height and the probability of a block splitting in a cycle are affected by the rock block aspect ratio, the rock strength, the


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presence of joints within a block, the cave pressure, the rate of draw and the presence of fine material between the blocks. In addition, the generation of fines through block rounding in the draw column and the splitting of rock blocks during temporary arch formation are considered. Each of these influences on secondary fragmentation will be discussed briefly below. Aspect ratio

If the aspect ratio of a block is 10 or greater it is assumed that it will split in a cycle. The probability of splitting reduces linearly to 10% when the aspect ratio is 1.0. Presence of joints

If a rock block is a combined block, it will contain joints and the probability of splitting will be much higher. The program assumes that the probability of splitting is doubled in combined blocks. Combined blocks having aspect ratios of 5.0 and higher therefore have a 100% probability of splitting in a cycle. Rate of draw

The rate of draw has an influence on the time for which a block remains in the draw column which, in turn, influences the secondary fragmentation. A faster draw rate is assumed to result in larger rock fragments. The program takes this into consideration by reducing or increasing the cycle height using an empirical relationship. Effect of cushioning

Cushioning takes place when fines prevent contact between larger blocks and act as a cushion, reducing the probability of a block splitting. Fines that exist during the primary fragmentation phase are considered to be active in cushioning. Fines that are generated during secondary fragmentation are assumed to flow out of the draw column more rapidly than the larger blocks and so do not cause any further cushioning. Caved height and cave pressure

When the user defines a draw height, the program calculates the associated height of caved material using the swell factor for the rock. The caved height will increase until it meets the ground surface or some other free surface. Once caving has reached this surface, the surface will start to subside. The height of caved material is used to determine the cave pressure. The cave pressure is calculated using the dead weight of the caved material above a rock block and considers the width to height ratio of the active draw zone. In a narrow draw zone, a large proportion of the weight of the caved material will be transferred to the surrounding rock, and the cave pressure will be reduced.


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Rounding of block corners

As a block moves down the draw column its corners will be rounded. The rounding of the corners produces material which is classified as fines. The amount of fines caused by the rounding of corners will depend on the acuteness of the corners of a block. The program makes use of the variation in joint orientation to determine the amount of fines generated. It is argued that if there is a large variation in joint orientations, acute corners will form more readily and rounding will generate more fines. The rounding of corners can generate a significant percentage of fines if the rock is drawn through a number of cycles. Arching

During the process of draw, temporary arching may take place in the draw column. An arch may be broken when a block in the arch splits under the influence of the arching stresses or when a block slides out of the arch. For splitting to take place, the stresses induced by arching must be greater than the block strength. Arching stresses are assumed to be generated by the pressure of the caved material above the block in question. Empirical relationships are used to determine the likely stress levels in the temporary arches and the numbers of blocks that will split due to arching. Example of secondary fragmentation results

The final result of a secondary fragmentation analysis is a reduction in the size of the primary rock fragments arising from the interaction of several factors. An example of a secondary fragmentation distribution is shown in Figure 4.9. The graph shows both the primary and secondary fragmentation of a rock mass after being drawn through 100 m vertically. The results show that very few fines were produced after primary fragmentation, but that after the secondary fragmentation process, about 5% of the rock may be classified as fines, that is smaller than 0.001 m3. As noted in Section 4.3, common indicators of fragmentation are often taken to be the fractions smaller or larger than 2 m3. The results in Figure 4.9 show that after the primary fragmentation about 60% of the fragments will be larger than 2 m3, and that after 100 m of draw, only 40% of the fragments will have to be broken by secondary breaking methods.


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0

20

40

60

80

100

0.001 0.010 0.100 1.000 10.000 100.000 1000.000 10000.000Block size (m³)

Secondary

Primary

Figure 4.9: Size distribution graphs for primary and secondary fragmentation calculated using BCF

4.5.4 Hangup Analysis

The hangup analysis module of BCF makes use of the results of the secondary fragmentation analysis to determine how many hangups are possible in a drawbell. The objective is to examine potential hangups at the top of the drawbell, classified as “high hangups”, and hangups lower down in the narrow part of a bell, called “low hangups”. The program assumes that if less than 25 blocks (5x5) are required to cover 40% of the area of the drawbell, a hangup is possible. The user enters the horizontal cross sectional area for each part of the drawbell. The program reads the secondary blocks from the results file and determines the approximate length and cross sectional dimensions from the block volume and aspect ratio. It then randomly selects two of the block dimensions to calculate a cross sectional area. The cross sectional areas are summed until either 40% of the bell area is exceeded or 25 blocks is exceeded. If 25 blocks is exceeded first, it is assumed that a hangup will not form. Good agreement was obtained between the results of this method and those of small physical models of drawpoints. The results of a hangup analysis may be used to estimate the production delays that will be incurred and the level of effort that will be required to bring down hangups.

Per

cent

Pas

sing


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4.5.5 Discussion

BCF is an expert (rule-based) system which provides a tool for the rapid assessment of the main factors that affect caving fragmentation. Sensitivity and comparative analyses may be carried out rapidly during the feasibility and planning stages of a project. The program is based largely on an empirical understanding of the caving process and users should be aware of the assumptions made and relationships used. BCF uses a Monte Carlo approach in which blocks are generated independently of each other using randomly chosen orientations, spacings, and trace lengths from the input joint set statistics. Estimates of the orientations and magnitudes of the stresses on some block surfaces are used to determine the possibility of shearing. The estimated rock strength and orientation of the block are also considered in this process. Stress fracturing of intact rock is considered by use of a factor of safety against failure based on the intact block strength and the estimated maximum principal stress in the cave face. These stress fractures are generated as an additional joint set. It is quite easy to envisage a situation in which the stress induced fractures in the cave back are neither sub-parallel nor of constant orientation with respect to the pre-existing discontinuities. Because of the way in which BCF combines a rock mass model and a fragmentation model into one process, the jointing statistics of the initial rock mass may not be reconstructed adequately in the model or maintained in the fragmentation process. The blocks are generated independently of the joint statistics. Even if this process produced a result that was consistent with the joint statistics of the rock mass, it is highly unlikely that the rock mass could be reconstructed without producing large amounts of void space. The method used to construct the in situ rock mass can introduce biases which may affect the predictive capabilities of the model. Consider two simple rock masses of the same overall size having identical joint shape, spacing and set orientation statistics but different joint sizes (Figure 4.10). Clearly, these two rock masses will fragment very differently following the initial stress redistribution in the caving process. The suggestion that when line mapping is conducted, only “block forming joints” should be considered is unrealistic. It is impossible to determine by inspection of a joint trace, on a censored wall, whether a given joint will or will not be a “block former”. The fact that each block is, and must be, defined in relation to all surrounding blocks in the interrelated network is implicit in the diagram and is not fully considered in the BCF model. The size of the discontinuities has a major influence on the fragmentation. As Figure 4.10 illustrates, for rock masses differing only in discontinuity size, the rock mass having the “smaller” discontinuities will be expected to produce the larger blocks on caving. The BCF methodology suggests that for larger trace lengths, there is a higher probability that larger blocks will be formed.


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Figure 4.10: Two joint models differing only in joint size

Considerations such as these suggest the need for the development of a modelling approach that has a more rigorous basis than BCF. Ideally, rock mass characterisation and modelling should precede, and form the basis of, fragmentation modelling. An approach to doing this is presented in Section 4.6. Esterhuizen (1999) made some modifications to the original version BCF to address some of the issues identified here and other limitations of the program. The modifications made to the BCF primary fragmentation module were:

• a more advanced method of natural block generation was developed allowing the faces of a block to be defined by up to six different joint sets. Previously, blocks could only be formed by three sets and it was assumed that the opposing block faces were parallel. This modification results in more realistic block formation;

• joint spacing and trace lengths are now more accurately simulated using an algorithm that

correctly simulates truncated exponential distributions. Previously, truncation of distributions resulted in mean values that were different from those defined as input;

• joint trace lengths may be used to limit block dimensions. Previously, block dimensions

were limited by the aspect ratio of a block or by shearing along the joint planes; • block accretion during primary fragmentation is now limited by the joint condition. Blocks

will accrete until a sheared joint is found, a joint with no tensile strength is found, or the end of a joint is encountered. Previously, joint condition was not considered in block accretion;

• variations in joint conditions are considered when assigning shear strengths to joints.

Previously, a single value was used;


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• stress fracture formation is inhibited when shearing occurs along two or more joint sets.

Previously, stress fractures where assumed to occur even if shearing is indicated; and • the user may specify stress fracture trace lengths. Previously, trace lengths were not

considered.

Modifications made to the BCF secondary fragmentation module include: • the block strength and aspect ratio relationship was changed to result in a better

representation of the likelihood that a block will fragment in the draw column; • the rate of draw is included as a factor affecting the degree of fragmentation; • the cave pressure is now based on the results of numerical models. Previously, cave

pressures were determined from extrapolation of sand box test results. This issue was explored in the International Caving Study Stage I as reported in Appendix B; and

• the cave pressure is updated as a block moves down a draw column. Previously, a constant

cave pressure was used.

Esterhuizen (1999) also made several changes to the BCF computer program itself:

• BCF was originally a DOS® based program. It was converted to run under the Windows® 32 bit operating system. This required re-coding of the entire user interface and has produced many small improvements that reduce the learning curve when starting to use the program;

• it is now possible to set up batch runs that allow multiple data sets to be evaluated

simultaneously; and • graphic viewing of results has been improved. Comparative viewing of several sets of

output data is now possible. 4.6 AN ALTERNATIVE METHOD OF ASSESSING IN SITU AND PRIMARY FRAGMENTATION

4.6.1 Methodology

As has been indicated, a new method of identifying in situ blocks and of predicting primary fragmentation for block caving applications was developed by Eadie (2002) as part of the International Caving Study Stage I. The model known as JKFrag attempts to address the


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problem fundamentally, using and further developing the discontinuity modelling approach presented in Section 2.6. It acknowledges the interdependence of the discontinuities in the network and uses a tessellation method to define the primary blocks. The basic assumption made is that the in situ blocks are defined by the network of discontinuities pre-existing in the rock mass. Each discontinuity is assumed to be a potential block face and discontinuity intersections are assumed to be potential edges of blocks. For a block to be formed there must exist a collection of discontinuities that intersect appropriately to define a fully enclosed region. As a result, each potential block may also contain discontinuities. It is therefore important to consider the complete network of discontinuities when identifying in situ blocks. As was discussed in Chapter 2, the network of discontinuities is measured and described statistically in terms of spacing, orientation, persistence and termination. In order to assess the primary fragmentation of a rock mass having given statistical properties, two processes are used: • the simulation of a rock mass model with properties statistically equivalent to those

obtained from field data; and • the determination of the in situ and primary fragmentation expected from that rock mass

simulation. This approach ensures that the given geometrical properties of the rock mass are maintained while providing a methodology that recognises the interdependence of blocks. The fragmentation model is independent of the method used to simulate the rock mass model. The algorithms used to determine the distribution of blocks expected from a given rock mass model are outlined below. A flow chart setting out the steps used in the approach is shown in Figure 4.11. The method applied to define blocks considers each of the discontinuities in relation to its neighbours. The in situ and consequently the primary blocks, naturally form a representation of the rock mass having polyhedral shaped building blocks. A mathematical method known as a constrained tessellation is used to divide the region using the simulated discontinuities as the primary inputs. This is achieved by initially constructing a Delaunay triangular grid of the region of interest.


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Figure 4.11: Simplified flow chart showing the steps in JKFrag (Eadie 2002) 4.6.2 Tessellation Procedure

The following procedure is used to efficiently construct the required triangular grid. Consider the given cross section of a jointed rock mass shown in Figure 4.12a. The end points of each of the joint traces within or dissecting that section are shown in Figure 4.12b. Note that in the case where traces either intersect or terminate on each other, the point of intersection is also considered to be an end point. These end points form the basis for an initial tessellation.


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(a) (b)

Figure 4.12: (a) Cross section of a jointed rock mass, and (b) end points of traces

0 2 4 6 8 10 12 140

2

4

6

8

10

12

Figure 4.13: Delaunay triangulation of the set of end points of the joint traces

A grid is generated using these end points such that the convex hull is completely covered by non-overlapping triangles. The resulting triangulation is shown in Figure 4.13. The Delaunay Triangulation has been used for the initial tessellation, because algorithms exist to accomplish the task efficiently. The algorithm implemented for the initial triangulation follows the method described by Tsai (1993).


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The traces in the given cross section may coincide with some of the edges that have been constructed in the tessellation. Traces that do not appear as edges in the tessellation need to be forced or constrained to become edges. To illustrate this point, consider the trace (or constraint) indicated in Figure 4.14a. It traverses several triangles. These particular triangles form what is termed the influence polygon of the constraint. Removing the triangles that comprise the influence polygon will produce two open regions, one on each side of the constraint.

(a) (b)

Figure 4.14: (a) A constraint with its associated influence polygon shaded, and (b) the

updated constrained triangulation

The process of adding the constraint is completed once each region has been triangulated as shown in Figure 4.14b. In the implementation of this algorithm a record is kept to indicate which triangle sides in the grid are constraints. Each of the remaining traces that do not coincide with edges in the initial triangulation are then added incrementally until the constrained tessellation shown in Figure 4.15 has been achieved. This construction using the traces as foundational edges of a grid has • utilised, without altering, the jointing statistics of the rock mass model; • provided information about each discontinuity and its relationship to its neighbours; and • provided a means of accessing each of the closed and partially closed spaces contained

within the bounds of the discontinuities for further fragmentation modelling. This approach is considered to provide a geometrically rigorous foundation for the subsequent estimation of in situ and primary fragmentation, particularly as it considers all of the simulated joints.


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2 4 6 8 10 12 14

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Figure 4.15: Completed constrained triangulation of convex hull of joint traces

4.6.3 In Situ Blocks

As shown in Figure 4.11, the next step in the process is to develop an in situ block size estimate. Once the constrained tessellation has been generated, the regions fully enclosed by eligible joint traces are then identified as cross sections of the in situ blocks. Joint traces that from the rock mass characterisation data are deemed to be open, are considered eligible. The regions of interest are determined by identifying the sets of adjoining triangles that are bounded by open discontinuities. An enclosed region can be identified by initially considering one of the triangles. If any of the sides of this triangle are listed as not being a constraint (eligible discontinuity), then any adjacent triangle sharing the common side must also be part of the enclosed region. Once all the triangles in the tessellation have been considered, all the enclosed regions have been identified. A record is kept of the triangles comprising each region. The area of each triangle is calculated using the coordinates of its end points. This provides the intersected cross sectional area of each of the established blocks in this plane. From this information a cumulative distribution of the cross sectional areas can be constructed. The two-dimensional information in this format has only limited practical application. However, in theory it provides a basis for the estimation of the volume distribution of the in situ blocks. As more sections at various orientations through the rock mass are examined, more information becomes available to confirm the three-dimensional properties of the rock mass.


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The nature of the discontinuity data available determines the accuracy of the predictions. Besides the three-dimensional nature of discontinuity orientation, core logging and face mapping can provide only one- or two-dimensional data. In particular, there is no current agreement on the shapes of discontinuities. Furthermore, as was discussed in Chapter 2, there is no full understanding of discontinuity termination in three dimensions. This also applies to the mode of joint extension in three dimensions. As volumetric information is required for practical purposes, some form of dimensional extrapolation must be made. Consider the case in which each slice taken through the rock mass displays the same intersected cross sectional area distribution. This rock mass may be defined, for the purposes of this discussion, as an isotropic rock mass. For each of the cross sectional areas of the block section in the plane, the volume is estimated as the volume of a sphere having as its radius the radius of a circle which has the same area as the block cross sectional area being considered. By applying this approach for various and appropriately chosen cross sections of the simulated rock mass, a series of distribution curves is obtained. Each curve represents an isotropic (as defined above) rock mass. The isotropic rock mass extrapolated from the cross section giving the finest block size distribution produces a lower bound curve for the anisotropic rock mass in question. Similarly, an upper bound curve can be found from the cross section giving the coarsest block size distribution. The two-dimensional fragmentation output provides both the information required for the stereological extrapolation to a three-dimensional volume distribution as well as a means of estimating upper and lower bounds for the fragmentation curve. Knowledge of the overall geometry of the rock mass then provides guidance for choosing the distribution within the predicted range. Even though the concept of spheres is used in this extrapolation, information regarding shape is not lost. The aspect ratio of each of the blocks in the plane can be calculated for each section taken. This provides information that may be used to identify the expected block shapes.

4.6.4 Primary Fragmentation

The primary fragmentation module makes an assessment of the blocks formed in the cave back before movement has occurred. It is assumed that the stresses induced in the cave back act on the in situ rock mass to produce further fragmentation. New fracturing of previously intact rock could be produced under high induced stress conditions as assumed in BCF. However, it is considered more likely that, in the general case, further fracturing contributing to the primary fragmentation will occur predominantly by: • the extension of pre-existing weakness planes; and • the fracturing of intact rock bridges.


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The tessellation presented naturally provides a means of simulating these two mechanisms. Recall the cross section illustrating the joint traces and the underlying triangular grid shown in Figure 4.12. Two situations arise in the tessellation that provide information for the model: • any joint trace that terminates in intact rock; and • intact rock bridges between joint ends.

The subsequent fragmentation will be influenced by • the magnitudes and orientations of the in situ stresses; • the strengths of the rock blocks; and • the geometry of the excavation. Eadie (2002) initially used a relatively simple index to account for these factors - the ratio of the maximum induced tangential stress, σθ, to the uniaxial compressive strength of the intact rock, σc (Mathews et al 1980, Grimstad and Barton 1993). This index is used as a basis for estimating the percentage extension of the unintersected trace length of each unterminated joint end point in the model. Grimstad and Barton’s guidelines on stress effects for a range of stress-strength ratios are used as the basis of a preliminary estimate of the extension of joints (see Table 4.2). Figure 4.16 shows a completed tessellation for the example previously considered in Figures 4.13, 4.14 and 4.15 following the extension of joints using the preliminary rules suggested in Table 4.2. The magnitudes of the joint extensions given in Table 4.2 require further refinement. This is likely to be best achieved by the numerical analysis of a range of plausible cases as illustrated by Eadie (2002).

Table 4.2: Estimated percentage trace length extensions for the stress – strength ratios defined by Grimstad and Barton (1993)

σθ/σc STRESS LEVEL AND EFFECT Joint

extension %

<0.01 Low stress, near surface, open joints 0

0.01-0.3 Medium stress, favourable stress conditions 0

0.3-0.4 High stress, very tight structure. Usually favourable to stability,

maybe unfavourable to wall stability.

0-10

0.5-0.65 Moderate slabbing after > 1 hour in massive rock. 30-50

0.65-1.0 Slabbing and rockburst after minutes in massive rock 50-75

>1.0 Heavy rockburst (strain-burst) and immediate dynamic

deformations in massive rock

75-100


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2 4 6 8 10 12 14

2

3

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5

6

7

8

9

10

11

12

Figure 4.16: Completed tessellation after the extension of joints

In the implementation of the primary fragmentation module, the block identification algorithm outlined for the in situ model is used to identify the regions fully enclosed by eligible joint traces. Here again, only eligible joints are considered to form possible block boundaries. The eligibility criterion is that the joints have to be open. A joint that has been extended is considered open. This block information is then analysed and presented as block volumes using the extrapolation procedure discussed for the in situ blocks.

The block identification algorithm also records information on the extent of internal jointing. This approach inherently respects the geometric characteristics of the set of all discontinuities. It naturally provides the ability to extend pre-existing weakness planes and fracture intact rock. The method also provides the ability to consider the extent of intact rock bridges and to make primary fragmentation estimates for any applied stress level or regime. For obvious practical reasons, field data can only be obtained to indicate the combined outcomes of primary and secondary fragmentation processes. Accordingly, the contribution made to the overall fragmentation in the primary stage cannot be isolated and measured in normal caving operations. Despite the difficulty of obtaining validation data, it is considered that the concepts and approach presented above provide a sound and rational basis for further development. For example, as knowledge of crack propagation mechanisms and modelling


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capability increase, it may become possible to develop a more deterministic model of joint extension. 4.7 CONCLUSIONS

In this chapter, the factors influencing fragmentation, methods of fragmentation measurement and methods of predicting fragmentation (most notably the BCF program) have been reviewed and a new method proposed for identifying in situ blocks and predicting primary fragmentation. The review of methods of fragmentation measurement concluded that digital imaging processing systems have potential for further development for use in cave mining applications if lighting and dust conditions can be controlled.

The BCF program is the only complete or near-complete method available for predicting the fragmentation produced in caving operations. Some improvements to BCF were made as part of the International Caving Study Stage I. It is likely that some of the assumptions and empirical rules used in the program could be further improved. For example, the program does not adequately represent and treat the geometrical properties of the full discontinuity network pre-existing within the caving mass. Accordingly, a new and more fundamentally based method of identifying in situ blocks and of predicting the primary fragmentation associated with the onset of caving has been developed by Eadie (2002). This method which is known as JKFrag is based on a rigorous two-dimensional tessellation procedure. It is considered to provide a reliable starting point for the development of a new or improved method of predicting caving fragmentation. The next stage of research on this problem will be to develop a method of predicting the secondary fragmentation produced within the draw column. As outlined in Section 4.2, secondary fragmentation may be produced by a variety of mechanisms, not all of which are well understood. The BCF approach to predicting secondary fragmentation uses some assumptions and rules that are considered unlikely to be universally applicable. It is considered unlikely that numerical modelling approaches will, of themselves, be able to fully resolve the problem in a practical way, although parametric numerical studies are likely to be very useful in developing improved empirical rules for primary and secondary fragmentation assessment. Therefore, the approach suggested for the development of an improved method of predicting secondary fragmentation, is to initially study the basic mechanics of the each of the major mechanisms involved using analytical and numerical models. From these studies and parametric numerical stress analyses, a set of empirical rules for each the mechanisms will be developed. In some cases, the rules used in BCF may provide an appropriate starting point, or indeed, the only practicable option. It is considered that the development of the empirical rules will also benefit from the results of distinct element numerical simulation studies of the type discussed in Appendix C.

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CHAPTER 5

CAVE INITIATION BY UNDERCUTTING 5.1 INTRODUCTION

s was indicated in Chapters 1 and 3, the caving of a block or panel is initiated by mining an undercut until its hydraulic radius reaches or exceeds a critical value. As broken ore is removed progressively from the critical undercut area, the ore above it

will collapse into the void so created. Vertical propagation of the cave will then occur in response to the continued removal of broken ore through the active drawpoints. Horizontal propagation of the cave will occur as more drawpoints are brought into operation under the undercut area. Experience has shown that undercutting makes a critical contribution to the success or otherwise of block and panel caving (eg Laubscher 2000). Poor planning, design, implementation and management of the undercut can jeopardise the ultimate success, productivity and costs of an operation. In particular, care must be taken to ensure that the undercut does not impose excessive abutment stresses on the surrounding rock mass and extraction level excavations which could cause delays in production and incur excessive costs through support and reinforcement requirements and rehabilitation. In summarising the nature and importance of undercutting, Butcher (2000a) has suggested that it has three aims: • to extract a void of sufficient dimensions to allow caving to occur; • to achieve the required undercut dimension to initiate caving with minimum damage to the

surrounding rock mass; and • to advance (in time as rapidly as possible) to caving hydraulic radius, initiate caving,

propagate the cave and consequently reduce undercut abutment stress. The successful implementation of the apparently simple undercutting concept in the range of circumstances met in practice requires that careful attention be paid to several factors. The examples of the background to the development of current caving practice at a number of mines given in Chapter 1 illustrate the effects of some of these factors which include: • the sequence of undercut and extraction level development; • the relative positions of, and distances between, the undercutting front, the front of

extraction level development and the extraction front;

A

Chapter 5: Cave Initiation by Undercutting

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• the starting point and direction of undercut advance; • the rate of undercut advance; • the height of the undercut; and • the shape of the undercut in both plan and vertical section. In this chapter, accounts will be given of • undercutting strategies; • undercut design and management including the influence of the factors listed above; • undercut shape and extraction methods; • the influence of undercut strategies on the stresses induced in the undercut and extraction

levels; and • drilling and blasting practices for the mining of the undercut. It is important to recognise that, although undercutting is the essential means of initiating caving, additional cave inducement measures may be required in order to initiate and sustain caving, especially in the higher strength rock masses for which caving methods of mining are now being used. Data collected by Flores and Karzulovic (2002b) indicate that drill and blasted slots were used to create release surfaces and assist cave initiation and propagation in more than 50% of the current caving mines studied. There is also interest in pre-conditioning rock masses by blasting or hydraulic fracturing, for example, to improve their caveability. This is one of the major topics being studied in the International Caving Study Stage II. 5.2 UNDERCUTTING STRATEGIES

5.2.1 Purpose

As the examples of current practice given in Chapter 1 and the numerical analyses to be presented in Section 5.5 show, the undercutting strategy adopted can have a significant effect on the stresses induced in, and the performance of, the extraction level installations and on cave propagation. Basically, three different undercutting strategies may be used – post-, pre- and advance undercutting. To these may be added other variants such as the Henderson method. The following accounts of these undercutting strategies and of their advantages and disadvantages are based on those of Bartlett (1998) and Butcher (2000a). 5.2 2 Post-undercutting

The post-undercutting strategy is also referred to as conventional undercutting. As illustrated in Figure 1.12, undercut drilling and blasting takes place after development of the underlying extraction level has been completed. Cones, drawbells or troughs are prepared ahead of the undercut and are ready to receive the ore blasted from the undercut level.


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The advantages of this system are that blocks can be brought into production more quickly than with some other methods, no separate ore handling facility is required on the undercut level, drifts on the undercut level are required only for drilling and blasting and so can be 30 m apart, and the probability of ore compacting on the undercut level is very small. The main disadvantages are that, other than in low stress environments, the rock mass between the undercut and extraction levels is subjected to high and variable stress levels and support and reinforcement must be installed well ahead of the undercut stress abutment zone. This can constrain the rate of undercut advance. Butcher (2000a) suggests that, as a general guideline, the use of a post-undercutting strategy should be assessed critically when the depth of the cave is greater than 500 m, when the caving area has a hydraulic radius of greater than 17 m, and when draw horizon extraction exceeds 50%. An example of the use of a post-undercutting sequence in the BA5 block of the Premier Diamond Mine is given in Section 1.3.3. 5.2.3 Pre-undercutting

In this approach, the undercut is mined ahead of extraction level development. In some instances, the term pre-undercutting is used to describe the case in which the undercut is completed before any extraction level development is carried out. On the other hand, pre-undercutting may also be considered to be a variant of the advance undercutting method to be discussed in Section 5.2.4 with the development of the extraction level lagging some distance behind the undercut. The minimum horizontal distance that the extraction level development lags behind the advancing undercut is often the separation distance between the two levels. This is sometimes referred to as “the 45 degree rule”. However, in higher stress environments, it may prove necessary to use larger lag distances than those given by this rule. Even with the 45 degree rule, it may be possible that the extraction level excavations will not be in a full stress shadow zone or may “see” some abutment stress concentration. When stresses are high, this may be sufficient to cause distress to the extraction level installations. In the Esmeralda sector of El Teniente, for example, it was found that difficulties arose when the 45 degree rule was applied and that a larger lag was required. In this case, the vertical separation between the extraction and undercut levels is 12 to 15 m but it has been found most satisfactory to maintain a horizontal separation of 22.5 m between the undercut front and extraction level completion (Jofre et al 2000). The production zone is located some 45 to 60 m behind the undercut front. Advantages of pre-undercutting are that the extraction level is developed in a de-stressed environment, the undercut can be mined independently of the extraction level, support requirements on the extraction level are generally lower than in the post-undercutting method, and the broken ore in the undercut level acts as rock fill reducing the abutment loads on the undercut face to some extent.


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The disadvantages of this strategy include the need for a separate ore handling facility on the undercut level, possible sequencing problems between extraction horizon development and undercutting, the need to develop drawbells from the extraction level into broken rock on the undercut level, possible high stress remnants arising from the compaction of blasted undercut ore making extraction horizon development difficult, drawpoint hangups resulting from ore compaction, and slower initial production arising from these various factors. An example of the use of pre-undercutting at El Teniente is given in Section 1.3.2 and illustrated in Figure 1.12. 5.2.4 Advance Undercutting

In the advance (sometimes called advanced) undercutting strategy, undercut drilling and blasting takes place above a partially developed extraction level. The partial development on the extraction level can consist of either extraction drifts only or extraction drifts and drawpoint drifts. Drawbells are always prepared in the de-stressed zone behind the undercut, usually adhering to the 45 degree rule. Figure 5.1 illustrates a conceptual advance undercut strategy proposed for the panel cave mining of future sectors of El Teniente, Chile (Jofre et al 2000). Butcher (2000b) notes that advance undercutting is essentially a compromise between the post- and pre-undercutting strategies, in that: • draw horizon damage is reduced because, compared with the post-undercutting strategy,

the extraction ratio on this horizon is decreased; • the cave is brought into production more quickly than with the pre-undercut strategy,

reducing the problems associated with increased development times; • the probability of the formation of stress-inducing remnants arising from muck pile

compaction is reduced; • a separate level is still required for undercutting but it will require a much more limited ore

handling facility than in the post-undercutting strategy; and advance undercutting is slower than post-undercutting because of the remaining extraction level development that is required after the undercut has advanced. The fact that drawbell development must be accomplished from the extraction level into broken ore in the undercut level contributes to this. On the other hand, this method obviates the need for the time-consuming and costly repairs to the extraction level drifts that are almost inevitably required with the post-undercutting strategy.


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Because of its inherent advantages, the current trend in block and panel cave design is to use an advance undercutting strategy with the disadvantages that have been outlined being reduced to tolerable levels by careful planning, support and reinforcement design and equipment selection.

Figure 5.1: Advance undercut panel caving, El Teniente mine, Chile (after Jofre et al

2000) 5.2.5 The Henderson Strategy

To the three main undercutting strategies must be added the Henderson or "just in time" method described in Section 1.3.3 and illustrated in Figures 1.17 and 5.2. In this approach, the drawbells are blasted with long holes from the undercut level just ahead of the blasting of the undercut itself (see Figure 5.2). This reduces the time during which the pillars and the extraction level installations are subject to high abutment stresses and damage may occur.

Undercut Sequence – Advanced U/C Panel Caving


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Laubscher (2000) suggests that this strategy does not work satisfactorily in squeezing ground conditions where hole closure occurs.

Figure 5.2: Drill layout for undercut and drawbell blast, Henderson Mine, USA (after Rech et al 2000)

5.3 UNDERCUT DESIGN AND MANAGEMENT

5.3.1 Purpose

As was noted in Section 5.1, a number of factors associated with the design and implementation or management of the undercut can have important impacts on undercut and overall caving performance. The purpose of the present Section is to discuss some of these issues. Among the most important of them are the method of formation and the cross-sectional shape of the undercut (flat, narrow or inclined). This factor will be discussed separately in Section 5.4. 5.3.2 Initiation and Direction of Undercut Advance

The choice of the starting or initiation point for the undercut and the preferred direction of undercut advance can be influenced by several factors including • the shape of the orebody; • the distribution of grades within the orebody;


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• the in situ stress directions and magnitudes; • the strength of the orebody and its spatial variation; • the presence and orientations of major structural features in the orebody; and • the presence of caved areas adjacent to the block or panel to be undercut. If the orebody is long and narrow in plan, a constraint will be placed on the possible directions of undercut advance. Under these circumstances, it will generally be necessary to open the undercut to the full width of the orebody and advance it in the longitudinal direction. There may be advantages in terms of productivity in retreating the cave in two directions away from a central slot or starting point as in the front caving example outlined in Section 1.2.1. However, it is more common for orebodies that may be mined by block or panel caving methods not to be of this elongated shape and to be either more approximately equi-dimensional or large in each plan dimension. In this more general case, the other factors listed above must be considered. In the case of an approximately equi-dimensional orebody, it is common for the cave to be initiated against a slot on the boundary of the orebody and advanced diagonally across the orebody as in Northparkes E26 block cave Lift 2 (Duffield 2000). Panels are usually similarly advanced on a diagonal front across the orebody as illustrated in Figure 1.3. Alternatively, the initiation point may be near the centre of the orebody with the undercut being developed progressively outwards towards the orebody boundaries as shown in Figure 5.3 for the Palabora block cave (Calder et al 2000). Operational factors as well as the distribution of grades and of the strength and caveability of the orebody need to be considered in establishing the starting point for undercutting in such cases. If, as in the case illustrated in Figure 5.3, the orebody is elongated in one direction, the issue of the minimum dimension required to achieve self-propagating caving referred to in Section 3.2 may arise.

Figure 5.3: Planned undercut sequence, Palabora block cave, South Africa

(Calder et al 2000)


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The influence on cave initiation of the in situ stresses and their redistribution around the undercut and the developing cave was outlined in Section 1.2.2. For the purposes of this discussion it will be assumed that the undercut drifts and the direction of cave advance are aligned with the principal horizontal in situ stresses as illustrated in Figure 5.4. If the direction of advance is perpendicular to the direction of the major principal horizontal stress, the levels of stress in the abutment ahead of the undercut will be high and will increase as the undercut advances. As the analyses and discussion to be presented in Section 5.5 will show, this will increase the likelihood of damage to the undercut drifts and the extraction level excavations. However, this effect may be an advantage in overcoming the strength and inducing caving of stronger rock masses as in CODELCO-Chile’s Andina and El Teniente mines.

Direction of cave advance

σh1

σv σh2

Figure 5.4: Direction of in situ stress relative to cave advance

Any spatial variation in the strength of the orebody can be expected to have an effect on the influence of the induced stresses on cave initiation and propagation. Because caving should be easier to initiate in weaker than in stronger ore, and because the stresses induced ahead of the undercut should increase as the undercut advances, it is often argued that mining should take place from a weaker to a stronger section of the orebody (Ferguson 1979). Other things being equal, there may be some advantages in terms of rates of return to offset capital costs of starting mining in any higher grade zones of the orebody. This may, however, have undesirable effects on the commissioning and operation of the processing plant. Major structural features such as faults and shear zones can have an influence on cave initiation and propagation and on the stability of undercut and extraction level excavations (eg Ferguson 1979, Laubscher 2000). A major circumstance to be avoided is the isolation of large wedges of rock that may fall or “sit down” under the influence of gravity, inhibiting cave propagation and imposing additional dead weight loads on undercut drifts and extraction level excavations. As a general rule, it is preferable to orient the advancing undercut face as close as possible to normal to the strike of any persistent structural feature or set of features.


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In many large caving mines, the orebody may be extracted in a series of blocks. Where possible in these cases, new blocks should be retreated from existing blocks rather than being advanced towards them. As illustrated in Figure 5.5, this prevents the creation of potentially highly stressed pillars between the two caves which can lead to the stress-induced failure of excavations in their vicinity. Ferguson (1979) gives an example of difficulties arising from this cause at the Shabanie mine, Zimbabwe.

(a) (b)

Figure 5.5: The initiation of caving of a block adjacent to an existing caved block for

(a) the preferred mining direction, and (b) an unfavourable mining direction (after Ferguson 1979)

5.3.3 Shape of the Undercut Face

Both mining experience and a consideration of the induced stresses suggest that sharp changes to, or large irregularities in, the shape of the advancing undercut face should be avoided and that the lead or lag between adjacent sections of the overall face should be minimised. Butcher (1999) has suggested that horizontal undercut lags should be generally less than 8 m to avoid significant undercut drift damage. A circular or square undercut will produce a larger hydraulic radius than a rectangular undercut of the same plan area and so should induce caving more readily. However, a flat undercut face is difficult to achieve in practice and sharp corners, especially re-entrant corners, are to be avoided. These factors argue for the adoption of a curved face with a large radius of curvature (Ferguson 1979). A face that is convex with respect to the cave should cave more readily than one that is concave. In panel caving operations, the cave front should advance across the orebody in a straight line as each of the adjacent panels is undercut as illustrated for the case of the Henderson mine in Figure 1.3.

Figure 5.6 illustrates some of these desirable and undesirable features of undercut face shape and orientation in an idealised case. It must be remembered that, in practice, it is not always possible to avoid the undesirable and adopt the desirable features. The orebody boundaries may


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not be of the conveniently regular shape shown in Figure 5.6, and changes of lithology such as the presence of more massive and stiffer rock (which could act as a stress concentrator and increase rock burst potential) can be expected to influence the preferred undercut shape. Furthermore, the preferred undercut orientation with respect to in situ stresses may have to be modified if major discontinuities leading to rock mass instability are present.

Figure 5.6: Idealised plan illustrating some of the desirable and undesirable

features of undercut shape and orientation

5.3.4 Rate of Undercut Advance

It is not an easy matter to establish the optimum rate of undercutting in a given case. To do so usually requires that a compromise be reached between a number of competing factors:

• reaching often ambitious initial production targets in order to achieve early financial returns can result in pressure to increase rates of undercutting;

• in the advance and post-undercutting methods, the rate of undercutting cannot exceed the

rate at which drawbells and drawpoints can be formed; • experience of the type reported in Section 1.3 shows that in high stress environments, high

rates of undercut advance can increase the levels of damage to undercut drifts, pillars and extraction level excavations and can lead to rock bursting in some cases. Reducing the rate of undercutting in these cases generally reduces the extent of damage and the incidence of rock bursts (Rojas et al 2000a);


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• the rate of advance must not exceed that which can be accommodated by the caving rate of

the orebody. At the present time, the rate of caving can be determined only from experience. It is a topic requiring further field and theoretical research. It is clear, however, that the caving rate is a function of the stresses induced in the cave back and of their relation to the strength of the rock mass. As a general rule, to ensure that the cave remains full of caved ore, the rate of drawing ore should not exceed rate at which bulking (the excess of caved volume over in situ volume) is produced by the “natural” caving process. If the rate of draw is too large, an air gap can develop between the back of the cave and the ore pile in the cave. Sudden or massive failure of the cave back can then lead to potentially disastrous air blasts. This topic will be discussed in Chapter 10;

• after caving has been induced, the rate of undercutting will be influenced by the height of

the ore column and the extraction level layout which, in turn, influence the rate of production;

• in some orebodies, usually those that are weaker and fragment more finely, excessively

slow rates of undercutting and of removing the blasted and caved ore can lead to compaction of the ore and difficulties in achieving uniform flow and extraction of the broken ore; and

• irrespective of the influence of these various factors, uniform temporal and spatial rates of

undercut advance should be achieved for the best results. Data collected by Flores and Karzulovic (2002b) show that undercutting rates in current block and panel caving mines may vary from 500 to 5000 m2 per month with the mean being in the range 2000 to 2500 m2 per month. Table 5.1 shows examples of rates of undercutting that have been used successfully in a number of recent cases.

Table 5.1: Examples of undercutting rates

Mine Undercutting rate (m2 per month)

Kimberley Mines* 2700

De Beers Premier mine BA5 panel cave 900

De Beers Premier mine BB1E block cave 1100

Esmeralda panel cave** 3000

Northparkes E26 block cave 1600 * The Kimberley rates are influenced by the need to minimize damage to extraction level pillars.

** The Esmeralda rates are influenced by the need to control seismicity


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5.3.5 Undercut Height

In the past, it has been variously thought that a high undercut would serve to limit the stresses induced in the extraction level excavations (Ferguson 1979) or to ensure that the cave propagated and the ore flowed as intended (Laubscher 1994). Furthermore, high undercuts have been seen as a source of early ore production as in the case of the Northparkes E26 mine, Lift 1 illustrated in Figure 5.7 where two undercut sublevels were used with a total height of more than 40 m (Dawson 1995). One of the assumptions has been that in the stronger orebodies in which drawbells are widely spaced to permit the use of large equipment, coarsely fragmented material would “sit” on the major apex between drawbells and would not gravitate into the drawbells so that uneven flow and draw of the broken ore would result.

Upper undercut

Lower undercutDrilled

18m

12m

Extraction level

30m

28m 4.2x4.2m

Figure 5.7: Northparkes E26 Lift 1 high undercut geometry (Vink 1995)

From a drilling and blasting perspective, higher undercuts are easier to break because of the increased free-face area available. The risk of not achieving full breakage is greater in narrow undercuts because of the higher confinement. In addition, any small amount of hole deviation can exacerbate the confinement problem. Experience with high undercuts, particularly in the stronger orebodies, has led to the conclusion that narrow undercuts do not have the previously assumed disadvantages (Laubscher 2000). As a consequence, high undercuts have been progressively replaced at a number of operations in recent years. Some disadvantages of high undercuts have been found to be:


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• the long holes necessarily used to mine the undercut can produce irregular cave backs resulting in poor caving and fragmentation characteristics being experienced;

• for the reasons referred to earlier, high undercuts lead to increased initial and overall costs; • the slow rates of advance achieved for high undercuts may produce some of the

disadvantages outlined in Section 5.3.4; and • mucking of swell from a high undercut can present operational problems particularly if an

advance undercut is used. Despite the advantages now being seen for narrower undercuts, several practical considerations help define a minimum undercut height in any given case. As well as the practicalities of drilling and blasting narrow undercuts to be discussed in Section 5.4 below, the degree of primary fragmentation achieved on the initiation caving is a most important factor. If the undercut is insufficiently high with respect to the block size produced in coarsely fragmenting ore, the likelihood of the formation of "pillars" of caved ore in the undercut is increased. Jofre et al (2000) describe the evolution of undercut heights and the drill layouts used in the mechanised panel caves at El Teniente. Ferguson (1979) discusses the earlier replacement of a double undercut at the King Mine, Zimbabwe, by a single narrow undercut. Despite the general recognition of the possible value of a narrow undercut, there are circumstances in which a high undercut may be beneficial. Rech et al (2000), for example, discuss the planned layout for Henderson Mines eastern section where the ore column height will be increased to 244 m. Together with this increased column height, a wider draw point spacing and a higher undercut than those presently used will allow the productivity per draw point to be increased about three-fold. Table 5.2 shows the undercut heights used at some current and recent block and panel caving operations, most notably in the various sections of the El Teniente mine. The highest undercut listed in Table 5.2 is the two-stage undercut used for Lift 1 of the Northparkes E26 mine described by Vink (1995). The undercut geometry and blasting patterns used in this case are illustrated in Figure 5.7.


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Table 5.2: Examples of undercut heights

Mine & Sector Undercut

Height (m)

Teniente (Ten 4 Sur B) 16.6

Teniente (Ten 4 Sur C) 13.6

Teniente (Ten 4 Sur D) 13.6

Teniente (Ten 4 Sur D) 10.6

Teniente (Ten 4 Sur Fw D) 3.6

Teniente (Esmeralda) 3.6

Teniente (Sub 6 Experiment) 16.6

Teniente (Tte. 5 Pilares) 7.0

Teniente (I-13 Tte 3) 8.6

Teniente (I-14 Tte 3) 8.6

Teniente (HP Tte 3) 4.0

Northparkes E26 Lift 1 42

Bell Canada 6

Palabora (inclined) 4

5.4 UNDERCUT SHAPE AND EXTRACTION METHOD

5.4.1 Introduction

In addition to the factors discussed in the previous section, the shape of the undercut (in vertical section) will have a major influence on the ease and effectiveness of its formation and on its effectiveness in initiating caving. There is an especially important relationship between the design shape of an undercut and the drilling and blasting practices used for its formation. Drilling and blasting for undercutting will be discussed in Section 5.6. Traditionally, undercuts have been designed with flat or undulating roofs and flat floors broken significantly by the drawbells. The tops of the pillars or major apices left between the drawbells were then rectangular and flat. To obtain better ore flow, at least one pair of sides of the drawbells are inclined downwards, reducing the extent of the floor left at the undercut level. The major apices may also be shaped to assist ore flow and to prevent ore from stacking on the tops of the apices which may inhibit effective blasting, the flow of ore and cave propagation. Two basic methods of forming traditional “flat” backed undercuts will be discussed here, the fan and the flat undercut.


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An alternative to shaping only the major apex is to incline both the roof and floor to form a narrow chevron shaped undercut. This type of undercut is a relatively recent innovation advocated by Laubscher (2000). To ensure adequate flow of the broken ore to clean the floor of the undercut, the angle of undercut inclination should be greater than the angle of friction generated between the broken ore and the in situ rock. Some advantages and disadvantages of this method and some examples of designs of this type are given in Section 5.4.4. The accounts of the fan, flat and narrow inclined undercutting methods given below are based on those of Butcher (2000a). 5.4.2 Fan Undercut

Fans have probably been the most common form of drill pattern used for undercutting in the history of block and panel cave mining. They have been used with grizzly, slusher drift and LHD caves, and have been implemented with and without the development of separate undercut levels. They are relatively flexible in that they can be drilled with a range of drilling equipment, their heights can be increased to produce additional undercut tonnage and they can be adapted easily if a mine changes its mining method from open stoping or sublevel caving to block caving. The backs of the undercuts produced by fan drilling may be undulating rather than flat. This may reduce the confinement in the areas of the back that are convex downwards and help induce caving. On the other hand, there is the potential for large, relatively unbroken masses of rock to be released from these areas causing stacking or cleaning problems. The major problems with fan undercuts arise from the formation of pillars due to blast hole loss and choke blasting conditions arising from inadequate cleaning of the undercut or the stacking of blasted ore on the major apex. These problems can be overcome with good design and blasting practice as discussed in Section 5.6. A traditional fan undercut producing an undulating back at the Bell mine, Canada, is shown in Figure 5.8. Fans are also shown in Figure 5.7 illustrating the two level undercut used at the Northparkes E26 mine, Lift 1.


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Figure 5.8: Fan undercut used with a grizzly system at the Bell mine, Canada

(Lacasse and Legast 1981)

5.4.3 Flat Undercut

As the term implies, a flat undercut is formed using flat lying drill holes rather than fans or steeply inclined holes. As a result, the undercut is narrow with a height not much greater than that of the drill drifts. A number of drilling patterns may be used in forming flat undercuts. Figures 5.9 and 5.10 show two different patterns used on the Esmeralda section of El Teniente. In the original half pillar or “John Wayne” method shown in Figure 5.9, holes were drilled from two adjacent drill drifts to meet in the middle. In the “full drilled pillar” design shown in Figure 5.9, the flat holes are drilled from one drill drift through to the adjacent drift. In either case, the best results are obtained if the holes are inclined rather than normal to the drift axes. In cases such as that illustrated in Figure 5.9, there is a danger that remnant pillars may result from poor blast hole toe breakage. This problem can be overcome by overlapping the ends of the holes in a chevron pattern. Butcher (2000a) suggests that flat or narrow undercuts are often used at deep levels because • they produce higher advance rates because less drilling and charging are required; • undercut blast hole loss is less because of the fact that fewer holes are required; and • lower undercut heights reduce the magnitudes of the induced stresses which may otherwise

cause problems.


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Undercut Drilled

15m

Extraction Level

4 x4 m

16m

DriftDrift

Drift Drift

Undercut Drilled

15m

Extraction Level

4 x4 m

16m

DriftDrift

Drift Drift

Figure 5.9: The original half pillar narrow flat undercut (“John Wayne”) at Esmeralda


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Undercut Drilled

15m

4 x4 m

Extraction Level

16m

HW FW

Drift Drift Drift

DriftDriftDrift

Undercut Drilled

15m

4 x4 m

Extraction Level

16m

HW FW

Drift Drift Drift

DriftDriftDrift

Figure 5.10: The “full-drilled pillar” undercut design at Esmeralda It is necessary that there be adequate cleaning of flat, narrow undercuts and that the stacking of blasted ore on the tops of the major apices (see Figure 5.11) be avoided. The stacking of ore during undercut formation can lead to choke blasting and the generation of remnant pillars. As shown in Figure 5.11a for the case of an advance or pre-undercut, when the swell is drawn from


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a flat undercut, loose blocks in the undercut back may rest down on the broken ore in the undercut. When the drawbell is developed and the drawing of the broken ore begins, the blasted ore and large blocks sitting on the major apex will not be drawn (Figure 5.11b) and a short-span stable arch may form in the cave back. Obviously, experience and high skill levels are required to satisfactorily implement flat, narrow undercuts. Flat undercuts produce lower tonnages than fan undercuts and so cannot be used as a method of ensuring high initial production rates.

(a)

(b)

Figure 5.11: The ore stacking problem with a flat undercut (a) before, and (b) after drawbell development (Russell 2000)


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5.4.4 Narrow Inclined Undercut

Narrow inclined undercuts are now being used or considered with a view to shaping the major apex so that the undercut will be self-cleaning and stacking and choke blasting will be avoided. The chevron shape of the undercut back also enhances instability and caving of the back as an extreme of the undulating back referred to in the account of fan undercutting. A corresponding disadvantage of the method is the increased likelihood that the initial fragmentation will be coarse. Narrow inclined undercuts are being used particularly where cave mining is taking place in deep, high stress environments as at Palabora (Calder et al 2000) and for Lift 2 of the Northparkes E26 mine (Duffield 2000). As indicated in Section 5.4.1, in order for the undercut to be self-cleaning, the inclination must exceed the angle of friction generated between the blasted and the in situ rock. Experience in this and other forms of mining suggests that this angle should be greater than 45o, usually 50 - 55o. There are several possible variants of the detailed design of narrow inclined undercuts. The designs for Palabora and Northparkes E26 Lift 2 use two undercut drill drifts per drawpoint which produce flat sections of the undercut above each drawpoint. Figure 5.12 shows the narrow inclined advance undercut proposed for Palabora. Figure 5.13 is a generalised representation of this type of design illustrating some of the problems that can occur (Butcher 2000a). A particular problem illustrated is the potential to leave a remnant pillar at the top of the undercut as a result of poorly controlled drilling and blasting.

Figure 5.12: Narrow inclined undercut design for the Palabora underground mine South Africa (Calder et al 2000)


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Figure 5.13: Potential problem areas in a narrow inclined undercut (Butcher 2000a)

Laubscher (2000) discusses two forms of an alternative an alternative design which uses only one undercut drill drift per drawpoint as illustrated in Figure 5.14. Laubscher (2000) argues that this design has advantages over the two drift design in that it provides assurance that the undercut has broken to the top of the major apex and that less development is required on the undercut level reducing costs and the induced abutment stresses. The major apex will also be higher than in the two drift design, providing more space for a secondary drilling level if required. In the first design shown in Figure 5.14a, the potential problem of leaving a remnant pillar at the top of the undercut as illustrated in Figure 5.13, again arises. This problem is addressed in the design shown in Figure 5.14a by over drilling from the advancing side. Even in this case, careful control of drilling is required. A more expensive solution is to develop a drift over the top of the major apex as illustrated in Figure 5.14b. This provides a check on the drilling accuracy and also serves as an anti-socket drift.


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(a)

(b)

Figure 5.14: Narrow inclined undercut with a single undercut drift, (a) without, and (b) with, an anti-socket drift above the major apex (after Laubscher 2000)

5.5 STRESSES INDUCED IN THE UNDERCUT AND EXTRACTION LEVELS

5.5.1 Introduction

The stability of the extraction level excavations and, to a lesser extent, that of the undercut, is critical to the efficient extraction of ore from caving mines. Observations and measurements indicate that the form and timing of the undercut has a significant influence on the stability of the extraction level drifts, primarily because of the high abutment stresses induced in the vicinity of an advancing undercut front (Bartlett 1998, Bartlett and Croll 2000, Butcher 1999, Laubscher 1994).


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Several factors discussed previously have the potential to influence the levels of stress induced in the extraction level excavations including the timing of the undercut relative to the extraction level development, undercut face shape, separation distance between the undercut and extraction levels, cave hydraulic radius, undercut direction and the in situ stress regime. The purpose of the analyses presented in this section is to quantify the effects of these factors on undercut and extraction level stresses. Knowledge of the likely levels of induced stress, combined with an estimate of the rock mass strength, will allow predictions to be made of the resulting levels of damage. On this basis, the undercut strategy, extraction layout design, and support design may then be optimised. Based on experience gained in cave mining operations, some undercut design guidelines have been established that are aimed at minimising stress-induced damage to extraction level excavations (Bartlett 1998, Bartlett and Croll 2000, Brumleve and Maier 1981, Butcher 1999, Lacasse and Legast 1981, Laubscher 1994). As was noted in Section 1.3.3, Bartlett and Croll (2000) describe how changing from a post- to an advance undercut resulted in reduced support requirements and a marked reduction in the rehabilitation of drifts in the BA5 panel at the Premier mine, South Africa. They also found that as the area of the undercut increased the stresses on both the undercut and extraction levels also increased. However, when continuous caving was initiated, the stress levels were observed to drop. It was also found that keeping the leads and lags between adjacent drifts to less than 8 m could reduce the damage to drifts. From observations made at the Bell mine in Canada, Lacasse and Legast (1981) note that the speed of retreat of the undercut is an important factor and that weak zones should be the starting point of an undercut where possible. As noted above, Rech et al (2000) describe the technique used at Henderson mine of developing drawbells from the undercut level close to the cave front. In this way these excavations are subjected to high abutment stresses for less time. From earlier stress measurements and observations made at the Henderson mine, Brumleve and Maier (1981) concluded that heavier support of drifts was necessary in poorer ground which had been exposed to abutment stresses for longer periods of time. This study is discussed in more detail in Chapter 8. Table 5.3 summarises five experiential design guidelines for undercutting developed by Butcher (1999). The ultimate objective of the guidelines is to reduce the level of, or minimise exposure to, high stresses in the vicinity of the undercut front. If this can be achieved, damage to the extraction or production level will be reduced.


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Table 5.3: Experiential design guidelines for minimising damage to the extraction level in caving mines (after Butcher 1999)

Guideline Reasons

1 Use advance undercutting. If advance undercutting is not possible, minimise the percentage extracted for drift and drawpoint development in the production level.

High stresses exist below the undercut front that can cause damage to pre-existing excavations in the production level.

Higher extraction percentages in the production level will increase stress levels there further.

2 Minimise the creation of horizontal irregularities in the undercut front.

Stresses concentrate in these irregularities and increase the level of damage experienced in the production level.

3 Prior to continuous caving being achieved, keep the rate of undercutting greater than the rate of damage to the extraction level.

The longer excavations are subjected to the high stresses below the undercut front, the greater the damage will be.

4 Place the undercut as high as practically possible above the production level.

Stresses decrease with distance below the undercut front.

5 Advance the cave from the weakest ground to the strongest ground to achieve continuous caving as early as possible.

Stresses at the undercut front increase with the hydraulic radius necessary to achieve continuous caving.

Stresses at the undercut front reduce once continuous caving is achieved.

A number of authors have used numerical models to study the stresses induced in undercut and extraction level drifts (eg Barla and Boshkov 1968, Chen 1996, Diering and Stacey 1987, Esterhuizen 1987, Flores 1993, Song 1989). Although these studies give important results, they do not allow adequately for caved block geometry or for a range of in situ stress regimes. A parametric numerical study was therefore carried out by Trueman et al (2002) as part of the International Caving Study Stage I to examine the influence of the following factors on the stresses induced in the undercut and extraction levels in a typical caving mine: • undercut sequence; • in situ stress regime; • separation distance between the undercut and extraction levels; • hydraulic radius of the cave; and • depth below the ground surface.


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5.5.2 Modelling Strategy

In the study carried out by Trueman et al (2002), numerical models were used to obtain predictions of the elastic stresses on the boundaries of the excavations on the undercut and extraction levels. The finite difference code FLAC3D (Itasca 1997b) was used to analyse a three dimensional model able to represent accurately the shapes of the cave and the extraction level excavations. Because of the physical size of a block cave, a two-stage approach to stress modelling was used. In the first stage, a large-scale model of the cave itself was used to determine induced stress levels in the vicinity of the undercut and extraction level drifts. The extraction and undercut level excavations were not included in this large-scale model but the volume of rock between the extraction and undercut levels was given a lower stiffness to account for the increased extraction there. The induced stresses from the large-scale model were then transferred to small-scale models of the undercut and extraction level drifts to obtain the maximum tangential stresses in the undercut, extraction and drawpoint drift roofs. While the stresses may be higher in other areas of the drifts (eg at drift intersections), the changes in the maximum stress in the drift roofs were considered to be representative and useful for illustrating the influence of the various factors on extraction and undercut level stresses. The change in stress around the drawbells was not examined as it requires the use of a more sophisticated method for transferring stresses from the large-scale to the small-scale model (to account for the gradient in stress that occurs below the undercut). The cases examined for post- and advance undercut sequences are summarised in Table 5.4. The directions of the in situ stresses were assumed to be as shown in Figure 5.4. A pre-undercut may be regarded as a special case of an advance undercut and so was not included as a separate case in this study. The extraction level layout assumed was that used at El Teniente 4 South (Flores 1993). In all cases, the undercut height was constant at 4 m. In order to eliminate mining depth as a variable, all stresses were normalised to the vertical stress. Tangential stresses in the roofs of excavations were computed at a range of points from 45 m in advance of the cave front to 45 m underneath the cave. The floor to floor separation between the undercut and extraction levels was 15 m in all cases for which results are presented. A number of runs were carried out at separations ranging from 10 m to 20 m. An approximately 10% difference in the boundary stresses was noted for every 5 m difference in separation. This suggests that the boundary stresses at a 20 m separation would be about 10% lower than those presented here and those at a 10 m separation would be 10% higher.


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Table 5.4: Cases examined in a parametric study of extraction and undercut level stresses (Trueman et al 2002)

Cave Geometry In Situ Stress Ratio

σv:σh1:σh2

(σh1 is parallel and σh2 is

perpendicular to the direction of

cave advance, see Figure 5.4)

Length x

Width (m)

Hydraulic

radius (m)

Height (m)

60 x 60 15 150

0

75

100 x 100 25

150

1:1:1

200 x 200 50 150

60 x 60 15 150

100 x 100 25 150

1:2:1

200 x 200 50 150

60 x 60 15 150

100 x 100 25 150

1:1:2

200 x 200 50 150

60 x 60 15 150

100 x 100 25 150

1:3:2

200 x 200 50 150

60 x 60 15 150

100 x 100 25 150

1:2:3

200 x 200 50 150

The effect of caving height on abutment stresses was investigated by comparing the stresses at the undercut front for cave heights of 0, 75 and 150 m. At a hydraulic radius of 25 m in a hydrostatic in situ stress field (σv=σh1=σh2), the maximum induced stress was very similar for cave heights of 75 m and 150 m. However, the maximum induced stress in both these cases was 15% lower than that predicted for the case in which no caving has occurred (a cave height of 0 meters). This agrees with experiential guidelines which suggest that cave abutment stresses will drop when the cave height increases at the onset of continuous caving. The modelling strategy was partially validated by comparing the modelled induced stresses with measurements made by Flores (1993) at El Teniente 4 South. Reasonable correlations were found.


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5.5.3 Extraction Level Stresses – Post-undercut Sequence

Small-scale models of an LHD extraction level were used to obtain estimates of the maximum induced stress around the production drifts for an undercut to extraction level separation of 15 m. Stresses from the large-scale models were used to specify initial conditions in the small-scale models. Figure 5.15 presents the calculated maximum tangential stresses normalised to the in situ vertical stress in drift roofs for each in situ stress regime studied plotted as a function of distance from the cave boundary for varying hydraulic radii. Figure 5.15 shows that the stresses in the extraction level production drifts generally increased with hydraulic radius. This is in accord with the experiential guidelines. In general, the maximum tangential stress in the roofs of production drifts increased by about 20% with a doubling of hydraulic radius to achieve continuous caving. Exceptions were for σv: σh1: σh2 = 1:2:3 and 1:1:2 (σh1 parallel and σh2 perpendicular to the direction of cave advance) where stresses in the production drift roof were high but largely unaffected by the size of the cave. For all of the in situ stress regimes modelled except those noted above, there was a significant fall in induced stress from the peak as the cave passed over the drifts. For the majority of in situ stresses, significant stress changes were therefore apparent, with the induced stress levels rising quite significantly as the cave approached that section of the drift and falling sharply as the cave passed over the top. 5.5.4 Extraction Level Stresses – Advance Undercut Sequence

An advance undercut sequence was examined for the in situ stress cases summarised in Table 5.4. Advance undercutting was examined by making two changes to the large-scale post-undercut model: • The undercut was extended out from the cave front by a specified distance. The broken

rock in the undercut was assumed to have the same properties as the caved material. • The stiffness of the extraction level ahead of the cave front was increased to represent

partial development. For this study, partial development ahead of the cave front consisted of production drifts only or production and drawpoint drifts.


218

Post undercut/σv:σh1:σh2=1:2:1Production drift roof

1

2

3

4

-60 -45 -30 -15 0 15 30 45 60Distance from cave boundary, x(m)

hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)Post undercut/σv:σh1:σh2=1:1:1

Production drift roof

0

1

2

3


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


2

3

4

5


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


0

1

2

3


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


2

3

4

5


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)

Figure 5.15: Predicted normalised tangential stresses in production drift roofs

for a post-undercut (Trueman et al 2002) The small-scale model with only the production drifts or the production drifts and drawpoint drifts excavated was used to examine the stresses in a partially developed extraction level below the advance undercut. The results are shown in Figures 5.16, 5.17 and 5.18. In general, the stresses induced at the excavation boundaries again increase with hydraulic radius. For most in situ stresses and cave orientations, a doubling of the hydraulic radius leads to an approximately 20% increase in induced stress. Exceptions can occur in the roofs of drifts where the major principal stress is horizontal and parallel to the direction of drift advance.


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Induced stresses in production drifts are similar whether drawpoint drifts are excavated in advance of mining or not. In most in situ stress regimes, with the exception of some in which the major principal stress is horizontal and perpendicular to the drift direction, significant stress changes occur. In the main, induced stress levels are significantly lower with the cave front 15 m in advance of the section of drift being investigated. This result gives some credence to “the 45 degree rule”. However, induced stress levels continue to fall under most stress conditions, albeit at a reduced rate, as the distance between the cave front and drift section increases. The extent of the stress changes decreases significantly for excavations that are not in advance of the cave front.

Advanced undercut without drawpoint drifts/σv:σh1:σh2=1:2:1Production drift roof

1

2

3

4

-60 -45 -30 -15 0 15 30 45 60Distance from undercut front, x(m)

hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)


0

1

2

3


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)


2

3

4

5


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


2

3

4

5


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


4

5

6

7


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)

Figure 5.16: Predicted normalised tangential stresses in production drift roofs

for an advance undercut without drawpoint drifts (Trueman et al 2002)


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Advanced undercut with drawpoint drifts/σv:σh1:σh2=1:2:1Production drift roof

1

2

3

4


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)Advanced undercut with drawpoint drifts/σv:σh1:σh2=1:1:1

Production drift roof

0

1

2

3


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)


1

2

3

4

5


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


2

3

4

5


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)


4

5

6

7


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)

Figure 5.17: Predicted normalised tangential stresses in production drift roofs for an advance undercut with drawpoint drifts (Trueman et al 2002)


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Advanced undercut with drawpoint drifts/σv:σh1:σh2=1:2:1Drawpoint drift roof

2

3

4

5

6


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


0

1

2

3

4


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)


4

5

6

7

8


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


1

2

3

4


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)


3

4

5

6


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s (σ

v)

Figure 5.18: Predicted normalised tangential stresses in drawpoint drift roofs for an advance undercut with drawpoint drifts (Trueman et al 2002)

5.5.5 Undercut Level Stresses

A small-scale model assuming 4 m wide and 4 m high drifts was constructed for the undercut level and stresses obtained from large-scale models at this level imposed upon the model for the in situ stress regimes noted previously. The maximum tangential stresses induced in the roofs of drifts are shown in Figure 5.19.


222

σv:σh1:σh2=1:2:1Undercut tunnel roof

1

2

3

4

5

6

7

8

0 15 30 45 60Distance from cave boundary, x(m)

hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)σv:σh1:σh2=1:1:1

Undercut tunnel roof

1

2

3

4

5

6


hr= 50 mhr= 25 mhr= 15 mT

ange

ntia

l str

ess

(σθ)

In s

itu v

ertic

al s

tres

s ( σ

v)


2

3

4

5

6

7

8


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


1

2

3

4

5

6


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)


1

2

3

4

5

6

7


hr= 50 mhr= 25 mhr= 15 m

Tan

gent

ial s

tres

s (σ

θ)

In s

itu v

ertic

al s

tres

s ( σ

v)

Figure 5.19: Predicted normalised tangential stresses in undercut drift roofs (Trueman et al 2002)

In general the magnitudes of the induced stresses increase with increasing hydraulic radius up to continuous caving. A doubling of the hydraulic radius leads to an approximately 30% increase in induced stress close to the cave front. Induced stresses are generally higher the closer the drift section is to the cave front and fall rapidly over the first 15 m away from the front. Drift sections therefore experience a gradual increase in induced stress as the cave front approaches. The magnitudes of the maximum induced stresses are, of course, higher on the undercut level than on the extraction level.


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5.5.6 Summary of Parametric Study Results

The results of the parametric study carried out by Trueman et al (2002) are generally in accord with the qualitative experiential guidelines summarised in Table 5.3. The study has permitted a number of these experiential guidelines to be quantified and bounded in the following ways.

For extraction level drifts: • as expected, the magnitude of induced boundary stresses in drifts is sensitive to the in situ

stresses and their orientations; • if the hydraulic radius to achieve continuous caving doubles, the maximum induced

stresses in the extraction level drifts increase by approximately 20% in most in situ stress environments. Exceptions to this occur in drift roofs where the major principal stress is approximately horizontal and perpendicular to the direction of drift advance;

• when the vertical separation between the undercut and extraction levels is in the range of

10 m to 20 m, an increase or decrease in the separation distance of 5 m leads to an approximate 10% difference in induced stresses at drift boundaries; ie an increase in separation of 5 m leads to a 10% decrease and a 5 m decrease leads to a 10% increase in boundary stresses;

• when continuous caving is achieved for a hydraulic radius of 25 m in a hydrostatic in situ

stress field, the maximum induced stress reduces by 15% and the vertical induced stress reduces by 30% at the undercut level. In the extraction level drifts, the maximum induced stress reduces by 2% and the vertical induced stress reduces by 14%;

• significant induced stress changes occur for many in situ stress states; ie stresses increase

as sections of drifts are approached by the advancing cave and decrease as the cave passes overhead. Exceptions to this general rule occur in drift roofs in which the maximum in situ principal stress is approximately horizontal and perpendicular to the direction of cave advance;

• in the main for an advance undercut, induced stress levels are significantly lower with the

cave front 15 m in advance of the section of drift being investigated. This gives some credence to “the 45 degree rule”. Exceptions occur again in drift roofs where the maximum in situ principal stress is approximately horizontal and perpendicular to the direction of drift advance. Induced stress levels generally continue to fall, albeit at a reduced rate, more than 15 m from the cave front; and


224

• for an advance undercut, any excavations in advance of the cave front are subjected to induced stresses similar to those in a post-undercut. Only the excavations formed behind the cave front significantly benefit from an advance undercut.

For undercut level drifts: • in general the magnitudes of the induced stresses increase with hydraulic radius up to

continuous caving being achieved; • induced stresses are generally higher the closer the drift section is to the cave front, falling

rapidly over the first 15 m away from the front. Exceptions are in drift roofs where the maximum principal in situ stress is horizontal and perpendicular to the direction of drift advance; and

• the magnitudes of the maximum induced stresses are higher on the undercut level than on

the extraction level. 5.5.7 Undercut Drift Support and Reinforcement

The provision of drift support and reinforcement forms an important part of undercut planning and design. Although they are not permanent excavations, the undercut drifts are vital components of an overall block or panel cave operation and must remain safe and accessible to men and machines throughout their design lives. As part of the International Caving Study, Stage I details were collected of a range of design and operational features of most currently operating caving mines. Summaries are given below of the support and reinforcement used in undercut drifts at four of these mines. Further details are given by Trueman et al (2002). Andina Panels II and III

Panels II and III at CODELCO-Chile’s Andina mine were both extracted using a post- or conventional undercut strategy. Panel II was located entirely in a relatively weak rock mass known locally as secondary rock. The separation between the undercut and extraction levels was 15 m. The in situ stress regime was σv: σh1: σh2 = 9: 18: 13 MPa and the average Q' of the rock mass was 0.4, equivalent to an average RMRL of 35. Continuous caving was reported to have occurred at a hydraulic radius of 26 m. The average uniaxial compressive strength of the intact host rock was estimated to be 108 MPa. Panel III has a mixture of secondary (weak) and primary (moderately strong to strong) ore. Initial continuous caving was achieved in the weaker rock mass (having an average Q' of 0.4) at a hydraulic radius of 11.7 m. The in situ stress regime was σv: σh1: σh2 = 17: 22:13 MPa. The average uniaxial compressive strength of the intact host rock was estimated to be 108 MPa.


225

Drift support and reinforcement was the same in both panels. In the undercut levels the drifts were generally reinforced only with spot bolting but wooden props were often erected in the immediate vicinity of the cave front. Generally drift conditions were good on both the undercut and extraction levels and, where placed, the support and reinforcement was adequate for the conditions in all sections of the drifts relative to the cave front. Esmeralda sector, El Teniente

The Esmeralda sector of the El Teniente mine, Chile, is being extracted using a pre-undercut strategy (Rojas et al 2000b). The rock mass had a Q' of 5.3 up to continuous caving being achieved at a hydraulic radius of 27 m. The in situ stress regime was σv: σh1: σh2 = 26: 34: 34 MPa. The separation between the undercut and extraction levels was 18 m. The average uniaxial compressive strength of the intact host rock was estimated to be 100 MPa. The undercut level drifts are 4 m wide and 3.6 m high. These drifts are reinforced with 2.3 m long resin anchored bolts on a 1.0 m spacing and mesh. Generally the reinforcement is adequate, with damage being confined to local areas in which the rock mass is more fractured. Northparkes E26, Lift 1

The Northparkes E26, Lift 1 block cave used an advance undercut, with the extraction level and drawpoint drifts only being developed in front of the advancing undercut. The in situ stress regime was σv:σh1:σh2 = 12: 23: 15 MPa. The average Q' of the rock mass in the vicinity of the drifts was 8.7 (or an average RMRL of 53). The undercut was extended to a hydraulic radius of 44 m but continuous caving was not achieved and the cave was induced until it reached an overlying weaker rock mass. As shown in Figure 5.6 a double undercut was extracted. The lower undercut was developed between a fully developed upper undercut and the extraction level. The stress charts presented in Sections 5.5.4 and 5.5.5 do not take such a scenario into account. The average uniaxial compressive strength of the intact host rock was estimated to be 110 MPa. In the upper undercut, the drifts were 4.2 m wide by 4.5 m high. Installed reinforcement consisted of 2.1 m long split sets on a spacing of 1.25 m with 8 bolts per ring. In general, experience showed the reinforcement to be adequate. Palabora The Palabora block cave is being extracted using an advance undercut (Calder et al 2000). The average Q' of the main host rock was estimated to be 23 with the average RMRL being in the mid-70s. In situ stress measurements had indicated a hydrostatic state of stress of 38 MPa. The average uniaxial compressive strength of the intact host rock was estimated to be 140 MPa.


226

The undercut level drifts are 4 m wide by 4 m high. Reinforcement consists of 2.4 m long resin grouted rebar on a spacing of 1.25 m. Additionally, weld mesh is installed with 1.0 m long split sets being used to secure the mesh in place. Fibre reinforced shotcrete nominally 50mm thick is used where ground conditions are considered to warrant it. Some spalling of the sidewalls is evident remote from the cave front even after bolts are placed but before mesh or shotcrete are installed. The shotcrete and mesh prevent further problems and the full support system appears to be working well close to the cave front at a hydraulic radius of 16 m. 5.6 DRILLING AND BLASTING FOR UNDERCUTTING AND DRAWBELL CONSTRUCTION

5.6.1 Introduction

The requirement for proper design and accurate implementation of drill and blast strategies and practices should be stricter in block caving than in other underground mining methods. The consequences of poor drilling and blasting practices during block cave construction can be severe and can impact on the initial performance of a block cave (cave initiation) and the subsequent integrity of the extraction level drifts and drawpoints. Proper selection of blasting parameters such as hole diameter, explosive type and initiation sequence is therefore crucial. Just as important is the selection of drilling and blast hole charging equipment. The blasting parameters should be selected to suit the geomechanical properties of the rock mass to be blasted. The drilling equipment should be able to adequately and efficiently drill the required undercut, drift and drawbell geometries. The explosive loading equipment should be able to load the blast holes to design. This is particularly important when horizontal, inclined and up-hole blast hole loading is required as is now common practice in block cave undercutting (see Section 5.4). Controlling blast performance and obtaining results to the level required during block cave construction was more difficult to achieve before the 1990s given the quality and reliability of some of the explosive products and accessories available at the time (Cameron and Grouhel 1990). However, the 1990s saw significant advances in commercial explosives technology including accessories. There is now available a wider range of gassed and variable density emulsions, low shock energy emulsions and ANFO/emulsion blends suitable for a wider range of hole diameters. In addition, very low velocity of detonation (VOD) products (eg diluted ANFO products and gassed emulsions) and highly accurate electronic and pyrotechnic detonators are now available. During the 1990s, significant advances were made in drilling equipment enabling easier alignment of underground drill rigs and more accurate drilling of holes. The technology and equipment for blast hole loading has also improved significantly as has understanding of rock breakage principles and explosive-rock interaction (eg Hustrulid 1999). With these developments, the opportunity now exists to “engineer” blasts to a level not feasible previously.


227

During block cave construction the drilling and blasting objectives should be to ensure: • full breakage of the undercut rings without leaving remnant pillars; • minimum damage to the perimeter walls during extraction level drift development; • minimum damage to pillars during drawpoint construction; and • achieving the required drawbell shape with minimum damage to the drawpoint brow. This means that proper rock breakage design principles and rules should be applied during the design of undercut and drawbell rings and basic wall control principles should be applied during extraction level construction. 5.6.2 Factors Affecting Drilling and Blasting Performance

Through its wide research and field experience, the JKMRC has developed a number of experiential or empirical drilling and blasting guidelines that are supported by more fundamental investigations (Scott 1997). These have been implemented successfully at a number of current block caving operations (Guest et al 1995). The following sections discuss some of the design guidelines and rules that are applicable to block and panel cave drilling and blasting. Additional and practical principles that are also applicable to underground blasting are given by Hustrulid (1999). Rock parameters influencing blast performance

The key intact rock and rock mass parameters that influence blast performance are: • rock density; • strength of intact rock blocks defined by the jointing; • jointing (intensity, orientation, persistence and condition); and • degree of attenuation and mode of failure. . In a given geological or geotechnical domain, any one or combination of the above parameters can have a significant influence on blast performance. The influence of these parameters is greater when blasting is carried out in confined environments such as undercuts and drawbells. In designing a blast for a given geotechnical environment, the engineer should assess which of the above factors are likely to control blast results and then take appropriate design measures such as those discussed below. Rock density

Rocks typically blasted in mining have densities in the range 2.1 to 4.8. In blasting, rock density has a pronounced effect on the initial rock displacement or movement and the subsequent in-flight particle velocities or throw. For a given amount of energy, it is generally more difficult to break efficiently and displace higher density rocks. In terms of blast design,


228

and where two or more rings are blasted at any one time, rock density should influence the choice of the inter-ring timing. Strength of intact rock blocks

The strength of the in situ rock blocks isolated by the jointing is usually described in terms of either static or dynamic uniaxial compressive strength (UCS) values. In terms of blast design, the intact rock blocks should influence the density of holes drilled (S/B) and the characteristics of the explosive required. The explosive characteristics include, for example, detonation pressure, VOD, weight strength and brisance. Brisance is a property which describes the explosive’s ability to shatter a rock. Jointing

As discussed in Chapters 2 and 4, the intensity, persistence and orientation of jointing define the sizes and distribution of the rock blocks that are either fully or partially formed in situ. The size and distribution of the in situ blocks determine the degree of breakage that might be required in order to achieve the desired fragmentation. Joint condition describes whether the joints are dilated, filled, closed or healed and in essence define joint strength. Joint strength determines how easily in situ rock blocks can be liberated during blasting. Healed joints can be as hard or harder to break as intact rock blocks. In terms of blast design, jointing and the corresponding size of the rock blocks formed should influence the characteristics of the explosives required (eg gas volume or VOD), the blast pattern (burden, B, and spacing , S,) and the design S/B ratio. The S/B ratio defines the density of holes which should provide the optimal distribution of explosive energy required to adequately fragment the in situ blocks. Where sub-optimal ratios are used, fragmentation can be structurally controlled with the resulting fragmentation distributions being equivalent to those of blocks in situ. In the case of high or double undercuts, achieving good fragmentation is important to ensure high loader productivity. Degree of attenuation and mode of failure

An important rock parameter often ignored during blast design is the rock’s ability to transmit and absorb propagating blast energy and the rate at which that energy is absorbed. In seismology, the term attenuation is often used to describe this property. Different rocks can be assigned different attenuation constants. Highly attenuating rocks absorb propagating blast energy readily. There is a relationship between the rock matrix (grain size) and attenuation. For example, fine grained materials tend to be more highly attenuating; that is, the rate at which they absorb blast energy is greater than for rocks with a larger grain size (Chitombo 1991).


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Highly attenuating rocks also tend to fail “plastically” during blasting and low attenuating rocks generally fail in a more brittle manner. Rocks showing brittle behaviour generally transmit blast energy over relatively longer distances. The corresponding energy is likely to be evenly distributed resulting in even breakage. On the other hand, rocks that behave plastically, often referred to in mining as “soft” or “spongy”, rapidly absorb applied blast energy in the near vicinity of detonating blast holes. The resulting distribution of blast energy over the rest of the blast volume is poor and the breakage uneven. This is particularly the case where large blast hole spacings are used. Rocks showing brittle and plastic behaviours often require the application of different blast parameters to achieve the same degree of breakage. Soft, plastic rocks are generally more difficult to fragment adequately. In terms of blast design, the rock parameters which influence attenuation should also influence the choice of explosives (VOD or rate of energy release) and blast hole timing. The characteristics of an explosive are more critical in soft or weak rocks than in stronger rocks. 5.6.3 Experienced Based Design “Rules of Thumb” for Rock Breakage Control

The following sub-sections discuss design guidelines or “rules of thumb” that should be considered during drill and blast design to ensure optimal rock breakage. The rules presented are applicable to undercut blast design. Optimum S/B ratio

Size of burden

The empirical formulae reported in the literature for selecting a suitable burden (the distance between rings) for different rock masses, are mainly for unconfined, free-face blasting and not directly applicable to confined blasting. However, from experience and depending upon hole size, it is suggested that the burdens used in confined situations such as during undercutting should, as a general rule, be kept to a minimum. For small hole diameters (51 mm, 64 mm and 76 mm), the burdens should be in the range 1.2 m to 2.0 m, particularly when blasting soft/plastic type rocks. For larger hole diameters (eg 89 mm and 102 mm) the burdens should still be kept small with a suggested maximum of 2.5 m (based on experience). Burdens of 3.0 m have been used with the larger hole diameters but toe breakage problems are often experienced. Maximum toe spacing

The choice of maximum blast hole toe spacing (S) should be governed by the size of ring burden (B) being blasted. For a given rock mass, there is an optimum S/B ratio that provides the optimum hole density, explosive energy distribution and therefore volume breakage. JKMRC experience is that the optimum S/B ratios applicable to a wide range of rock types and


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conditions are in the range of 1.1 to 1.4. This is consistent with figures given by Persson et al (1994). For soft, massive and/or sparsely jointed rock masses, the recommended S/B should be small and approaching 1.1. Under no circumstance should it be less than 1.0. S/B ratios of less than 1.0 tend to produce the so-called pre-split effect in which the breakage of the burden is poor. Ratios equal to 1.0 are not necessarily optimal in terms of the resulting breakage. Ratios of up to 1.3 have been found to be adequate for blasting some of the soft De Beers kimberlite rocks (Guest et al 1995) and the soft ultramafics rocks in Western Australia. For heavily jointed rock masses the S/B can approach 1.4. The influence of structure becomes prominent where ratios are equal to or greater than 1.4. When determining toe spacings, it is important that a consistent method of calculating toe spacing is used. Figure 5.20 shows examples of conventional methods of calculating toe spacing.

Figure 5.20: Conventional method for calculating toe spacing


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Choice of explosive type

Based solely on the manner in which commercial explosives deliver energy into rock, the general explosive-rock interaction criteria shown in Table 5.5 can be used as a guide.

Table 5.5: General explosive-rock interaction criteria

Type of Rock Explosive Characteristic Required

Hard (and brittle) Rock Emulsions or High VOD Watergels

Soft (and plastic) Rock ANFO or Heavy ANFO

Very Soft Rock Diluted ANFO

The terms hard and soft rock relate not only to the intact rock strengths and mode of failure but also to the spacing and condition of the jointing. High VOD emulsion and water gel explosives have relatively high weight/bulk strengths and, more importantly, deliver their energy early and relatively quickly. They are therefore suitable for use with hard and brittle rocks. ANFO and heavy ANFOs tend to deliver their energy over longer periods. This makes them more suited to softer and more plastic rocks. However, before any pure emulsion or emulsion / ANFO blend type explosives products are considered, it is imperative that suppliers are asked to provide the following specifications for their products: • the critical diameters of the products; • sensitivity of the products in the hole sizes to be used. This is particularly important for

small diameter holes; • for emulsions and emulsion blends sensitised using micro-balloons, the likelihood of the

products being shock desensitised should be quantified. This is particularly important given the size of hole diameters and the relatively small blast hole spacings being currently applied; and

• the VOD versus hole diameter relationships of the products (including ANFO). Any one of these factors can affect efficient detonation of the products, inevitably influencing blast performance and results. Average powder factors (kg/t) and corresponding energy distributions

During undercutting, the optimum powder factor (kg/t) is that which ensures full and proper breakage of a ring, in particular the “toe” and “shoulder" sections. However, the breakage of the full blast volume is dependent upon the distribution of the explosive energy and not necessarily on the average powder factor. The concept of explosive distribution is discussed later.


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In practice, both the average powder factor and the optimum explosive distributions are obtained through trial and error methods. To speed up this optimization process, the JKMRC has developed software (JKSimBlast) to calculate the average powder factor and the corresponding energy distribution. Energy distribution is dependent upon rock type, hole diameter, the S/B ratio, explosive type and timing. Experience in blasting soft and plastic kimberlite rock at the Kimberley and Koffiefontein Mines in South Africa, showed that a powder factor of 0.35kg/t at the toe ensured full breakage of relatively confined undercut rings. During the mining of the Northparkes E26 Lift 1 block cave, powder factors in the range 0.25kg/t – 0.3kg/t provided the required undercut ring breakage. In most hard rock and confined blasting situations, powder factors in the range 0.25kg/t – 0.35kg/t are generally recommended. The charging geometry is also important in achieving the required explosive distribution. A recommended charging configuration and rule is given in Figure 5.21. The recommended charging also ensures that the concentration at the collar is reduced thereby minimizing the potential for brow damage.

Figure 5.21: Recommended blast hole

charging patterns

Combination of patterns used to distribute the charge


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Blasthole timing

The use of high powder factors does not necessarily ensure the efficient breakage of a volume of rock. Efficient breakage of a blast volume is better promoted by the manner in which blast energy is distributed and imparted into the rock. This is controlled by blast hole timing. Single hole or hole-by-hole firing with an inter-hole timing in the range 17 to 30 ms is generally recommended for improving breakage in most hard rock blasting. This ensures that the energies from adjacent holes contribute to the breakage of the volume of rock straddled by the adjacent holes before displacement. With electronic or any super accurate system where overlapping of delays is unlikely, the inter-hole timing can be in the range 9 to 25 ms. In general, faster times are more suited to hard and brittle rocks and slower times to more “plastic” and high density rocks. Firing ring holes on the same delay (cluster blasting) has been shown through experience to produce poor breakage particularly in soft rocks. In such cases the fragmentation is generally uneven, consisting mainly of fines and coarse fragments in a bi-modal distribution. Cluster blasting in not a recommended practice during undercutting or drawbell blasting. This method of blast hole firing may also cause damage to the adjacent rock. The influence of timing on fragmentation is discussed in a paper on the evolution of compound rings by Guest et al (1995). Based on the JKMRC’s experience in blasting TKB kimberlite, the inter-hole delay should be less than 40 ms. Analysis of high speed films taken at Koffiefontein and Finsch mines in the early 1990s showed that 3.2 m burdens using 89 mm holes charged with an emulsion product, moved after 40 ms to 50 ms of detonation. This means that if delay times of ±40 ms are used, then the adjacent holes will work independently resulting in poor fragmentation and an increased likelihood of hole cut-off. Historically, it has been difficult to control fragmentation because of the scatter or dispersion associated with delay elements. However, there has been significant improvement in pyrotechnics and the delay scatter in current systems is minimal thus making it possible to better control blast fragmentation. In selecting delay elements it is important that suppliers are asked to provide relevant performance data. Hole diameter

The selection of hole diameter is generally influenced by the available equipment and, more importantly, by the lengths of the holes to be drilled. For the hole length generally drilled in most block cave undercuts (ie less than 30 m), 64 mm, 76 mm and 89 mm are suitable diameters with 102 mm diameter holes considered the recommended maximum. Larger diameter holes also mean that the potential for increased powder factors and kilograms of explosive per delay, which increases the potential for blast damage.


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5.6.4 Undercut Drilling and Blasting

As has been discussed earlier in this Chapter, the successful development and extraction of an undercut is essential for the initiation of the caving process. The best drilling and blasting practices must be implemented to avoid poor breakage resulting in the formation of remnant pillars (Figure 5.22) which not only inhibit the initiation of caving but also act as loading points or “stress channels” thereby increasing the likelihood of damage to extraction level drifts, drawpoints and support. The planned geometry of an undercut should govern the drilling and blasting parameters and practices to be used. The following criteria should be considered during design: • ease of drilling to ensure proper distribution of holes required to completely break the rock

volume thereby avoiding the formation of pillars; • ease of loading or clearance of blasted material to avoid blasting of subsequent rings under

severe choked conditions; • potential for hole losses or closures during undercutting due to either blast damage, stress

abutment effects or relaxation of the rock mass; • ease of managing the undercutting sequencing and achieving the required undercut front

shape; • problems likely to be encountered during undercutting and ease of recovery from such

problems; and • potential for large failed blocks reporting early into the blasted undercut immediately

following undercutting. The types of undercut geometries used have been discussed in Section 5.4. Regardless of the type of undercut geometry used, the following critical drilling and blasting factors should be considered: • height of undercut; • hole inclination; • amount of explosives or explosive energy distribution; and • amount of blasted material to be loaded. Height of undercut

The influence of the height of the undercut on undercut and caving performance was discussed in Section 5.3.5 and examples of undercut heights used, in particular at El Teniente, were given in Table 5.2. Purely from a breakage view point, higher undercuts are easier to break because of the increased free-face area. The potential for not achieving full breakage is greater in narrow undercuts because of confinement. Any small amount of hole deviation can also exacerbate the confinement problem.


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Figure 5.22: Examples of remnant pillars A number of block caving operations have encountered significant drilling and blasting problems during the mining of narrow undercuts. In Esmeralda, El Teniente, changes to the drilling geometries were made to improve breakage as shown in Figures 5.9 and 5.10. Bell Mine increased the undercut height from approximately 3.2 m to 6.0 m also as a way of improving breakage (Figure 5.23).


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Figure 5.23: Undercut geometry at Bell Mine, Canada Hole inclination and ring geometry

Narrow flat undercut

For a narrow flat undercut, parallel or near parallel holes are considered ideal. However, the ability to do this will depend upon the capabilities of the drilling equipment available. Where the drilling equipment is unable to drill parallel holes, then a fan from a single drilling point can be utilized. These alternative drilling geometries are shown in Figure 5.24.

Figure 5.24: Typical drilling geometries for a narrow flat undercut


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Parallel holes provide the best explosive energy distribution. Fan drilling from a single point will require slightly more holes to provide the same explosive distribution. To achieve the required throw with either drilling configuration the holes should be angled slightly forward. From experience the recommended angle range is 20° to 30°. Uphole or SLC undercut rings

The recommended practice is to slightly dump the uphole rings forward by 10° or 15°. Toe breakage has been demonstrated in practice to be better with the slightly forward dumped rings. Vertical or 90° rings often result in “crown formation”. The impact of hole inclination is well illustrated in the Block Cave Manual (Laubscher 2000) by the work of Bell at Shabanie Mines (Figure 5.25).

Broken ground

Planned limit of break

Area unbroken justshatteredProbable break outline

Direction of retreat Development

Undercut with 90 degree rings

Broken ground

Planned limit of break

Area unbroken justshatteredProbable break outline

Direction of retreat Development

Undercut with 70 degree rings

Figure 5.25: Impact of ring inclination (Laubscher 2000)


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Explosive energy distribution

The powder factor is an index often used to determine the amount of explosive required to adequately fragment a mass or volume of rock. It is simply the number of kilograms of explosive required per tonne or cubic metre of rock to be broken (ie kg/t or kg/m3). In practice, engineers experimentally determine a range of powder factors or the average powder factor required to achieve optimum breakage and fragmentation. Powder factors to suit a given rock mass can also be approximated using blastability indices (eg Lilly 1986). Unfortunately, an average powder factor assumes that the explosive energy is uniformly distributed within the blast volume. While this may be a good approximation in a bench or parallel hole geometry, the concept is not strictly correct in the case of a ring or fan blast hole geometry. The JKMRC has introduced the concept of the three dimensional explosive energy distribution as a better way of accounting for the explosive energy within the rock volume to be blasted (Kleine 1988). The methodology considers that each point in the rock will have an energy value arising from each detonating blast hole, determined by the distance between that point and all detonating holes and the amount of explosive in each hole. With reference to Figure 5.26, the powder factor calculation was extended by considering a small infinitesimal segment of charge and writing the equation for the resulting explosive concentration at a point P for a sphere centered at the charge segment. The general form of the equation is:

( )

∫+πρ

⎟⎠⎞

⎜⎝⎛πρ

=2L

1L 32

22r

2

e

dllh

34

2D..1000

P (5.1)

Equation (5.1) can be integrated and rewritten as:

⎟⎟⎠

⎞⎜⎜⎝

⎛−

ρρ

=1

1

2

22

2

r

e

rL

rL

h1D5.187P (5.2)

Special conditions apply to the above relationships at the charge axis (ie h = 0) and at very large distances (ie h ∞). The explosive concentration at any point in three dimensions is determined by solving the appropriate integrated form of the equation for each explosive charge and summing the values.


239

Figure 5.26: Calculation of the 3D explosive energy concentration at point P

The explosive energy concentration calculated by this method will generally be slightly higher than the conventional powder factor. However, the values will be similar at the blast’s designed burden. The highest concentration of energy is in and around each blast hole. The JKMRC blasting software incorporates this model. The energy distribution can be calculated in any plane in three dimensions. This approach has been applied successfully to undercutting at the De Beers caving operations and at the Northparkes E26 Lift 1 block cave. In both instances, the undercuts were successfully mined without leaving pillars. Figure 5.27 shows an example of the analysis used to calculate the distribution of explosives in some of the undercut rings at Northparkes. From the analysis it is possible to assess the effect of increased burden or number of holes. The idea is to achieve the required powder factor in the toe regions in the case of fans.


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Reference undercut ring - NORTHPARKES Lift 1

Toe Spacing = 3m, Explosive = Emulsion 1 g/ccRock SG = 2.8

1.6 m Burden 2.0 m Burden

2.5 m Burden

3D Explosive Energy Distribution

Figure 5.27: Use of the 3-D explosive energy distribution to optimise ring design


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Amount of blasted materials to be loaded

Conventional and recommended practice during undercutting is to remove some material to avoid blasting under “choked” conditions. As a general rule, 20% to 40% of the blasted volume is removed which is equivalent to removing the “swell”. This ensures that the subsequent rings are not blasted under full confinement. Laubscher (2000), however, suggests that in the case of narrow flat undercuts, sufficient material should be left to support the backs thereby reducing abutment stresses. This is analogous to the effect of backfilling stopes in narrow gold reef mining, a practice that has been shown to decrease abutment stresses. Experience has shown that when excessive material is removed, abutment loading can result. 5.6.5 Drawbell Blasting

Figures 5.28 and 5.29 provide examples of some of the drawbell shapes currently used. In designing the most optimal drawbell shape and size, the following drilling and blasting issues should be considered: • ease of drilling the patterns considered optimal for achieving the bell size and shape

required to induce flow; • the ability to shape the major apex pillar so as to minimize the width of the apex ; • the practicality of drilling wider fan holes required to shape the drawbell to the preferred

dimensions from the relatively short base of the drawbells; and • ease of charging the rings to achieve the required powder factors and explosive energy

distributions for efficient drawbell blasting without excessive charging of ring collars given the concentration of holes.

It is also recommended that the number of blasts or stages used to open a drawbell be minimised. An added benefit of the reduced stage drawbell opening sequence is increased safety for operators when charging. The time spent within the blasted drawbell area is significantly reduced. The impact of vibrations from the different stages is also minimised. The recommended stages are: • raising; • slotting; • stripping of fans into the slot on one side; and • stripping of fans into the slot on the second side.


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A common and recommended drawbell raising practice is “blind raising” using large diameter (660 mm and 1200 mm) holes drilled using equipment such as the RD 2000 from Master Drilling. The recommended hole diameters for drawbell blasting are 64 mm and 76 mm with 89 mm being the recommended maximum size. A recent innovation in drawbell blasting was made at the Premier Mine’s BA5 block cave where the drawbells were blasted using electronic delay detonators. This enabled hole by hole firing which effectively ensures that each hole is primed and detonated and that the resulting amount of explosive detonated per delay is greatly reduced. This is important for blast damage considerations.

Strip A Strip BSlot

Raise

Stage 1- Stripping

Stage 2- Slot

Stage 3 and 4- Strip A & B

14m

5.2m

3.6m

5.2m

10m

Figure 5.28: Skull shape drawbell for Northparkes E26 Lift1 block cave and the blasting sequence used


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Excavation 1° - 2° Phase

Cross Section

Figure 5.29: Drawbell shapes and sizes used for the Esmeralda section,

El Teniente Mine, Chile 5.6.6 Drilling Equipment Selection

Proper selection of drilling equipment is crucial for producing the required undercut, drawpoint and drawbell shapes. The selection criteria should include: • size of unit; • boom coverage; and • ease of drilling horizontal and angled holes.


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Almost all major equipment manufacturers now have or can configure rigs to meet most requirements for undercut and drawbell drilling. Examples of rigs that can be considered for block cave applications are described below. Mercury 1LC10 (D4 – E50)

This is a high power compact top-hammer long hole drill, with a four wheel drive articulated hydraulic jumbo, for fully mechanised production from medium sized drifts. This Tamrock rig can be easily operated in 4.2 m wide drives and has a good boom coverage because of its flexible boom head. This means that horizontal holes can be drilled more easily and closer to the floor. Standard 1.8 m rods can be used. The rig is capable of drilling 76 – 89 mm holes when equipped with the Hydrastar 300 SR. Solomatic 620

The Tamrock Solomatic 620 is a flexible one boom electrohydraulic long-hole rig for production drilling underground. The boom has a wide parallel drilling coverage, long boom extension, 360° rotation and wide tilt angle ranges forwards and backwards, thus offering wide drilling variety. The rig is capable of drilling 64 – 89 mm holes up to 32 m long. This is more suited to wider tunnel sizes, nominally 4.5 m, if conventional 1.8 m rods are to be used. If this unit is to be used without modification, then undercut drives have to be mined to a width of 4.5 m. Other suitable rigs include the Atlas Copco Simba 1354 Series. It is important that the rigs are adequately equipped with appropriate angling devices to permit the proper drilling of angled holes. Accessories

Rig and hole alignment systems now exist that can be mounted on drilling rigs. One such device is the Transtronic angle indicator type 6H. The system can be installed in other types of rigs. It comprises two gravity type sensors that measure boom inclination and feed rotation. The device can measure side angle measurements to 360° with an accuracy of ± 0.2°. The inclination angle measurement range is ± 60° with an accuracy of ± 0.8°.

The application of tubes is also recommended for accurate hole drilling. Tubes are now available to suit 76 mm and 89 mm diameter holes.

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CHAPTER 6

EXTRACTION LEVEL DESIGN 6.1 PURPOSE

he operational efficiency and cost effectiveness of block and panel caving mines depend, among other things, on the design and performance of the extraction or production level excavations. As earlier discussions have indicated, extraction level

design and performance are influenced by the degree of fragmentation achieved and by the undercutting strategy adopted. The three-dimensional geometries of the excavations between the extraction and undercut levels can be very complex. They must be designed to ensure that they remain stable and conducive to efficient production operations throughout their design lives which can be several years, or even decades in some cases. As in the example shown in Figure 1.9, the early block caving mines used gravity loading systems via grizzlies and a combination of finger and transfer raises to the haulage level. This system is best suited to ore that fragments finely but is labour intensive requiring significant development. In other cases, including some of the South African diamond mines, a slusher system was used to transfer the ore from the drawpoints to the haulage. A good account of gravity and slusher draw systems is given by Pillar (1981) and summarised by Brady and Brown (1993). An example of a slusher draw system is shown in Figure 6.1. Although grizzly and slusher systems still find some use, they have been almost completely replaced by mechanised methods of drawing and moving the ore on the extraction level using Load-Haul-Dump (LHD) vehicles. Accordingly, extraction level layouts for only mechanised methods of loading will be considered here. The purpose of this chapter is to discuss the factors influencing extraction level design and performance, the advantages and disadvantages of a number of types of extraction level layout, the design of drawbells and drawpoints and the stability, support and reinforcement of extraction level installations. Consideration will be restricted to horizontal extraction level layouts. Discussions of a range of inclined drawpoint layouts are given by Carew (1992), Jakubec (1992), Laubscher and Esterhuizen (1994) and Laubscher (2000).

T

Chapter 6: Extraction Level Design

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Figure 6.1: Slusher draw system (after Pillar 1981) 6.2 FACTORS INFLUENCING EXTRACTION LEVEL DESIGN AND PERFORMANCE

In order to introduce what can be a complex topic, the major factors influencing extraction level design and performance will be listed and discussed briefly before more detailed consideration is given to some of them later in this chapter. Discussions of many of these factors are given by Esterhuizen and Laubscher (1992) and Laubscher (1994, 2000).


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Fragmentation The fragmentation of the ore produced in the draw column influences the choice of draw system used. Gravity systems require finely fragmenting ores while LHD systems are the natural choice for the more coarsely fragmenting ores commonly encountered in modern block and panel caving mines. The degree of fragmentation determines the size of the draw zone and hence the drawpoint spacing. It also influences the height of the drawpoint, the need for access for secondary breaking, the shape of the major apex, the LHD size and crushing requirements. Undercut Strategy and Design

As has been discussed in Chapter 5, the undercutting strategy adopted (post-, pre- or advance undercut) influences the stresses induced in extraction level excavations, the need for support and reinforcement, the rate at which drawpoints can be brought into production and their long-term performance. The design and performance of the extraction level excavations are also influenced by the detailed design of the undercut as discussed in Section 5.4. For example, the undercut shape can influence the tendency for ore to stack and induce excessive loads on the pillars between the drawpoints (see Figure 5.11). Geotechnical conditions

Because the percentage of excavation on and above extraction levels is so high, and the performance of the excavations is critical to the continuity and efficiency of production, it follows that the prevailing geotechnical conditions have major influences on extraction level design and performance. Depending on the undercut strategy and design adopted, the stresses induced in the excavations can be expected to be high and to change throughout the history of development, undercutting and production. The geotechnical characteristics of the rock masses discussed in Chapter 2 (eg major and minor discontinuities, intact and rock mass strengths) and their relation to the in situ and induced stresses will be major factors to be taken into account in the design. They can influence the sizes, shapes and the need for support and reinforcement and repair of the excavations. These issues will be discussed in more detail in Section 6.5. Operational factors

The extraction level layout must be conducive to the efficient removal of the broken ore. The ease and speed of development must also be taken into account in design. A complex layout may require twice as much time to develop as a simple layout (Esterhuizen and Laubscher 1992). The sizes, shapes and geometrical relationships of the extraction level excavations must allow ease of access of LHDs to drawpoints, loading, reversing, turning into and out of drifts, travel to unloading points and unloading. Access of men and machines to deal with hangups in the drawbells or stacking on the major apex must also be considered.


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Major operational hazards

The major operational hazards to be discussed in Chapter 10 – major rock falls, rock bursts, mud rushes, air blasts and water and slurry inflows – can all have significant impacts on the extraction level. The prevention of their occurrence is the preferred way of alleviating the effects of these hazards. In some cases, the provision of appropriately designed support and reinforcement can alleviate the impact of rock falls, rock bursts and air blasts on extraction level installations. These issues will not be considered further here but will be discussed in Chapter 10. Drawpoint brows

As in some other mining methods the stability, reinforcement, wear and repair of drawpoint brows can have major influences on the efficiency of production operations in block and panel caving mines. The issues associated with drawpoint brows could be considered to form part of the other design and geotechnical considerations outlined above but they are considered to be so important that they have been listed here separately. Major influences on brow stability and performance are brow orientation, the presence and orientation of discontinuities (because the brow shape provides release surfaces not present in other excavations), the impact of stress abutments during undercutting, the nature of the reinforcement used and the timing of its installation, and brow wear or deterioration during production. 6.3 EXTRACTION LEVEL LAYOUTS

6.3.1 Scope

Laubscher (2000) has identified 10 different horizontal LHD layouts as having been used in block caving mines. Given that some are local variants of the major types, only five major types of layout will be considered here. In this context, the layout will be taken to refer to the arrangement of production drifts, cross-cuts or drawpoint drifts and drawpoints on the extraction level. The layouts to be considered here are the continuous trough, herringbone, offset herringbone, Henderson or Z design, and parallelogram or El Teniente layouts, and some of their variations. The issues of ore passes, separate haulage levels and crusher locations are discussed generically in Section 6.3.7. The discussion is based largely on those of Esterhuizen and Laubscher (1992) and Laubscher (2000). 6.3.2 Continuous Trough or Trench Layout

Weiss (1981) describes the continuous trough or trench layout used under high stress conditions at the Austro-American Magnetite Company’s mine in Austria. This layout has also been used at the Shabanie Mine, Zimbabwe (Laubscher 2000), the San Manuel Mine, USA (Stevens et al 1987) and considered at El Teniente (Jofre et al 2000). Figure 6.2 shows the layout used at the


249

Austro-American Magnetite Company’s mine. The longitudinal undercut trough drift was developed and blasted before the associated production drifts and drawpoints on the same level. This system involves no minor apices.

1 = ore, 2 = caved in area, 3 = surrounding rock, 4 = head slot, 5 = slots, 6 = drawing drift,

7 = fan drills, 8 = strike drift, 9 = main hauling drift, 10 = drill holes for dewatering

Figure 6.2: Continuous longitudinal trough layout used at the Austro-American Magnetite Mine, Austria (Weiss 1981)

Weiss (1981) suggests that this layout aided the stability of the major apex under the prevailing high stress conditions, although as Laubscher (2000) points out, there is no lateral restraint to the major apex as is normally provided by the minor apices. The lower percentage of excavation on the extraction level than in other layouts should help avoid some stability problems. Selection of the optimum orientation of the trough with respect to the major horizontal principal stress direction is especially important with this layout. If the undercut trough is elevated above the extraction level then a small minor apex is created. Two-dimensional numerical stress analyses carried out by Jude (1990) for the San Manuel trial showed that an elevated trough resulted in the elimination of a destressed zone which developed in the non-elevated case and so helped maintain a compression arch over the production drift. A disadvantage of the continuous trough layout is that the drawpoints associated with a given production drift use the same draw cone so that problems with the cone can affect production from several drawpoints.


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6.3.3 Herringbone Layout

A herringbone layout with directly opposite drawpoints as illustrated in plan view in Figure 6.3, has been used at King and Shabanie mines, Zimbabwe (Laubscher 2000, Wilson 2000). In this case, the minor apices are in line with each other and at right angles to the major apex. Compared to other layouts of similar dimensions (see Figures 6.4, 6.6 and 6.7, for example), this produces unfavourable stability conditions because of the two acute corners (which will be rounded in practice) opposite each other and the large effective spans created by the opposing cross-cuts (Esterhuizen and Laubscher 1992). Operationally, this layout does not have the advantage that a loaded LHD can back readily into the opposite cross-cut or drawpoint drift for ease of turning. This can have safety implications in mud rush prone mines, for example. For these various reasons, the original herringbone layout has been modified to improve its performance.

Figure 6.3: Typical herringbone layout analysed by Esterhuizen and Laubscher (1992) 6.3.4 Offset Herringbone Layout

In the offset herringbone layout illustrated in Figure 6.4, the drawpoints on opposite sides of a production drift are offset or staggered. This helps improve both the stability conditions and the operational efficiency over those applying in the symmetrical herringbone layout discussed above. This system was used initially at the Henderson Mine, USA, and at the Bell Mine, Canada. It has become the layout most commonly adopted in the newer block and panel caving operations including Northparkes (Duffield 2000), Palabora (Calder et al 2000), Premier BB1E and C-cut (Bartlett and Croll 2000) and Freeport Indonesia’s Deep Ore Zone (Barber et al 2000).


251

Figure 6.4: Typical offset herringbone layout analysed by Esterhuizen and Laubscher

(1992)

A plan of the full extraction level layout adopted at Palabora is shown in Figure 6.5. Calder et al (2000) report that this layout was chosen principally on the basis of LHD manoeuvrability and the potential that it provided to use electric LHDs with trailing cables. With the crushers located on one side of the production area, all of the drawpoints point towards the incoming LHD. In addition, with the return ventilation on the opposite side of the production area to the crushers, any dust from the loaded LHD bucket is coursed away from the LHD driver. It is instructive to note in this regard that the Northparkes E26 Lift 1 block cave used crushers on opposite sides of the production area but that the layout for Lift 2 has been simplified through the use of crushers on one side only (Duffield 2000).

Figure 6.5: Extraction level layout, Palabora Mine, South Africa (Calder et al 2000)


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6.3.5 Henderson or Z Layout

In this design illustrated in Figure 6.6, the opposite drawpoint drifts are in line but inclined to the production drift, and the drawpoints and drawbells are at right angles to the production drift. The minor apices are again in line and at right angles to the major apex but, in this case, the acute corners are diagonally opposite each other. This layout was first used at the Henderson Mine, USA, and so is sometimes referred to as the Henderson layout. Interestingly, it is not being used in the newer parts of the Henderson Mine where the simpler layout shown in Figure 1.16 has been adopted. No other current cases of the use of this layout are known.

Figure 6.6: Typical Henderson layout analysed by Esterhuizen and Laubscher (1992)

6.3.6 El Teniente Layout

In the layout developed at the El Teniente Mine, Chile, illustrated in Figure 6.7 (with a more detailed recent example shown in Figure 1.11), the drawpoint drifts (called zanjas) are developed in straight lines oriented at 60o to the production drifts (or calles) and the major apices. The minor apices are short and inclined to the major apices. Figure 6.7a shows the original layout with square drawbells at right angles to the drawpoint drifts. Figure 1.16 shows a similar, but not identical, layout being used at the Henderson Mine, USA, in which the angle between the production and drawpoint drifts is 56o and the drawpoint is divided into two by a small pillar (Rech et al 2000). Figure 6.7b shows one of a number of revised layouts developed at El Teniente in response to a range of mining conditions (Jofre et al 2000). The drawbell shape shown in Figure 6.7b was developed to increase the interaction between drawpoints and to improve downhole blast design.


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Figure 6.7: (a) Typical El Teniente layout analysed by Esterhuizen and Laubscher (1992), and (b) modified El Teniente layout with decahedral drawbells (Jofre et al

2000) Numerical modelling carried out by Esterhuizen and Laubscher (1992) showed the El Teniente layout to be “stronger” than the comparable herringbone, offset herringbome and Henderson layouts. The minor apex in the El Teniente layout is shorter and more stable than those in the other layouts. The orientation of the layout with respect to the major horizontal principal stress direction has a critical influence on the performance of the El Teniente layout in high stress conditions. The most favourable stress conditions are likely to be achieved when the major principal horizontal stress is oriented parallel to the production drifts. An advantage of the El Teniente and new Henderson layouts is that the LHD can back into the opposite drawpoint drift for turning or for straight-on loading when there is brow wear. The only real disadvantage of the layout is that it is not suitable for electric LHDs using trailing cables (Laubscher 2000). 6.3.7 Ore crushing and transportation

Although ore crushing and transportation are not central to the concerns of this book, they will be discussed briefly here for completeness. As illustrated in Figure 1.9, the early block caving mines generally used finger and transfer raises to transport the broken ore from the grizzly to the haulage level where it was loaded onto a tracked transport system. Underground crushers were not required for many of the weak, finely fragmenting ores for which this mining method was developed. Different approaches are required for medium and coarsely fragmenting ores and for the mechanised loading of ore. A modern example of the Henderson Mine, USA, is summarised in Section 1.3.4 and illustrated in Figures 1.14 and 1.15. Full details are given by Rech et al (2000). In this case, the ore is transported by LHDs to ore passes on the extraction level and then transferred through the ore pass system to the haulage level 194 m below. The ore is then loaded into side-dumping trucks and hauled to the crusher dump on this level. It then passes down to the crusher and, after crushing, is transported out of the mine and to the mill by a very long conveyor system.


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A number of recently developed block and panel caving mines use more compact layouts. As noted in the discussion of the offset herringbone layout in Section 6.3.4, the Northparkes E26 Mine Lift 1 uses crushers on the extraction level on opposite sides of the production area. From the crushers, the ore is transported to surface via a conveyor system and a hoisting shaft as illustrated in the schematic vertical section shown in Figure 6.8. The deeper Lift 2 will use a single crusher on the extraction level and an inclined conveyor system to the existing hoisting shaft (Duffield 2000).

Figure 6.8: Schematic vertical section, Northparkes E26 Mine (Dawson 1995)

Figure 6.5 shows a plan of the extraction level layout for the larger Palabora block cave. Here, four crushers are installed on one side of the production area to optimise LHD travel distance and provide flexibility if a crusher becomes unavailable (Calder et al 2000). The crushed ore is then transported via an inclined conveyor system to the hoisting shaft. The planned block caving operation at Bingham Canyon, USA, will use a different system again. In the two planned block cave areas, LHDs will muck ore from drawpoints to ore passes which will then feed the ore to a wide heavy-duty conveyor belt. The conveyor will transfer the ore to a central crusher. The crusher will discharge onto a series of long inclined conveyors which will deliver the crushed ore to the existing surface stockpile (Carter and Russell 2000). With the increasing sizes and speeds of LHDs, and the introduction of tele-operated and fully automated machines, the quality of extraction level roadways has become of increasing importance in caving mines. As well as helping improve productivity and reduce LHD maintenance costs, good roadways can also help manage problems of water ingress and contribute to the development of a closed support ring in high stress conditions as discussed in


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Section 6.5.2. The quality and performance of roadways are influenced by their design, construction and maintenance, and by the treatment that they receive during production operations. There are differing views in the industry about the value of certain roadway types and designs, the levels of expenditure merited on roadways and the desirable forms and standards of roadway maintenance. Some of these issues are canvassed by Laubscher (2000). While this is not the place to explore these issues in detail, a summary of the advantages and disadvantages a range of roadway types prepared by N J W Bell is reproduced in Table 6.1. 6.4 DRAWPOINT AND DRAWBELL DESIGN

6.4.1 Gravity Flow of Caved Ore

Drawpoint and drawbell design are related to, among other things, the degree of fragmentation of the ore and its flow characteristics. Although the topic has been studied almost since caving methods of mining were introduced (eg Lehman 1916), the gravity flow characteristics of caved ore are still not well understood. Based on these studies and accumulated mining experience, a number of principles and guidelines to drawpoint spacing and design can be offered although a formulaic approach is still not available. What has come to be regarded as the classical approach to describing the gravity flow of ore involving the concept of the flow ellipsoid, was developed and applied initially to sub-level caving by Kvapil (1965, 1992) and Janelid and Kvapil (1966). Although it is known to have some deficiencies (Just 1981, Rustan 2000), this approach will be adopted here for purposes of illustration. A useful summary of this approach has been given by Otuoyne (2000). On the basis of a range of laboratory and field tests, it was postulated that if the ore is contained in a bin or bunker and a bottom outlet is opened, the material that will have been discharged after a given period of time will have all originated from within an approximately ellipsoidal zone known as the ellipsoid of motion, draw or extraction (Figure 6.9). Material between the ellipsoid of motion and a corresponding limit ellipsoid or loosening ellipsoid will have loosened and displaced but will not have reached the discharge point. The material outside the limit ellipsoid will remain stationary. As draw proceeds, an originally horizontal line drawn through the broken material will deflect downwards in the shape of an inverted “cone”. The shape of this draw cone indicates how the largest displacements occur in a central flow channel.


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Table 6.1: Advantages and disadvantages of some extraction level roadway types (Laubscher 2000)

ROADWAY TYPE ADVANTAGES DISADVANTAGES Conventional concrete A good finished surface that can be well

controlled to give a strong floor in intermit contact with the hard rock beneath. Can be reinforced with weld mesh panels and protected from LHD digging by inclusion of rails.

Requires a minimum of seven days curing time with no traffic on it, so the end is not available for other work during this time and the floor should perforce be dug out to solid or the material left on the floor properly compacted before the final concrete is poured. Repairs are very difficult as the concrete has to be dug out and a further seven days waited for the replacement concrete to cure.

LHD interlocking roadway bricks

Readily laid on a compacted floor, which does not have to be dug to solid. No delays to travel. Readily repairable in case of damage e.g. floor lift. Concrete is guaranteed to be of good strength to stand the wear and tear of the LHD travel.

Manpower to install the bricks. The costs of the bricks and their moulds etc, which are very high. Can’t be used in drawpoint loading areas without concrete or steel plate.

Compressed bricks – e.g. G pattern, cubic or other varieties.

Same as the interlocking brick, except easier and cheaper to make.

Same as the interlocking brick, but as laborious if not more so to place and tend to tear out on corners.

Reinforced concrete with rails/RSJ’s

As for concrete floors, however the rails have to be put in place before hand and this increases the costs. However, it does improve the wear characteristics and maintenance. These are particularly recommended for loading areas where the wear and tear of the bucket into the floor, if not done, can be horrendous leading to large holes being dug in the floor.

If footwall heave occurs the whole floor lifts and access into the area can be prevented or huge damage to tyres results.

Roller compacted concrete

If the logistics can be overcome and controls put in place this could well be a very useful material for initial floors and for repairs of existing concrete floors or other floors.

Underground this has proved difficult to install as the controls have to be exceptional to make the concrete to the required strength.

Steel plate in the loading area

These have been successfully used in development and trials are to be conducted at King in the next set of draw points to see if this idea has merit.

Run of mine gravel Easy to lay. High rolling resistance to vehicles. Maintenance costs ongoing and higher. Material must be suitable.


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Figure 6.9: The flow ellipsoid concept of the gravity flow of broken ore (Kvapil 1992) A number of more recent field and laboratory studies have shown that the “ellipsoid” is not always a true ellipsoid (eg Just 1981, Kvapil 1992, Rustan 2000). Its shape is a function of the distribution of particle sizes within the flowing mass and of the width of the discharge opening. The smaller the particle size, the more elongated is the ellipsoid of motion for the same discharge opening width. The upper portion of the ellipsoid of motion tends to be flattened or broadened with respect to a true ellipsoid and be shaped more like an inverted tear drop, particularly for large draw heights and irregular particle sizes. Furthermore, as will be discussed and illustrated in Chapter 7, more recent studies (Gustafsson 1998) have shown that for coarser, angular and gap-graded materials, more irregularly shaped flow patterns may develop. Nevertheless, for ease of calculation and explanation, the original flow ellipsoid theory will be developed here. The shape of a given ellipsoid of motion can be described by its eccentricity

( ) 2/12N

2N

N

1 b - a a

=ε (6.1)

where aN and bN are the major and minor semi-axes of the ellipsoid of motion as shown in Figure 6.9. The composition of the ore and its particle size distribution, mechanical properties and moisture content will all influence the shape of the ellipsoid and its eccentricity. Smaller


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particles will produce a larger eccentricity than larger particles. The eccentricity may also be influenced by the rate of draw with a higher rate of draw producing a smaller draw width and therefore higher eccentricity (Otuonye 2000). It is assumed that all horizontal cross-sections of the ellipsoid are circular although this has now also been questioned with an elliptical cross-section being suggested (Laubscher 2000). Janelid and Kvapil (1966) suggested that, in practice, ε varies between 0.90 and 0.98 with values in the range 0.92 to 0.96 being most common. If EN is the volume of material discharged from an ellipsoid of motion of known height, hN, then the corresponding value of the semi-minor axis of the ellipsoid can be calculated as

2/1

N

NN ⎟⎟

⎠

⎞⎜⎜⎝

⎛=

2.094hE

b (6.2)

or

( ) 2/1NN 1

22 -

h b ε= (6.3)

For this ellipsoid there will be a corresponding limit ellipsoid of volume EG, beyond which the material remains stationary. The material between the boundaries of the two ellipsoids will loosen and displace but will not report to the discharge point. Janelid and Kvapil (1966) represent this loosening factor as

NG

G

E - EE

=β (6.4)

They found that β varies between 1.066 and 1.100, but that for most broken ores, β tends towards the lower end of the range so that EG ≈ 15 EN (6.5) Assuming that the limit ellipsoid has the same eccentricity as the ellipsoid of motion, equations 6.2, 6.3 and 6.5 can be used to calculate its height as hG ≈ 2.5 hN (6.6)


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As material is discharged progressively, the sizes of the ellipsoid of motion and of the corresponding limit ellipsoid, continue to grow. A useful dimension in the design of caving layouts is the radius of the limit ellipsoid at the height hN (Figure 6.9),

( )( )[ ] 2/1N 1r 2

NG h - hh ε−= (6.7)

In practice, for the range of assumptions made and equations 6.1 to 6.6 given by Janelid and Kvapil (1966), r ≈ bG as shown in Figure 6.9. Otuonye (2000) gives a further useful equation that may be used to calculate the volume of the limit ellipsoid, EG, from a measurement of the ellipsoid’s width, u, at any height, h:

⎥⎥⎥

⎦

⎤

⎢⎢⎢

⎣

⎡

⎟⎟⎠

⎞⎜⎜⎝

⎛⎟⎟⎠

⎞⎜⎜⎝

⎛⎟⎟⎠

⎞⎜⎜⎝

⎛ +π=

21

2G

221

2G

2

2G

222GG b4

1b4

14

2311 w - u - -

b w u - - hb E (6.8)

where w is the width of the bottom opening and the value of the semi-minor axis of the limit ellipsoid, bG, is known or estimated. It must be pointed out that the approximate relationships and parameters involved in this approach were established for sublevel caving and are now quite old. They require validation for the coarser fragmenting ores now being mined by block and panel caving methods. Some authors postulate cylindrical rather than ellipsoidal draw zones (eg McCormick 1968) but the design principles remain much the same in the two cases (eg Hustrulid 2000). For the very high draw columns now being used or proposed for some operations, the differences between cylindrical and ellipsoidal draw zones are minimal. 6.4.2 Drawpoint Spacing

Three different definitions of drawpoint spacing may be used. The first of these spacings is that between the draw zones within a drawbell worked from two sides as in the typical layouts illustrated in Figures 6.3, 6.4, 6.6 and 6.7. The second is the centre-to-centre spacing of drawpoints in a direction parallel to the production drift axes across the minor apex as in the cases of the 15 m spacings shown in Figures 6.3, 6.4 and 6.6. Unless otherwise indicated, it is this spacing that will be referred to in this section. A third spacing that must also be considered in design is that across the major apex, referred to frequently by Laubscher (1994, 2000). In the herringbone layout shown in Figure 6.3, this spacing is approximately 20 m. In the offset herringbone layout shown in Figures 6.4, the drawpoint spacing “on the diagonal” across the major apex is more like 22 m. Establishing the “correct” spacing of the drawpoints requires a careful analysis of the interactions of several factors including the fragmentation and flow characteristics of the ore


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(including the changes that occur with continuing draw), the method and planned rate of draw, geotechnical and design factors influencing the strengths of pillars, and production and cost considerations including the numbers of drawpoints required to achieve the required levels of production. Obviously, as the drawpoint spacing increases, the number of drawpoints that must be developed and the cost of development will decrease. Increasing drawpoint spacing will also increase the sizes of the pillars between drawpoints and allow larger drifts and equipment to be used in an attempt to increase the efficiency of production. However, the drawpoint spacing is also a function of fragmentation. The ellipsoid of draw concept can provide the basis for the selection of an initial spacing. The application of this approach requires a knowledge of the shapes and dimensions of the ellipsoid of draw and/or the limit ellipsoid. This knowledge may be obtained from model experiments (eg Heslop and Laubscher 1981) or from measurements made in full-scale trials or operations (eg Alvial 1992, Gustafsson 1998, Laubscher 2000). Figures 6.10 and 6.11 taken from an account given by Richardson (1981), illustrate the principles involved for the case of a draw zone with a circular cross-section. If the draw zone has a cross-section of another shape, possibly an ellipse, similar but slightly more complex considerations will apply (Laubscher 2000). In the case shown in Figure 6.10a, the draw zones for adjacent drawpoints do not overlap, isolated draw zones develop and pillars of undrawn ore are left between the drawpoints. This results in the potential loss of ore and in the application of excess weight to the major apex with the associated potential for damage to the extraction level excavations. On the other hand, in the case shown in Figure 6.10b there is a significant overlap between the draw zones. No ore is lost or additional loads imposed, but when the draw zone intersects the waste overlying the orebody, there is potential for waste to be pulled down between the drawpoints resulting in dilution. This effect can be exacerbated if the waste comminutes readily and can flow more easily than the ore (Richardson 1981).

(a) (b)

Figure 6.10: Idealized vertical section showing (a) excessive drawpoint spacing with non-overlapping draw zones, and (b) close spacing with overlapping draw zones ( after

Richardson 1981)


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The conclusion to be reached from these simple considerations is that the drawpoint spacing should be such that the draw zones just overlap. Thus, as a first approximation, the drawpoint spacing should be slightly less than twice the value of the semi-minor axis of the limit ellipsoid, bG, shown in Figure 6.9. However, in block and panel caving mines, the drawpoint width, w, or more correctly, the active drawpoint width, wa, will have a finite value which should be taken into account. This is done by taking the drawpoint spacing required for adjacent draw zones to just overlap as S = wa + 2 bG (6.9) A further consideration, that of the layout of the drawpoints in plan, is illustrated in Figure 6.11. If the cross-sectional area of the draw zone is circular, the best “theoretical” arrangement is for the drawpoints to be placed on a hexagonal pattern as shown in Figure 6.11a as this minimises the amount of undrawn ore left between draw zones. Figure 6.11b shows a square pattern which produces comparatively greater zones of undrawn ore than a hexagonal arrangement with the same spacing. Reducing the drawpoint spacing to reduce these “dead” zones, has other undesirable consequences in terms of increased development costs, reduced pillar sizes and operational inefficiencies.

(a) (b)

Figure 6.11: Idealised cross-section showing (a) hexagonal, and (b) square drawpoint

spacings (after Richardson 1981)

Figure 6.12 illustrates the application of this simple approach to the four main types of extraction level layout discussed in Section 6.3. For purposes of comparison, it is assumed that the draw zones have a constant diameter of about 16 m throughout so that adjacent draw zones just overlap across the minor apices for each type of layout. Clearly, the pair of draw zones generated by the two drawpoints served by the same drawbell overlap to a significant extent. In addition to the advantages and disadvantages discussed in Section 6.3, Figure 6.12 shows that the El Teniente and, to a lesser extent, the offset herringbone layouts have advantages over the other layouts in that they leave smaller “dead” zones between draw zones across the major


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apices. However, in practice the situation is not as simple as that shown in Figure 6.12. Figure 6.13 shows the progressive development of draw zones in an offset herringbone layout as draw progresses from an isolated drawpoint, to two drawpoints serving a drawbell, to adjacent drawbells and then to the next line of drawbells.

Figure 6.12: Idealised cross-section showing draw zones for four major types of

extraction level layout Flores (1993) describes the application of this approach as presented by Kvapil (1992) to the development of a revised mining plan for El Teniente 4 South using mechanised panel caving. In this case, experience with existing layouts could be used to obtain the parameters required to calculate drawpoint spacings. For example, Flores (1993) was able to estimate the value of the eccentricity as 0.96 for both the draw and limit ellipsoids. Since the height of the panel concerned was 260 m, the height or major axis of the ellipsoid of draw or extraction ellipsoid, hN, was estimated from earlier measurements to be one quarter of the panel height or 65 m so that the major semi-axis, aN, was 32.5 m. From equation 6.6 the height of the limit ellipsoid or ellipsoid of loosening, hG, is 2.5 hN or 163 m and its semi-major axis, aG, is 81.5 m. Given that ε = 0.96, and aN = 32.5 m, the value of bN can be calculated from equation 6.1 as 9.25 m. Thus, the diameter of the extraction, dE or the drawpoint spacing for touching draw zones, is given by equation 6.9 as 21.5 m with the effective width of the drawpoint (or the undercut in the specific case analysed by Flores), wa, taken as 3 m.


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Figure 6.13: Progressive development of draw zone interaction in adjacent lines of

drawbells (Laubscher 2000)


264

Figure 6.14 shows the draw or extraction and limit ellipsoids superimposed on the undercut layout being used before it was revised. There is no overlap of the extraction ellipsoids which means that passive pillars are left between the undercut drifts imposing high loads on the extraction level development below. In the design proposed by Flores (1993), the initial geometry was revised to that shown in Figure 6.15.

Figure 6.14: Gravity flow parameters and undercut geometry, El Teniente 4 South (Flores 1993)


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Figure 6.15: Gravity flow parameters and proposed revised undercut geometry, El Teniente 4 South (Flores 1993)

In the preceding paragraphs, the classical ellipsoid of draw and limit ellipsoid concepts were presented and applied to the determination of centre-to-centre drawpoint spacings. However, as will be discussed in Chapter 7 and illustrated in Figures 7.1 and 7.2, the flow of broken material during draw can be more complex than assumed in this simple model. In particular, the shapes and sizes of the “ellipsoids” are greatly influenced by the shapes, sizes and grading of the fragmented particles. In many cases, there may be inter-mixing of material between the draw zones of adjacent drawpoints even when the theoretical ellipsoidal draw zones do not overlap in the manner illustrated in Figures 6.10 and 6.11. The diameter of the draw zone for an isolated drawpoint, or the isolated draw zone (IDZ) diameter, D, is a function of the size distribution of the broken rock. A common rule of thumb used in layout design is that interaction can generally occur if the drawpoint spacing is less than 1.5 D (Laubscher 2000). Figure 6.13 illustrates how interaction can be developed progressively in a group of drawpoints.


266

Drawpoint spacings of 15 m as illustrated in the examples shown in Figures 6.3. 6.4, 6.6 and 6.7 have been commonly used in modern block and panel caving mines. However, with some of the newer mines and the newer parts of established mines operating in more coarsely fragmenting ores, drawpoint spacings have increased. For example, the spacing at Northparkes E26 Lift 2 will be 18 m (Duffield 2000) and that at Palabora 17 m (Calder et al 2000). As was noted at the beginning of this Section, in practice drawpoint centre to centre spacings are not always on a square plan. There is usually a larger spacing across the major apex than across the minor apex. Figure 6.19 to be introduced below shows examples of spacings of from 14 m x 15 m to 15 m x 20 m used in the different sectors of the El Teniente mine. 6.4.3 Drawpoint Size, Shape and Orientation

From a stability viewpoint, the drawpoint and the drawpoint drift should be as small as possible. In practice, drawpoint size will be determined by the particle sizes of the broken ore and the sizes and operational requirements of the loading equipment which, in turn, should be determined by the characteristics of the ore. The blocking of drawpoints by oversize blocks should be avoided but cannot always be guaranteed. The percentage of broken ore reporting as blocks of more than 2 m3 in volume is often used as an indication of the propensity of the orebody to produce large fragments (Laubscher 2000). (This size is equivalent to a cubic block having a side of 1.26 m.) It has been suggested on the basis of model experiments and field experience that the drawpoint size should be three to six times that of the largest fragment (Kvapil 1965, Oyuonye 2000). In practice, modern drawpoint drifts in mechanised layouts in the stronger orebodies are typically 4 m or more in both width and height with the effective sizes of the drawpoints themselves being smaller. Drawpoint drifts and drawpoints are usually of the typical shapes of drifts in metalliferous mining with rounded top corners and possibly roofs. In most circumstances, these curved shapes are more stable than flat roofs especially as spans increase. They are also easier to construct using drill and blasting techniques. Furthermore, the shotcrete lining and steel set support sometimes used at the drawpoint are more conducive to the curved shape. However, as illustrated in Figure 6.16, drawpoints have been constructed with flat roof sections, most notably in what is sometimes known as the Henderson design. The orientation of the drawpoint brow with respect to the major and minor discontinuities present in the rock mass is an important factor in determining the stability of the brow. Figure 6.17 illustrates the point for an idealised case in an offset herringbone layout in a rock mass containing one or more sets of relatively steeply dipping discontinuities striking in a preferred direction. For the relative orientations of discontinuity strike and the excavations shown on the left hand side of Figure 6.17 the discontinuities strike perpendicular to the plane of the drawpoint brow. The horizontal or sub-horizontal stresses induced across the drawpoint brow will tend to clamp the discontinuities in place and minimise the amount of slip and block fall-out that is possible. In conventional rock engineering analyses (eg Hoek and Brown 1980), this


267

orientation would be regarded as unfavourable in terms of production drift stability although this is not likely to be the critical design issue in this case.

Figure 6.16: Concreted drawpoint with flat roof (Butcher 2000b) On the other hand, the orientations shown on the right hand side of Figure 6.17 are more favourable for production drift stability but are relatively unfavourable for drawpoint brow stability. The discontinuities strike parallel to the plane of the brow and parallel to the stresses induced across the brow. This means that they will not be clamped by the stresses and that much larger blocks will be free to slide or fall from the brow than in the former case.


268

Figure 6.17: Illustration of the influence of discontinuity and excavation orientation on drawpoint brow stability (after Laubscher 2000)

6.4.4 Drawbell Geometry

Laubscher (2000) points out that the term drawbell is descriptive in that the ideal shape of of a drawbell is presumably like an inverted bell in order to obtain the best possible flow of ore to the drawpoint. Improved ore flow is often reported with shaped drawbells. However, there must be a compromise between the shape of the drawbell and the strengths of the major and minor apices which must remain stable throughout the life of the drawbell and drawpoint. As well as the two primary factors of ore flow and pillar strength, the drawbell shape will also be influenced by the undercut design and the practical drilling and blasting issues discussed in Section 5.6.5. Figure 6.18 shows sections through the drawbells across the major apex for three possible designs. The three cases illustrated show how inclining the drawbell sides should produce better flow characteristics than vertical sides and a large, flat top to the major apex. In the design illustrated in Figure 6.18a, poor interaction of the draw zones is obtained and production is likely to be lost. The issue of ore stacking discussed in Section 5.4.3 will also arise. In the second case illustrated in Figure 6.18b, the sides of the drawbell are inclined with the other design parameters remaining constant. The interaction obtained between draw zones should be improved over that in the first case and the extraction of ore should be more efficient. There is, however, a decrease in the size of the pillar. A further improvement to the flow characteristics may be achieved by using an inclined undercut with an increased slope and a corresponding increase the overall drawbell height as illustrated in Figure 6.18c. The strength of the pillar is further reduced in this case and would have to be evaluated before such a design was adopted.


269

Figure 6.18: Influence of drawbell shape on ore flow and pillar shape (after Laubscher

2000) As illustrated in Figures 5.28 and 5.29, the shapes of drawbells in plan or horizontal cross-section can be quite complex in modern designs. In the El Teniente layout as illustrated in Figure 6.7, for example, a number of choices exist for the orientation of the drawbell sides with


270

respect to the production and drawpoint drifts. Figure 6.19 shows a number of designs that have been used in a range of particular circumstances at El Teniente, mainly in the Teniente 4 South sector (Jofre et al 2000).

Figure 6.19: Some drawbell shapes used at El Teniente (Jofre et al 2000)

6.5 SUPPORT AND REINFORCEMENT

6.5.1 Terminology

Because of their vitally important roles in maintaining production and the high and changing stresses imposed on them, much attention has been traditionally paid to the support and reinforcement of extraction level excavations. Before discussing the issues involved and the


271

techniques used, it is important to clarify some aspects of the often confused and loose terminology used in this area. The terms support and reinforcement have been used throughout this book and are defined in the glossary given in Appendix A. The definitions adopted distinguish between support and reinforcement and are due to Windsor and Thompson (1993) and Windsor (1997). Support is the application of a reactive force at the face of the excavation and includes techniques and devices such as timber, fill, steel or concrete sets or liners and shotcrete. Reinforcement on the other hand, is a means of improving the overall rock mass properties from within the rock mass by techniques such as rock bolts, cable bolts and ground anchors. Reinforcing elements may be tensioned or untensioned on installation. In the latter case, they are referred to as dowels. Support or reinforcement may also be described as being either active or passive. Active support or reinforcement imposes a predetermined load to the rock surface at the time of installation. It may take the form of tensioned rock bolts or cables or, in other applications, hydraulic props, powered supports or segmented concrete linings. Passive support or reinforcement is not installed with an applied loading but develops its loads as the rock mass deforms. Passive support or reinforcement may be provided by steel arches, mass concrete or untensioned rock bolts, reinforcing bars or cables. The timing of the application of support and reinforcement can have a major influence on its effectiveness. Of particular interest in the current context is the concept of pre-reinforcement which is the application of reinforcement prior to the creation of the excavation. This applies some constraint to deformation and increases the rock mass strength before the loosening often associated with excavation can occur. On the other hand, post-reinforcement is the application of reinforcement at an appropriate time after the creation of the excavation. As the discussion of undercutting strategies in Section 5.2 should indicate, the post-reinforcement or post-support of extraction level excavations will be too late in many circumstances. An alternative approach to terminology developed for the rock burst conditions encountered in the deep level gold mines of South Africa ( Ortlepp et al 1999, Stacey and Ortlepp 1999, 2000) is useful for some applications in caving mines, particularly where large amounts of deformation and damage, or even rock bursts (see Chapter 10), are induced. Here, the support and reinforcement systems used may be described as being either retainment or containment support. In this terminology, retainment support refers to those elements and systems which act within the rock mass to reinforce it for some distance in from the excavation boundary in the same way as reinforcement in Windsor and Thompson’s terminology. Containment support refers to those elements and systems such as wire mesh, straps, lacing, shotcrete and other sprayed membranes, used to contain the broken rock mass around the excavation.


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6.5.2 Principles

The principles of good support and reinforcement practice have been given in a range of textbooks (eg Brady and Brown 1993, Hoek and Brown 1980, Hoek et al 1995) and other publications (eg Hoek 2001, Windsor 1997, Windsor and Thompson 1993). The principles to be outlined here apply to the provision of support and reinforcement for extraction level excavations which, as has been established in Chapter 5, are subject to high induced stresses and susceptible to rock mass failure. The continuity of production operations on extraction levels is so critical in block and panel caving mines that a conservative approach to the provision of support and reinforcement is often adopted. While this approach can be effective in ensuring that the excavations remain stable throughout their design lives and, in some cases require no repair, it does not represent best practice because it may involve more expenditure on extraction level installations than is absolutely necessary. The general principles of good support and reinforcement practice may be illustrated through ground-support interaction, characteristic line or convergence-confinement concepts as in the well-known diagram shown in Figure 6.20. Although expressed in terms of “support”, these concepts also apply to what is referred to here as reinforcement. The details of this approach have been given elsewhere (eg Brady and Brown 1993, Hoek and Brown 1980, Hoek et al 1995, Hoek 2001) and will not be repeated here. The essential feature of this approach as illustrated in Figure 6.20, is that it shows how the support or reinforcement helps mobilise and conserve the inherent strength of the rock mass surrounding the excavation even when it is in a yielded or broken state. The support or reinforcement helps the rock support itself. This approach also illustrates clearly the importance of the timing of installation and the stiffness and yield characteristics of the support and reinforcing elements. A practical example of the use of this approach in establishing extraction level support requirements given by Lorig (2000) and Leach et al (2000) will be summarised in Section 6.5.6.

Figure 6.20: Simplified ground-support interaction diagram for a circular tunnel

excavated in a hydrostatic stress field (Hoek 2003)


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Based on considerations of ground-support interaction mechanics, Brady and Brown (1993) developed a set of general principles for good support and reinforcement practice. These principles were not meant to apply to the case of providing support for the self-weight of an individual block of rock, but to the more general case in which yield of the rock mass surrounding the excavation may be expected to occur. They do not refer specifically to the provision of containment support which will always be required in the application being considered here. These principles as adapted by Brown (1999) are reproduced here with further modification and expansion as necessary to take into account the special features of extraction level installations in block and panel caving mines. It is interesting to note that despite their differing origins, these principles have many points in common with the guidance offered by Laubscher (2000). In this list of principles, the word support is used throughout in the interests of simplicity but should be taken to refer to both support and reinforcement as defined here. 1. Install the support close to the face soon after excavation. In some cases, it is possible and

advisable to install the support (or, as is more likely, the reinforcement) before excavation or before the excavation is complete. In others, usually involving high “squeezing” pressures, it may be advisable to permit some displacement to occur before the support is installed.

2. There should be good contact between the rock mass and the support. If this is not

achieved, it has the effect of reducing the effective stiffness of the support in which case excessive displacement and loss of rock mass strength may occur.

3. The deformability of the support should be such that it can conform to and accommodate

the displacements of the rock mass. This means that in high stress and dynamic loading environments, the support should be capable of yielding and retaining load carrying capacity.

4. Ideally, the support system should help prevent deterioration of the mechanical properties

of the rock mass with time due to weathering, repeated loading or wear. The application of this principle to the support and reinforcement of drawpoints will be illustrated and discussed in Section 6.5.5 below.

5. Repeated removal and replacement of support elements should be avoided. 6. The support and reinforcement system should be easily adaptable to changing rock mass

conditions and excavation cross-section. This principle does not assume the importance in block and panel caving extraction level design that it does in other mining methods because the rock mass conditions will generally be reasonably constant so that the same design can be used for a given block or panel.


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7. The support and reinforcing system should provide minimum obstruction to the excavations and the working face. In the case of caving mines this principle can be extended to account for the need for the support and reinforcement system, and particularly the containment support system, to be able to withstand impact and abrasion by loading equipment and blocks of rock.

8. The rock mass surrounding the excavation should be disturbed as little as possible during

the drill and blast excavation process so as to conserve its inherent strength. A further ground support principle that was not included in Brady and Brown’s list but is well known in civil engineering tunnelling practice (eg Brown 1981, Kovari 2001), is that under high stress conditions, support and reinforcement performance can be improved by “closing the ring” of shotcrete or a concrete lining across the floor of an excavation. In modern mechanised block and panel caving mines, providing and maintaining a good concrete floor on the extraction level is important in controlling costs and maintaining the efficiency of the extraction operations. This brings with it the opportunity to “close the ring” by making the concrete floor integral with the overall support and reinforcement system as illustrated in Figure 6.16. 6.5.3 Support and Reinforcement Elements

The support and reinforcement elements and systems used on the extraction levels of block and panel caving mines will be familiar to most readers of this book. They will be listed and briefly described here for completeness. Fuller details of most of them are given by Hoek et al (1995), Laubscher (2000), Ortlepp (1997) and Wilson (2000). Rock bolts consist of an anchorage, a shank, a face plate, a tightening nut and, in some cases, a deformable bearing plate. By definition rock bolts are tensioned and they may be cement or resin grouted along their lengths. In the applications being considered here, the practical lengths of rock bolts are usually restricted by space limitations in the excavations from which they are installed. Dowels, usually consisting of grouted reinforcing bar, have been widely used for reinforcing extraction level excavations as in the example shown in Figure 6.26. Cable bolts can be longer than rock bolts and can sustain much greater axial loads. For the best results they should be grouted and tensioned, although untensioned grouted cables have also been used. Cable bolts are especially useful for reinforcing the large spans created at excavation intersections as illustrated in Figure 6.31 below. Wire mesh usually consisting of 4 mm diameter wire either welded on a 75 or 100 mm square pattern or in the form of chain link mesh, is used to contain broken surface fragments and as reinforcing with shotcrete.


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Shotcrete or sprayed concrete is used to provide passive support to the rock surface and forms an integral part of most support and reinforcement systems used in modern mines. It may be applied in several layers of tens of centimetres in thickness using either a wet-mix or a dry-mix process. The former process is the most common in modern underground mining. Shotcrete may be reinforced with wire mesh of with steel or other fibres to improve its toughness, durability, shock resistance, and shear and flexural strengths. Sprayed membranes of other types are being used increasingly in metalliferous mining instead of mesh and for other special purposes (Spearing and Champa 2000). Because of the high stresses and the need for absolute security in extraction level excavations, it is considered unlikely that they will find much use in block and panel caving, at least until their mechanical properties have been improved considerably. Mass concrete has long been used as a support element in block caving mines (eg Butcher 2000c, Gallagher and Loftus 1960). In the more traditional block caves in weak rocks, mass concrete could be expected to be stronger than the rock mass that it replaced. This cannot always be expected to be the case in modern mines in stronger rock masses. In these cases, mobilising and conserving the inherent strength of the rock mass using other support and reinforcement systems and the principles outlined in Section 6.5.2 is to be preferred. Straps and lacing made from flat steel plate, tendons or mine rope have been widely used to contain the surface rock between rock and cable bolts and the rock at the acute (bull nose) and obtuse (camel back) angles at the intersections of production and drawpoint drifts. Wilson (2000) gives descriptions of a range of these measures used at the King and Shabanie mines, Zimbabwe, including a technique known as rock stapling which uses steel cable or rope. Steel arches were used extensively in the early days of block caving and, as illustrated in Figure 6.16, are still used especially in drawpoints and in conjunction with shotcrete. Experience in block caving and other forms of mining has shown that, when used alone, steel sets are not as effective as might be supposed unless the principles of good support practice outlined in Section 6.5.2 are followed. The Toussaint-Heintzmann yielding arch which includes elements which yield at constant load have been found useful in controlling large deformations in some cases. 6.5.4 Stress – Strength Analyses

As in other forms of underground rock engineering, stress-strength analyses may be used in design to study the likely responses of extraction level excavations. Examples of the two dimensional calculation of the elastic stresses induced on the boundaries of extraction level drifts under a range of scenarios were given in Chapter 5. However, because of the truly three dimensional nature of the problem, the stress paths followed by the rock on the boundaries of excavations through their lives, and the likelihood that plastic deformation of the rock mass will


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occur, the situation in practice is likely to be more complex than that shown in these analyses and more difficult to replicate numerically. Because of the difficulties involved in carrying out the necessary stress-strength analyses in many cases, recourse is often made to precedent practice and practical guidelines in extraction level support and reinforcement design. For example, Trueman et al (2002) describe the use of a method based on the Q system of rock mass classification introduced in Section 2.7.3. A simpler and more preliminary assessment of likely support requirements in a drift may be made using Table 6.2 which shows approximate tunnel support design categories for a range of rock mass strength to in situ stress ratios. The use of even this simple approach requires a knowledge of the in situ stress regime and an estimate of the rock mass strength that is usually made using the Hoek-Brown criterion (Hoek and Brown 1997).

Table 6.2: Approximate support design categories (Hoek 2003) Rock mass strength/

in situ stress Support category

> 0.5 No serious problems anticipated; support usually chosen to deal with

local safety issues; detailed support design studies not necessary

0.3 to 0.5

Routine support design usually adequate; rock bolts or shotcrete or

light steel sets are normally adequate. Relatively simple design

approaches, such as rock support interaction using ground reaction

curves, are normally adequate

0.17 to 0.3 Careful design of support required; reinforcement (rock bolts) with

mesh- or fibre-reinforced shotcrete required

0.05 to 0.17

Serious instability anticipated; very detailed design of support

required; yielding steel sets may be used; numerical analysis is

advisable

<0.05 Practically impossible to maintain tunnel stability

Despite the difficulties involved, it is instructive to carry out three-dimensional stress analyses in some cases. Figures 6.21, 6.22 and 6.23 show the results of a three-dimensional elastic stress analysis of a hypothetical extraction level layout carried out by Wattimena (2003) using the finite difference code FLAC 3D (Itasca 1997b). The vertical separation of the undercut and extraction levels is 15 m as is the centre-line separation of the drawpoint drifts. The results shown in Figures 6.21, 6.22 and 6.23 are for a case in which the horizontal stress parallel to the drawpoint drifts is twice the vertical in situ stress and that parallel to the production drifts (or parallel to the longer axis of the major apex), is three times the vertical stress. The results presented were obtained using the modelling technique described in Section 5.5 in which a


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large-scale problem is first solved in order to obtain boundary stresses for application to a smaller-scale problem. The example considered here is for a post-undercut in which the extraction level and drawbells are developed ahead of undercutting. Figure 6.21 shows the results for an area 45 m ahead of the cave front. Figure 6.22 shows the results for an area located 7.5 m behind the cave front, while Figure 6.23 shows the results for an area under the cave, 45 m behind the cave front. The coloured contours in the figures represent values of the maximum principal (compressive) stresses on the excavation boundaries normalised with respect to the vertical in situ stress.

Figure 6.21: Three-dimensional stress analysis of extraction level excavations 45 m

ahead of the cave front for post-undercutting (Wattimena 2003) Examination of the figures shows that the excavations 45 m ahead of the cave front are subject to significant stress concentrations. The maximum stresses in the vertical pillar at the end of the drawbell, in a zone above the intersection of the production and drawpoint drifts, and near the bottom of the drift walls, are 7 to 8 times the vertical in situ stress. There is a thin skin above the intersection for which the stress concentration reaches 10 - 11. Most of the minor apex has a boundary stress concentration of 2 – 3 with limited areas at 3 – 4. The boundary stress concentration over most of the major apex is 3 – 4 with the stress concentrations being significantly higher around the extraction level excavations.

FLAC3D 2.00

JKMRC

Step 4574 Model Perspective09:59:13 Thu Aug 30 2001

Center: X: 1.900e+001 Y: 2.900e+001 Z: 2.000e+000

Rotation: X: 20.000 Y: 0.000 Z: 0.000

Dist: 2.695e+002 Mag.: 1.8Ang.: 22.500

Plane Origin: X: 0.000e+000 Y: 9.000e+000 Z: 0.000e+000

Plane Normal: X: 0.000e+000 Y: 1.000e+000 Z: 0.000e+000

Block Contour of Min. Prin. Stress Plane: on behind

-1.2000e+001 to -1.1000e+001-1.1000e+001 to -1.0000e+001-1.0000e+001 to -9.0000e+000-9.0000e+000 to -8.0000e+000-8.0000e+000 to -7.0000e+000-7.0000e+000 to -6.0000e+000-6.0000e+000 to -5.0000e+000-5.0000e+000 to -4.0000e+000-4.0000e+000 to -3.0000e+000-3.0000e+000 to -2.0000e+000-2.0000e+000 to -1.0000e+000-1.0000e+000 to 0.0000e+000 0.0000e+000 to 0.0000e+000

Interval = 1.0e+000


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Figure 6.22: Three-dimensional stress analysis of extraction level excavations 7.5 m behind the cave front for post-undercutting (Wattimena 2003)

The maximum stresses in this example occur in the case shown in Figure 6.22 which is 7.5 m behind the cave front. Here, the boundary stress concentrations above the intersection of the production and drawpoint drifts reach 11 – 12. There are significant volumes of rock under the drawpoint drift floors having stress concentrations of 4 – 5, while the boundary stress concentrations at the drawpoint floor-wall junctions are 6 – 7. The major apex has boundary stress concentrations of at least 3 – 4 throughout with significant areas being at 4 – 5 and higher. The boundary stress concentrations in the minor apex are at least 2 – 3 with significant areas at 3 – 4 and higher. Figure 6.23 shows that although the area 45 m behind the cave front is less heavily stressed than the areas in front of and near the cave front, it is still subject to some stress concentration. The extent of the zone of elevated stress above the production and drawpoint drift intersection is reduced but the maximum boundary stress concentration is still 8 – 9. Similarly, there is a reduced zone with a stress concentration of 4 – 5 in the floors of the drawpoint drifts. The stress concentrations over most of the minor and major apices are 1 – 2 and 3 – 4, respectively.

FLAC3D 2.00

JKMRC


Center: X: 1.900e+001 Y: 2.900e+001 Z: 2.000e+000

Rotation: X: 20.000 Y: 0.000 Z: 0.000

Dist: 2.695e+002 Mag.: 1.8Ang.: 22.500





Interval = 1.0e+000


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Figure 6.23: Three-dimensional stress analysis of extraction level excavations 45 m

behind the cave front for post-undercutting (Wattimena 2003) This example illustrates not only the elevated stresses produced in the extraction level excavations ahead of and near the cave front in conventional or post-undercutting, but the effect of the high horizontal in situ stresses on the stresses induced in the major and minor apices, and the fact that the stresses at a given point on the extraction level change as the cave front approaches and passes over the point. 6.5.5 Support and Reinforcement of Drawpoints

As the example of the Bell Mine, Canada, to be introduced in Section 6.5.6 below shows, before the modern understanding of support and reinforcement outlined in Section 6.5.2 was developed, there was a tendency to support drawpoints and brows with passive support systems consisting of mass or reinforced concrete and steel sets. There was also a tendency to used flat roofs and drawpoint brows as illustrated in Figure 6.16. It has now been shown that, in general, it is better to use curved profiles to improve the induced stress distribution (eg Hoek and Brown 1980) and to use active reinforcement, including where possible pre-reinforcement. However, where the rock mass is closely jointed and weak, it may not be possible to reinforce it adequately and so the use of passive, and usually massive, containment support will remain necessary. Figure 6.24 shows an example of a drawpoint that was reinforced successfully by means of untensioned cement grouted reinforcing bars installed during its development. The 3 m long

FLAC3D 2.00

JKMRC


Center: X: 1.900e+001 Y: 2.900e+001 Z: 2.000e+000

Rotation: X: 20.000 Y: 0.000 Z: 0.000

Dist: 2.695e+002 Mag.: 1.8Ang.: 22.500





Interval = 1.0e+000


280

bars were grouted into the rock above the brow, from the drawpoint and from the trough drive, before the final blast of the brow area was carried out. This means that the rock mass was pre-reinforced and kept tightly interlocked throughout the life of the drawpoint (Hoek et al 1995).

Figure 6.24: Reinforcement of a drawpoint (Hoek et al 1995) Hoek et al (1995) also suggest that, in general, attachments such as face plates should not be used on the ends of reinforcement elements exposed in the drawpoint brow area. Face plates, straps and mesh will tend to be ripped off during production and may pull the reinforcement with them. They consider grouted reinforcing bar to be a good choice for drawpoint reinforcement when the rock is hard, strong and massive. When the rock is closely jointed and there is a possibility of inter-block movement during operation of the drawpoint, rebar may be


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too stiff and the use of grouted birdcage (Hutchins et al 1990) or nutcage cables (Hyett et al 1993) should be considered. Drawpoints and their brows are subject to impact and wear during production. They may also be required to sustain secondary rock breakage activities if hangups occur. Welded steel plates may be installed to reduce wear in drawpoints in cases such as that illustrated in Figure 6.16. Some wear is inevitable and can be tolerated. However, if the damage from both wear and induced stresses becomes too great, the drawpoint may have to be rebuilt. This can be a time-consuming, costly (especially in terms of lost production) and difficult task (Laubscher 2000), and is best avoided by the application of the support and reinforcement principles outlined here. 6.5.6 Examples

Bell Mine, Canada

Figure 6.26 shows an example of the support and reinforcement system used on the extraction level, Bell Mine, Canada, when the LHD system illustrated in Figure 6.25 was introduced in 1980. Table 6.3 provides an instructive comparison of the designs for the earlier grizzly method and the LHD method of extraction. In the LHD method, geomechanical evaluation was used and active support (grouted reinforcing bars, cables, mesh and shotcrete) was preferred to the old passive support systems (steel sets and mass concrete).

Figure 6.25: Caving steps and design dimensions for the LHD method, Bell Mine,

Canada (Lacasse and Legast 1981)


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Figure 6.26: Extraction level support and reinforcement for LHD method, Bell Mine, Canada (Lacasse and Legast 1981)


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Table 6.3: Comparison of grizzly and LHD methods of extraction, Bell Mine, Canada (Lacasse and Legast 1981)

Grizzly Method LHD Method

Structural Geology Rock Mass Competence Geometry of the Pillar Orientation of Pillar Major Apex Dimension Drawpoint Spacing Drift Support Drawpoint Support Pillar Support Undercut Retreat Direction Undercut Speed Undercut Position Undercut Blasting

Not considered State of the art evaluation No interaction, Undercut trough Parallel to boundary line 6 x 6 m 7.6 m Passive, square shaped, massive concrete Passive, H beam reinforced concrete None Normal to extraction drift access Adapted to production needs Between pillars Ring blast – No vibration control

Structural mapping and joint survey Geomechanic classification Interacting pillar (major/minor Apex) Largest angle to weakest joint system 12 x 13 m 12.2 m Active, arch shaped, grouted rebars, mesh and shotcrete Active, Drift support plus reinforced shotcrete Grouted cables From weak to competent ground Adapted to rock mass competence Above pillar (major apex) Vibration control with 32 kg max charge per delay

Ground reaction curves, C-cut, Premier Mine, South Africa

In the design of the possible new C-cut area at the Premier Mine, South Africa, outlined in Section 1.3.3 and illustrated in Figure 1.13, the numerical code FLAC3D was used to carry out stress analyses to study the influence of a number of design parameters. As part of these studies, ground reaction curves were calculated. This brief account of that part of the study is taken from a paper by Leach et al (2000) and a report by Lorig (2000). The ground reaction curves were calculated from an elastic-plastic three dimensional model of the extraction level that included the production and drawpoint drifts and the overlying undercut (Figure 6.27). In the model, excavation of the extraction level was simulated by deleting the material representing the drifts and incrementally reducing the pressure on the walls of the drifts from the pre-mining level to zero. At the end of each incremental relaxation, the displacement of the drift wall was recorded. The ground reaction curves shown in Figure 6.28 were generated by plotting the pressure against the wall closure of drift 2 at the points shown in Figure 6.27. Because of the large zone size used in the model, the magnitudes of the deformations are likely to be underestimated. Nevertheless, the model is still able to give a reasonable indication of the support pressures required to control deformation (Lorig 2000).


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Figure 6.27: Plan section through the three dimensional model used to calculate ground reaction curves (Lorig 2000)

Figure 6.28: Ground reaction curves calculated for drift 2 (Lorig 2000)


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The pressure at which drift wall deformations begin to increase rapidly is the point at which severe damage might be expected to be initiated. The computed ground reaction curves fall into three groups defined in terms of the location relative to the undercut front and hence the stress conditions under which they are mined. These three types of behaviour corresponding to the three areas indicated in Figure 6.29 are: • High stress conditions (> 30 MPa) ahead of or directly below the undercut face.

Deformation of the drift walls increases rapidly when the support pressure falls below 6 MPa. This level of support pressure is impractical to supply and any drift development in this zone would require substantial yielding support and reinforcement.

• Moderate stress conditions (20-30 MPa) where production drifts are effectively less than

10 m in plan from the edge of the undercut. Deformation increases rapidly when the support pressure falls below 2-3 MPa. Again, this level of support pressure is probably not feasible with conventional techniques but may be achievable with mass concrete placed once the undercut is advanced.

• Lower stress conditions (<20 MPa) where the production drifts are outside the high

abutment stress zone near the undercut face and are more than 10 m in plan behind any point on the undercut face. The deformation rate increases at support pressures below 0.1 to 0.5 MPa. It is possible to provide support pressures of these levels with a combination of rock bolts, cables and shotcrete.

The conclusion reached from the ground reaction curve calculations is that development should only be carried out in those areas where the stresses due to the undercut are below 20 MPa. In this case, stability of the drifts could be expected to be maintained with conventional support and reinforcement systems.

Figure 6.29: Schematic diagram showing regions of rock behaviour defined on the

basis of the calculated ground reaction curves (after Lorig 2000)


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El Teniente 4 South, Chile

The history of block and panel caving at Codelco-Chile’s massive El Teniente mine was outlined in Section 1.3.2 where it was noted that the mine has experienced some severe rock bursts. The rock burst hazard will be discussed further in Chapter 10. As might be expected in such a large mine having a long history, a wide range of extraction level support and reinforcement designs have been used. For purposes of illustration, the designs used on Teniente 4 South as reported by Flores (1993) will be reproduced here. Figure 6.30 shows a standard 4 m by 4 m drift design using cement grouted reinforcing bar, chain link mesh and shotcrete. Figure 6.31 shows the cable reinforcement design used at three different cross-sections in the production drifts. Figure 6.32 illustrates the support system used for pillar corners at the intersection of production and drawpoint drifts and the steel sets used with cables for drawpoint support and reinforcement.

Figure 6.30: Standard drift support and reinforcement system, El Teniente 4 South,

Chile (Flores 1993)


287

Figure 6.31: Production drift cable reinforcement, El Teniente 4 South, Chile (Flores 1993)


288

Figure 6.32: Pillar and drawpoint support, El Teniente 4 South, Chile (Flores 1993)


289

Henderson Mine, Colorado, USA

This final example is chosen to illustrate some desirable features of mining engineering and geomechanics practice in a panel caving mine rather than to provide details of the support and reinforcement designs used. In all forms of underground mining, it has been found advantageous to implement mine planning, design and production management systems which link geology, resource evaluation, geomechanics data (including rock mass quality, discontinuity and ground fall data), support and reinforcement requirements, and production planning. Such systems have long been computer-based but there is still great value in preparing hard copy overlays to scales that are not readily managed on a computer screen. Rech et al (1992) provide an excellent example of some practical applications of a geostatistical and engineering database to the management of the large scale panel caving operations at the Henderson Mine, Colorado, USA. The applications discussed by Rech et al (1992) include: • interfacing with the draw control system in the modelling of vertical cave advance. (Draw

control will be the subject of the next chapter); • three-dimensional modelling of data on argillisation which has an impact on production

planning; • analysis of ground support requirements based on ore grade; and • modelling of other quantities of geomechanical interest such as RQD and silicification, as

an aid to ground support design. At Henderson, as at a number of other operations, the various systems and engineering processes linked to the database are interactive, and custom-designed plots or plans can be produced at short notice. Figure 6.33 shows a typical overlay of geological structure and ore zones on the drawpoint pillar layout on the 7700 level. It had been found during earlier mining of the 8100 level that there was a high degree of correlation between molybdenum ore zoning and the resulting rock mass quality and support and repair requirements. This is understandable because molybdenum disulphide is a very low shear strength material, and a lubricant, which occurs naturally as deposits along fracture surfaces within the host rock mass. Figure 6.34 shows the support plan prepared for the same area on the basis of information such as that shown in Figure 6.33.


290

Figure 6.33: Overlay of geological structure and ore zones on drawpoint pillar layout,

7700 level, Henderson Mine, USA (Rech et al 1992)

Figure 6.34: Support plan, 7700 level, Henderson Mine, USA (Rech et al 1992)


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El Salvador Mine, Chile

This account of some difficulties experienced on the production level of the El Salvador copper mine, Chile, and their resolution through the use of rock bolts and mesh-reinforced shotcrete, is based on the paper by Van Sint Jan et al (1987) and information provided to the author by Flores and Karzulovic (2002a). This case history provides an excellent example of the application of the modern support and reinforcement principles and practices outlined earlier in this Section. The El Salvador copper orebody which is extracted by block caving, is located in granodioritic porphyries which have been intruded into andesites and rhyolites. The depths from the ground surface to the complex array of excavations between the undercut and haulage levels ranges from 260 to 700 m. The intact rock compressive strength is approximately 50 MPa and the rock mass is intersected by three major joint sets. The most important structural features in the mine are steeply dipping faults and dykes that are parallel to an almost vertical joint set having a strike of 045o, and are usually associated with heavily fractured rock. At the time of concern in the 1980s, the typical practice was to support the 4.0 m wide by 3.4 m high production level drifts by 44 kg/m yielding steel arches on 0.75 m centres, designed largely on the basis of precedent practice. Heavy rectangular steel frames were used at production and drawpoint drift intersections. Delays in placing and fully blocking the steel sets resulted in the loosening and fall of rock, producing overbreak which had to be filled with rock cribbage and wooden lagging. As a result of this process, the radial stiffness of the support system was much smaller than it would have been had the principles of good support and reinforcement practice outlined in Section 6.5.2 been followed. When the steel sets were subjected to large loads and bending moments as cave mining proceeded, they underwent significant deformation and major failures were common (Figure 6.35).

Figure 6.35: Failure of yielding arch support at a drawpoint, El Salvador Mine, Chile (Flores and Karzulovic 2002)


292

Because of these difficulties and the associated cost of the yielding steel arch support, it was decided to investigate the use of a new support and reinforcement system which could control ground displacements and be repaired or augmented easily. It was recognised that the rational selection and design of such a system would have to take account of the high in situ horizontal stresses and would require that the rock mass strength be estimated. A program of geological mapping, in situ stress measurement and rock mass characterisation was the starting point of this process. It was found that discontinuities and planar or linear structures such as the dykes, faults and the associated zones of fractured rock, rather than the intact rock, would exert the major influences on ground behaviour. Analyses of a number of candidate support and reinforcement systems were carried out using the rock-support interaction analysis for axisymmetric excavations and loading developed by Brown et al (1983) using the initial version of the Hoek-Brown strength criterion. Although it was recognised at the outset that the application of this model represented a major simplification of the real problem, it was also recognised that insufficient data were available to justify carrying out a more comprehensive analysis. The results of these analyses suggested that the ground displacements around drifts could be controlled and stability achieved by the use of a system consisting of 25 mm diameter and 2.5 m long grouted rock bolts on 1.0 m centres complemented by 100 mm of mesh-reinforced shotcrete. In order to allow for the effects of the non-hydrostatic stress field, bending moments were estimated on the basis of lining deformation. Since that time, a semi-analytical method of solving the non-hydrostatic loading case has been developed by Carranza-Torres and Fairhurst (1999). On this basis, the flexible lining system was implemented and performed satisfactorily. Van Sint Jan et al (1987) reported that it had resulted in savings in support system costs of in the order of $US1000 per metre of drift compared with the former steel set system. As shown in Figure 6.36, the flexible support system was also most effective in dealing with the complex geometries at intersections and drawpoints where the rock bolts may be installed as pre-reinforcement ahead of excavation.

Figure 6.36: Drawpoint with flexible support and reinforcement system, El Salvador

Mine, Chile (Flores and Karzulovic 2002)

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CHAPTER 7

DRAW CONTROL

7.1 INTRODUCTION

he major geomechanics-related issues discussed in earlier chapters of this book have been concerned with the investigation, planning and design of a potential block or panel caving operation. If the geotechnical investigation, caveability and fragmentation

studies, the development of an undercutting strategy and design, and the extraction level layout and support and reinforcement designs are not adequate, the operation may be put at risk in a number of ways. The analysis of risk in cave mining was introduced in Chapter 1 and will be discussed further in Chapter 11. However, if all of these geomechanics-related issues are addressed adequately in the investigation and design stages, one major source of risk to the operation remains (putting aside for present purposes political risks, broadly defined, and the economic risks associated with commodity prices and financing issues). If the recovery of caved ore from the large number of drawpoints necessarily involved in the operation, is not carried out in a well planned and controlled manner, that is, if adequate draw control strategies and production plans or schedules are not in place, the operation may be susceptible to a further range of risks. These include failure to recover parts of the resource, excessive dilution, non-uniform or arrested cave propagation, major rock falls and associated air blasts (to be discussed in Chapter 10), excessive loading on pillars around the extraction level, and uneven caving to surface including chimneying (to be discussed in Chapter 9). Laubscher (2000) defines draw control as “the practice of controlling the tonnages drawn from individual drawpoints with the object of: • minimising overall dilution and maintaining the planned ore grade sent to the plant; • ensuring maximum ore recovery; • avoiding damaging load concentration on the extraction horizon; and • avoiding the creation of conditions that could lead to air blasts, mud rushes etc.” The ways in which the broken ore and waste flow during draw, exert a major influence on the development and effectiveness of a draw control strategy and operational plan. This topic of

T

Chapter 7: Draw Control

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gravity flow was introduced in Chapter 6 in the context of drawpoint spacing and design and will be discussed further in Section 7.2. Subsequently, Appendix C will present a computational method of simulating particle flow in draw columns and its application to understanding the particle flow and draw processes and the factors influencing draw system design and operation. As Laubscher (2000) points out, the draw control strategy can be expected to change during the life of a block or panel. The largest change will occur after caving has been established by undercutting and the block is ready to go into production. In the undercutting and cave establishment phase, draw must be controlled to influence the shape of the cave back and to keep the distance between the cave back and the caved ore pile small enough to prevent unintended rilling and air blasts should the back collapse. Draw control in these early stages of the life of a block or panel will be discussed in Section 7.3. Laubscher (2000) suggests that by the time caving is established and full-scale production has begun, it will be necessary to have: • investigated the local factors that will influence the caving and draw down process; • calculated the potential tonnages and grades that will be available from each drawpoint; • developed an overall draw control strategy and production plan, including the timing of

initiation of drawing from individual drawpoints and the future rates of draw; • developed an operational method of managing the draw, including a system for recording

and analysing the results obtained; and • developed a method for estimating at any stage the remaining tonnages and grades for

future production planning and scheduling. These issues associated with draw control in the production phase will be addressed in Sections 7.4 and 7.5. After the general principles and approaches have been outlined in Section 7.4, some of the computer-based tools available to assist in the draw control process will be introduced in Section 7.5. It should be emphasised that some of the issues to be discussed in this chapter do not fall within the commonly accepted scope of mining geomechanics, although many of them do have geomechanical origins and influences. Furthermore, draw control was not a subject of research in the International Caving Study Stage I, nor is it a subject of which the author has any direct practical experience. Accordingly, the material presented in this chapter represents a digest of the literature, relying heavily on the papers presented at the three international mass mining conferences held to date (Chitombo 2000, Glen 1992, Stewart 1981) and Laubscher’s Block Cave Manual produced as part of the International Caving Study Stage I (Laubscher 2000).


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7.2 DRAW MECHANISMS

7.2.1 Basic Studies

Despite its importance in block and panel caving, sublevel caving and other underground mass mining methods, the mechanisms of flow and mixing in caved or otherwise broken ore and waste rock under draw are not well understood. As was noted in Section 6.4.1, the topic has been studied since shortly after block caving methods of mining were introduced in the USA (eg Lehman 1916). Historically, the major methods of study have been laboratory model tests using sand, gravel or crushed rock, and the more difficult and less generic field studies using markers of various kinds. Many of the most important studies have been associated specifically with sublevel caving which does not always involve the scale and numbers of drawpoints associated with block and panel caving. Some of the most important studies have been those of Cox (1967), Gustafsson (1998), Heslop and Laubscher (1981), Janelid and Kvapil (1966), Just (1972, 1981), Kvapil (1965, 1992), Mujuru (1995), Peters (1984) and Yenge (1980, 1981). Recent reviews, particularly with respect to sublevel caving have been given by Bull and Page (2000), Otuonye (2000) and Rustan (2000), and with respect to block and panel caving by Laubscher (2000). In addition to these laboratory and field experiments, there have been a number of attempts to study this and similar problems using theoretical methods usually based on stochastic processes (eg Chen, 1997, Gustafsson 1998, Jolley 1968, Litwiniszyn 1964) or continuum mechanics stress-strain analyses involving plasticity (eg Jenike 1966, Pariseau and Pfleider 1968). A brief summary of these approaches is given by Just (1981). The continuum plasticity approach can apply only when the particle size is small and the material may be considered to act as an equivalent continuum. There is the further difficulty of determining a range of required material properties. For this and other reasons, the way forward is considered more likely to involve other approaches. Stochastic methods still have some attraction for use in flow modelling as shown by Gustafsson (1998). However, the most promising method of simulating the flow of broken rock is considered to be the Distinct Element Method developed by Cundall (1971) for the numerical modelling of the mechanical responses of particulate materials. The further development and use of this approach has been reported in the present and other contexts by, for example, Cundall (2001), Cundall and Strack (1979), Lorig et al (1995) and Lorig and Cundall (2000). This approach is currently best represented in the code PFC3D (Particle Flow Code in 3 Dimensions), developed by Itasca (1998a). Essentially, PFC3D models three-dimensional assemblies of particles that may move and interact under the laws of motion and laws relating to the conditions at the particle contacts. PFC3D has been used to model the flow of particles into a drawpoint during caving, but the utility of this approach is limited by the long simulation times required. In an attempt to


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overcome this practical difficulty, as part of the International Caving Study Stage I, Lorig and Cundall (2000) developed the code REBOP (Rapid Emulator Based On PFC3D). REBOP seeks to reproduce the mechanisms observed in PFC3D models in a more efficient (although approximate) manner, so that simulations can be performed quickly, even for problems involving multiple, interacting drawpoints. An introduction to gravity flow modelling using PFC3D and REBOP is given in Appendix C. At the time of writing, the development of REBOP is continuing as part of the International Caving Study Stage II. The expectation is that REBOP simulations will not only improve basic understanding of the complex issue of draw mechanics, but will also prove able to provide replacements for the empirical mixing rules currently used. As a result of these various studies, it is now thought that three basic draw mechanisms operate in caved material – mass flow, granular or gravity flow and void diffusion (Laubscher 2000). Brief descriptions and illustrations of each of these mechanisms will be given in the succeeding sub-sections.

7.2.2 Mass Flow

The concept of mass flow is illustrated in Figures 7.1 and 7.2. Mass flow occurs in the upper portion of an established cave where the subsidence is uniform. Particle trajectories are nearly vertical, converging or diverging according to the shape of the cave, responding to any side pressures and deflecting downwards towards lower density areas above groups of drawpoints being drawn at higher rates than the other drawpoints. Particle flow is generally not influenced by the rates of draw from individual drawpoints, but there may be steps in the subsidence profile if zones of drawpoints are drawn at higher rates than the norm. If individual drawpoints are worked in isolation, the pattern of orderly mass movement and uniform subsidence may be broken by chimneying, piping, funnelling or rat-holing through the caved material. Other than for such occurrences, there is no horizontal or vertical mixing in the mass flow zone and the rates of flow of fine and coarse materials are the same (Laubscher 2000). The mass flow zone is underlain by a zone of interaction and intermixing in which the other two mechanisms of draw may operate.

7.2.3 Granular or Gravity Flow

Kvapil’s classic concept of the gravity flow of granular materials through a single drawpoint was introduced in Chapter 6. The key concepts of the ellipsoid of motion or draw, the limit ellipsoid and particle flow trajectories are illustrated in Figures 6.9 and 7.1a. Although modifications to the elliptical shapes of the draw and limit “ellipsoids” have been found to be necessary by some investigators (eg Just 1981, Kvapil 1992, Rustan 2000), the general concepts appear to hold reasonably well for two-dimensional flow of finely fragmented, well graded and roughly equidimensional material to a single drawpoint.


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However, as illustrated in Figure 7.1b, when large numbers of drawpoints are worked concurrently in a similar material, the draw characteristics can change dramatically. There is a zone of mass flow and uniform subsidence in the upper portion of a high draw column. In the lower portion of the draw column, there is a zone of stress interaction that may induce the lateral migration of material from slowly drawn drawpoints and inter-column areas into the more active draw columns. This mechanism is often referred to as interactive flow. The lateral movement of material helps even out the rate of subsidence in the overlying mass flow zone (Laubscher 2000).

Limit Ellipsoid

Elipsoid of draw

Mass Flow

Interactive zone

Even draw down

Surface Rill

Dilution

Ore

Low pressure

High pressure

Rapid Draw in Centre Height of Interaction Zone

Figure 7.1: Schematic illustration of granular flow paths for (a) an isolated drawpoint,

and (b) several drawpoints worked concurrently (Laubscher 2000)

7.2.4 Void Diffusion

If, as is often the case in block and panel caving operations, at least in the early stages of caving, the material is coarser, more angular and more poorly graded, the classic flow “ellipsoid” may not develop. Rather, a more irregular flow pattern in the form of “fingers” pointing upwards may develop as illustrated in Figure 7.2. The void diffusion mechanism first postulated by Jolley (1968), is one in which voids associated with large, angular particles are formed. These voids may become filled with finer material from above or from one or both sides, or they may collapse under the influence of arching forces of the type illustrated in Figure 1.7 and then reform at successively higher elevations in the draw column. Although such mechanisms may be simulated using particulate numerical models, confirmation of the operation of the void diffusion mechanism in practice has been obtained by Gustafsson (1998) in marker experiments in Swedish sublevel caving operations. Gustafsson (1998) modelled this draw mechanism stochastically using a probability function which determined whether a void


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was filled from above or from one side or the other. By applying a modifying flow factor to the finer and more mobile waste, he was able to reproduce the runs of waste observed in the field marker experiments. As Laubscher (2000) suggests, it is easy to envisage that the void diffusion mechanism demonstrated in sublevel caving, will also operate in block and panel caving. As illustrated in Figure 7.2a, if a drawpoint is operated in isolation, the passage of successive voids above the drawpoint may be expected to produce a zone of less dense, more mobile material, while the surrounding material becomes compacted under the influence of continuing static and some dynamic loading. Chimneying, piping or funnelling may occur when a succession of cavities work their way through the caved material to surface. Once the path is established and filled with mobile fine material, it becomes a preferred flow channel providing a path for the migration of dilution deep into the ore (Laubscher 2000). If, on the other hand, a number of drawpoints are worked concurrently as in Figure 7.2b, the overlapping “fingers” may be expected to produce both vertical and horizontal mixing of the caved material. As the voids diffuse upwards through the caved mass, it is likely that they will become smaller and possibly more stable. At some elevation in the draw column, the effect of a large number of voids from several drawpoints will be to de-stabilise adjacent voids, limiting the opportunities for faster flow of finer material and promoting more closely vertical flow. This will effectively transform the draw mechanism into one of mass flow as shown in Figure 7.2b.

Mass FlowEven draw down

Surface Rill

Dilution Ore

Rapid drawdown of waste

Rat h l

Preferred path

Drawpoint worked in isolation Intermixing Zone

Height of Intermixing Zone

Figure 7.2: Schematic illustration of the void diffusion mechanism for (a) an isolated drawpoint, and (b) several drawpoints operating concurrently (Laubscher 2000)


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7.2.5 Practical Implications

As Laubscher (2000) suggests, the granular or interactive and void diffusion flow models represent the ends of a probable spectrum of caved material behaviour in practice. If the caved material is made up of particles that are predominantly equidimensional and rounded with a well graded range of sizes, the material may be expected to behave in a similar manner to sand in physical model studies (eg Heslop and Laubscher 1981) and follow the interactive model in block caving. The presence of soft or low friction materials will also favour the development of this mechanism. If, on the other hand, the material is gap-graded or has a wide range of sizes and a high proportion of larger angular or slabby blocks, the draw mechanism is more likely to follow the void diffusion model. In practice, many materials will lie between these extremes and exhibit composite draw behaviour. A given orebody is more likely to exhibit a void diffusion mechanism earlier than later in the life of a block or panel. Table 7.1 provides a comparison of some of the features and practical implications of the granular or interactive flow and void diffusion draw mechanisms. On a practical level, hangups in drawpoints are more likely in materials and circumstances that produce a void diffusion mechanism. This means that more secondary breakage will be required and that a greater number of drawpoints will be required to produce the same tonnage than with interactive draw. For void diffusion draw, there may be a temptation for LHDs to work those drawpoints running finer material in preference to those to which larger blocks report or which suffer hangups (Laubscher 1994, 2000). An important consideration in draw control is the height of the interaction zone (HIZ) referred to in Table 7.1. The HIZ is the vertical height of the zone at the base of the cave within which adjacent draw zones interact and produce lateral migration of broken rock. In both draw mechanisms, the HIZ depends on the range of material sizes present and the maximum drawpoint spacing. If the block is worked so that the active drawpoints are more widely spaced than the layout spacing, the HIZ will increase accordingly. In both mechanisms, other things being equal, the higher the interaction zone, the earlier dilution will appear and the more dilution will be drawn eventually. Figure 7.3 demonstrates a manual method of calculating the HIZ developed by Laubscher (1994). There is another version of this method based on fracture frequency of the rock within adjacent draw column (eg Diering 2000) rather than the Rock Mass Rating, RMRL, as in the example given in Figure 7.3. Given that the RMRL incorporates factors other than the fracture frequency which might be expected to influence fragmentation and draw characteristics, the RMRL version of the method will be presented here. The range of RMRL for the in situ rock in the draw column is plotted on the left hand side of the diagram as in the case of examples A and B. The vertical line C is located at the highest RMRL value for the draw column. The curves 1,2,3 and 4 on the left hand diagram apply for particular ranges of RMRL. The intersection of the relevant curve with the vertical line C is projected horizontally to the curve representing the


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draw zone spacing across the major apex in the diagram on the right hand side of Figure 7.3. The HIZ is then read off the horizontal axis on the right hand side.

Table 7.1: Comparison of draw characteristics for granular flow and void diffusion (Laubscher 2000)

Condition Narrow range of material sizes,

multifaceted blocks. Granular and Interactive Flow

Wide range of material sizes, large interlocking blocks Void Diffusion (VD)

Isolated drawpoint Regular shaped ellipsoid of draw; all material is drawn sequentially, with the highest rates of subsidence in the core of the draw column decreasing outwards to the limit ellipsoid.

Voids diffuse upwards and outwards according to a probabilistic function that favours the less dense areas and fine materials. Eventually “ratholes” form as preferred channels for drawing in fine mobile waste, into the ore

Simultaneous draw from adjacent drawpoints

Draw produces low-pressure zones above a drawpoint and high pressures in the surrounding ground. When many drawpoints are worked at the same time, the stresses above each working drawpoint interact; increasing the pressures in inter-drawpoint material are greatly increased. The pressure differences produce lateral migration in the form of plastic deformation from high-pressure zones to the low-pressure live draw columns, effectively widening the draw from each drawpoint.

Voids formed by working of drawpoints or blasting of hang-ups diffuse upwards and outwards, into the inter-drawpoint material. Working adjacent drawpoints at the same time will increase the number of voids in the inter-drawpoint ground. These de-stabilise each other and this effectively widens the draw from each drawpoint. As the voids diffuse upwards they tend to split, become smaller and more numerous and are more likely to disturb the stability of other voids. This favours vertical flow and limits opportunities for fine material to enter cavities. At a particular elevation in the draw column the material above subsides as in mass flow.

Uneven draw and slow or drawpoints stopped for repair

Interaction between stress fields produces higher pressures in the inter- drawpoint material and in the adjacent slower or temporarily stopped draw-points; this promotes lateral migration into live draw columns, evening out the draw. This action occurs below the HIZ. No “catch up” tonnage should be drawn from the slow drawpoint.

Faster drawn drawpoints produce more voids that diffuse upwards and outwards into adjacent drawpoints. This outward diffusion is not compensated for by a similar number of voids diffusing from slower neighbours. This tends to even out the draw rate as stress interaction does in interactive draw, but it favours lateral migration of the fine fraction. “Catch up” tonnage should only be drawn if the ore has a high proportion of large material sizes.

External influences (peripheral structures and lateral pressures) – inclined draw

Inclined draw due to sliding on peripheral structures or lateral pressures that can widen the draw column.

VD is probabilistically controlled and is not influenced by lateral pressures, but may be constrained by peripheral structures. In the mass flow it would be influenced by lateral pressures as in interactive draw


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Figure 7.3: Method of calculating the height of the interaction zone (HIZ) in a draw column (Laubscher 1994)

7.3 DRAW CONTROL DURING UNDERCUTTING AND CAVE INITIATION

The factors influencing the initiation of caving by undercutting and subsequent cave establishment were discussed in Chapter 5. Some of the important factors are: • the starting point for undercutting and the initial direction of undercut advance; • the shape of the undercut in both plan and vertical section; • the geomechanical nature of the orebody, the induced stresses and the associated

mechanics of caving. These factors will influence, among other things, the fragmentation and the flow behaviour of the caved material;

• the undercut area required to initiate caving and the possibility of extending it if necessary, generally in the direction of the minimum span;

• the number of active drawpoints required to meet the production schedule; • the draw control strategy and production plan; and • the further extension of the initially caved area, by either a block or panel caving method. Most of these factors have been discussed in earlier chapters of this book. The purpose of the present Section is to discuss the influence of draw control practice on the establishment and propagation of the cave. It is essential to recognise that there is a relationship between the natural rate of caving and the permissible rate of draw of the caved material. Consider a vertical slice of width, w, in a cave developing in an area under draw as shown in Figure 7.4a. If the slice is drawn down by a


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vertical distance, d, as shown in Figure 7.4b, caving can then occur above the slice until the space created by drawing is filled again as shown in Figure 7.4c. During the process of caving, the ore will increase in volume, or bulk, with an in situ volume, V, becoming a caved volume of V(1+B) where B is the bulking factor. The term swell factor is sometimes used to represent (1+B). Laubscher (1994) suggests swell factors of 1.16 for fine fragmentation, 1.12 for medium fragmentation and 1.08 for coarse fragmentation. It is suggested that, in some cases, the swell factors may, in fact, be greater than these values. If the in situ unit weight of the ore is γ, the overall unit weight of the caved ore will be γ / (1+B).

Figure 7.4: Vertical slice through a draw column in the early stages of draw showing (a) caved ore to the cave back, (b) the formation of an air gap following draw, and (c)

the filling of the void following the next episode of caving.

It will be noted that, in order for the cave to propagate, it is necessary that an air gap be created above the ore pile by drawing caved ore. However, in order to ensure that the cave remains filled after each successive episode of caving and that an excessive air gap does not develop, the rate of draw must be related to the rate of caving and to the bulking factor, B. The volume of the air gap BCEF in Figure 7.4b and the volume of in situ ore BCC′B′ in Figure 7.4c, must together equal the bulked volume of the newly caved ore EFC′B′ in Figure 7.4c. Thus (c + d) w = c (1 + B) w (7.1) or, d = c B (7.2)

This means that for caving to proceed as illustrated in Figure 7.4 so that a permanent air gap is not created, the volume drawn from the slice after each episode of caving should be only the difference between the in situ and bulked volumes of the newly caved ore, sometimes referred to as the swell. In other words, the time rate of volume draw,

•d , should be B times the time


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rate of caving by volume, •c . If the time rate form of equation 7.2 is written in terms of

tonnages using the relationship volume = tonnes / unit weight, we obtain the result

( )B 1B/c d tt +=••

(7.3)

where tc•

and •

td are the time rates of caving and draw in terms of tonnages. The same

considerations apply to the drawing of blasted ore during the undercutting phase. There is sometimes a temptation to draw excessively from the undercut or from drawpoints that come into early production in a new operation or block in order to begin generating cash flow and return on the high initial investment. This may have at least three undesirable consequences: 1. it could lead to the development of an uneven profile in the cave back and the possible

arrest of caving as a result of the new, uneven distribution of stresses induced in the cave back;

2. in weak materials, if draw is continued or if a path for the gravity flow of fines is

established, it could lead to chimneying through the orebody and early dilution as illustrated in Figures 7.1a and 7.2a; and

3. if a group of drawpoints is drawn excessively with respect to the rate of caving, or if cave

propagation is arrested and draw is continued, an air gap may be created providing the potential for a damaging air blast to be generated should a major collapse of the cave back occur. The issues of major collapses, air gaps and air blasts will be discussed in Chapter 10.

In orebodies with no distinct near-vertical geological boundary between the payable ore and the barren country rock, it is especially important to avoid excessive early draw from boundary drawpoints because this could lead to the early introduction of dilution. A further problem occurs at the advancing cave front during the establishment of caving in a block cave and throughout a panel caving operation. As described by deWolfe (1981), this problem was well identified and addressed by draw control practices in the early stages of panel caving at the Henderson Mine, Colorado, USA. This account of the issue is taken from the paper by deWolfe (1981). It also demonstrates the practical application of the principles discussed earlier in this Section. Since the cave front in a panel caving operation is constantly advancing, it is necessary to be continually initiating a new cave at one end of a panel while exhausting or depleting it at the


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other as illustrated in Figure 7.5. Therefore, an inclined ore-waste contact must exist within the caved mass of rock from initiation to exhaustion. There is a second inclined contact between caved and uncaved material, sloping in the obverse direction to the first as shown in Figure 7.5. Draw control procedures should aim to keep the caved ore-waste contact as smooth and as even as possible so as to avoid dilution, particularly near the time of exhaustion. The method of managing this issue at Henderson was to assign to each row of drawpoints maximum permissible draw tonnages expressed as percentages of the available tonnage in the draw columns. As illustrated in Figure 7.5, these percentages were increased in 10 or 15 % increments, working away from the cave line. As the cave front advanced, these tonnages were increased progressively for a given line of drawpoints. However, deWolfe (1981) reports that, at the time concerned, the actual tonnage drawn from each drawpoint was kept to about 50% of the allowable maximum in order to maintain adequate tonnage in the cave to sustain production for several months if caving were to be arrested.

Figure 7.5: Ideal draw control in panel caving, Henderson Mine, USA (deWolfe 1981)

The average rate of draw for the operating drawpoints was 0.3 m per day, with draw assignments ranging from 0.6 m per day for drawpoints that were behind schedule down to a minimum of 0.15 m per day for drawpoints that were ahead of schedule or in the process of being exhausted. It was assumed that, at any time, one third of all active drawpoints would be unavailable due to repair or hangups, and that the two-thirds in operation would have to meet the daily production requirements. The average draw rate of 0.3 m per day was based on the caving rate and was strictly adhered to through draw control procedures. DeWolfe (1981) notes that “a proper draw rate should allow a small void to be formed between the broken ore and the in-place ore allowing caving action by gravity to continue. This void must not be allowed to become so large as to create the possibility of an air blast caused by massive rock falls. Large voids running the length of the cave could also permit waste material to roll along the contact between broken and unbroken


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rock diluting higher-grade ore columns nearer to the cave line.” Figure 7.6 illustrates this possibility which must also be avoided during the development of the cave in a block caving operation. DeWolfe (1981) also refers to the practice of drawing only the swell from undercut blasting.

Figure 7.6: Formation of an air gap and waste rilling when the draw rate exceeds the caving rate swell in panel caving (deWolfe 1981)

7.4 DRAW CONTROL DURING PRODUCTION

7.4.1 Manual Calculation of Draw Tonnages and Estimation of Dilution

In order to provide an easily understood introduction to the factors and steps involved, manual methods of calculating draw tonnages and estimating dilution will be outlined in this sub-section. These calculations may be carried out manually or using computer-based spread-sheets. In practice, the calculations are usually carried out as part of an integrated computerised geological model and mine and production planning system. Two examples of such approaches will be given in Section 7.5. Both manual and computerised methods of calculation use as their starting points geological block models showing the distribution of economic and geotechnical parameters within the orebody. The details of these models lie outside the scope of this book. The calculation methods also involve assumptions about the draw mechanism and require a method of modelling the mixing of caved material. The manual method of calculation involves the following steps (Laubscher 2000): 1. Decide upon the potential direction of draw and define the draw columns. The direction of

draw can be influenced by such factors as geological structures and topography which constrain the subsidence zone (to be discussed in Chapter 9), and the material properties,


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caving mechanism, methods and rates of draw and the stress regime operating on the draw column. In many cases, especially in deep caves, the draw can be expected to be mainly vertical if uniform draw down (see Section 7.4.2) is used.

2. Transform the block model to suit the drawpoint layout and draw inclinations. 3. Divide the block up into like drawpoints and like areas of inclination of draw and grade

distributions to ease the burden of calculations. 4. Calculate the HIZ, for example by using Figure 7.3. 5. Depending on the draw mechanism considered likely to apply, select an appropriate grade

analysis graph from Figures 7.7 to 7.10. It will be noted that, in these figures, the draw column is divided into slices of equal height with the HIZ occurring at the top of slice 7 and the ore-waste boundary occurring at the top of slice 12. The vertical composition axis gives the percentage of ore in the material drawn for a particular slice. The lower graphs in Figures 7.8 to 7.10 represent linearisations of the “theoretical” curves for ease of calculation.

6. Calculate the average grade for increasing stages of draw from the graphs until the grade

drops below the grade cut-off value. The draw column remaining above this point should be tested to determine whether there is a local drop in grade and/or whether there is more economic ore above this point. The grades may be modified using Figure 7.11 to take into account the quality of the draw control achieved.

Figure 7.7: Grade analysis graph for an isolated drawpoint in granular material

(Laubscher 2000)


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Figure 7.8: Grade analysis graph for interactive draw (Laubscher 2000)

Figure 7.9: Grade analysis graph for an isolated drawpoint in void diffusion draw (Laubscher 2000)


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Figure 7.10: Grade analysis graph for void diffusion draw (Laubscher 2000)

Figure 7.11: Dilution curves for varying draw control quality (Laubscher 2000)


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Another useful calculation is of what Laubscher (1994) calls the dilution entry or the percentage of the ore column that has been drawn before dilution appears at the drawpoint. Obviously, the dilution entry is a reflection of the amount of mixing that occurs in the draw column. The percentage dilution entry is calculated as D = (A – HIZ) / A x C x 100 (7.4)

where A = draw column height x swell factor and C = draw control factor. The draw control factor, C, is a measure of the variation in the tonnages drawn from the working drawpoints (presumably those likely to have an influence on the drawpoint concerned.) When Laubscher (1994) used this factor for the first time, he proposed that it be a straight line function of the standard deviation of the tonnages drawn from the group of working drawpoints over a selected period of time and varies from 1.0 when the tonnages are equally distributed down to 0.3 when they are distributed very irregularly (Laubscher 1994). 7.4.2 Draw Control Strategies and Procedures

The following strategies are commonly employed in drawing block or panel caves (Laubscher 2000): Uniform draw down is generally the best option to ensure good recovery and dilution control. Towards the end of the life of a block this strategy may leave drawpoints with initially high draws to be worked in increasing isolation. Height of draw may be used as a scaling factor for draw rates. This is intended to ensure that all drawpoints in an area are depleted at about the same time. However, if there are large differences in draw heights, the drawpoints that have to be worked faster may draw more from their neighbours than allowed for in the simple calculations. This may require the draw columns to be redefined and the tonnages available to be recalculated. High grading may be practised through earlier or faster draw of the higher grade drawpoints. As well as possibly leading to uneven or isolated draw, faster draw may again mean higher rates of migration from neighbouring drawpoints and the need to recalculate the tonnages available for draw. The inclined ore-waste interfaces should be controlled in panel caving and the early stages of block caving as discussed in Section 7.3 and illustrated in Figures 7.5 and 7.6. Having established the draw control strategies and calculated the expected tonnages and grades available from each drawpoint, procedures to measure and control the draw process must then


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be put in place. Here again, these procedures may be manual or, as is now usually the case, they may utilise modern information and communications technologies (eg Knights et al 1996). The draw control procedures should include methods of: • communicating to operating staff the production requirements (or calls) from an area and

from individual drawpoints for the day or shift. It is essential in this regard that all drawpoints and production drifts are clearly numbered;

• measuring and recording the tonnages actually produced from individual drawpoints and areas; and

• reporting and analysing these production figures. In practical terms, it is necessary to ensure that the draw strategies and production schedules can be implemented with the numbers of drawpoints available. Julin (1992) suggests that there should be an 80-90% availability of new drawpoints with this figure decreasing with the period of operation. Theoretically, it may be advantageous to work all drawpoints in a given production area on every shift. However, this may not always be practicable because of the unavailability of drawpoints as a result of hangups or the need for repair, or when the tonnages scheduled to be drawn from individual drawpoints (or calls) are small. Alternatives may be to work half the drawpoints on one day or shift and the other half on the next. This may be achieved by working alternate drawbells along each production drift or alternate production drifts. Those drawpoints not being worked at a given time should be closed off in some way (eg by chains). It is also necessary to ensure that operators do not draw excessively from particular drawpoints in which the fragmentation and/or ease of access may make them easier to work than other drawpoints. Equally, it is necessary to ensure that occurrences which jeopardise the pulling of scheduled drawpoints (eg hangups, remedial support and reinforcement) are kept to a minimum and of minimum duration. A significant number of practical issues and constraints must be taken into account in developing production schedules for areas and individual drawpoints. Guest et al (2000) give a detailed list of these issues and constraints in the context of diamond mining operations. In addition to the unavailability of drawpoints because of repair or hangups just discussed, these factors include: • allowances for undercut material in the early stages of block caving or in panel caving; • the rate at which new drawpoints may be commissioned (from one to, say, six or more per

month); • maximum permissible draw rates for individual drawpoints or groups of drawpoints

resulting from considerations such as those outlined earlier in this chapter; • extraction equipment (usually LHDs) productivity and availability; • the condition, availability and productivity of ore transportation systems, including the

roadways used by LHDs;


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• similarly, the condition, availability and productivity of any crushers and ore passes used in the production process;

• the stage in their lives or the tonnages already extracted from the drawpoints before the start of the production schedule in question. This may affect the grade (including dilution) and size distribution of the material and drawpoint availability;

• any constraints on the grades or the maximum or minimum tonnages required for ore stockpiles or the processing plant in a particular period; and

• more general economic constraints such as the market supply and prices of the commodities being produced. The case of molybdenum production at the Henderson Mine was referred to previously in this regard.

7.5 EXAMPLES OF COMPUTERISED DRAW CONTROL SYSTEMS

7.5.1 PC-BC

PC-BC is a program developed for use in block caving design and operations and now marketed by Gemcom Software International (Gemcom 1999). This brief account of PC-BC is taken from a paper by the program’s developer (Diering 2000). The program is integrated into a general purpose geological modelling and mine planning system so that it can be used for studies ranging from pre-feasibility to daily draw control. It follows a similar process to that outlined for manual calculations and applies empirical rules of the type developed by Laubscher (1994, 2000) to model material mixing. An overall flow chart for PC-BC from initial block models to mineable reserves or daily production schedules is shown in Figure 7.12. As Diering (2000) indicates, PC-BC has been used on a wide range of studies for planned and operating caving mines and has been improved progressively as a result. Input Data

The starting point for PC-BC is the assembly of the required input data about the orebody, the waste rock and the cave design: • A geological block model with block sizes in the range 5-25 m should provide data on

grades, in situ unit weight, a rock code (used to distinguish ore from waste and, in some cases, previously caved material), and the percentage of fines in each block (typically taken to be material that is smaller than about 200 mm).

• A fracture frequency rating which ranges from 0-40 and is obtained as part of the rock mass classification process, is used to calculate the HIZ in the manner described in Section 7.4.1.

• Drawpoint locations may be taken to be at the extraction level and directly below the brow of the drawpoint or they may be on the undercut level. Drawpoints are also grouped into types depending on the method of loading the caved material or other design features.


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• A draw cone is defined for each drawpoint type. The draw cones may be vertical or inclined columns of rock above each drawpoint having radii of from 5 m up to 20-25 m in some special cases. The draw cone radius is chosen so that adjacent draw cones will just overlap at the undercut level. The draw cone is given a slight conical shape with an outwards inclined surface of 1o-3o up to the top of the ore zone, typically 200-300 m above the undercut level. At the top of the cone, the radius is taken to be greater than or equal to that at the top of the ore column.

Figure 7.12: Overall flow chart for PC-BC (Diering 2000)


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Computation of Drawpoint Reserves

As in the manual calculation method, the next step is to compute an inventory of the volumes of rock that potentially could be extracted from each drawpoint. Material may be unique to a drawpoint or it may be potentially shared between drawpoints. The method used to compute block fractions associated with each drawpoint is outlined by Diering (2000). Once these drawpoint fractions have been computed, a drawpoint reserve is constructed for each drawpoint. The draw column above each drawpoint is divided into typically 50 slices having a vertical dimension matching that of the block model. Each slice is then divided into several categories for subsequent modelling and reporting. The material is classified as being shared or unique, coarse or fine, and ore or dilution. For each slice and for each of these categories, data on tonnages, volumes and grades are stored in a slice file as illustrated in Figure 7.13. The effect of the surface topography must be taken into account in determining draw column heights and constructing slice files.

Figure 7.13: Components of the slice file in PC-BC (Diering 2000)

In addition to the basic grade and tonnage data for each drawpoint, a height to volume curve is also constructed from the block model fractions previously computed. This allows for the fact that, because of its larger diameter, a slice of a given height higher in the column will have a greater volume than a slice of the same height lower in the column. At this point, the slice file contains information about the draw column associated with each drawpoint before any mixing has taken place. As shown in Figure 7.12, this represents a major branch point in the program. The basic options are: • pre-compute the vertical mixing for each column before it has been drawn so that heights

of draw and approximate mineable reserves can be computed for overall planning purposes (Option A); or


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• start a mining simulation extracting material from drawpoints sequentially so that the mixing of material may take place progressively (Option B). This better represents the actual caving process and is the option that will be described here.

Material Mixing

The simulation of both vertical and horizontal mixing is an important part of PC-BC. The modelling of vertical mixing initially used the linear form of Laubscher’s grade analysis curves discussed in Section 7.4.1. A number of short-comings were identified in the original approach, as a result of which a new algorithm was developed (Diering 2000). It should be noted that this revised method of modelling vertical mixing still relies to a significant extent on Laubscher’s concepts and methods based on the granular or interactive flow mechanism. The revised vertical mixing algorithm involves the following steps: • The draw column for each drawpoint is divided into typically 50 slices, each 10-15 m

thick. The volume, thickness and the components of the slice shown in Figure 7.13 are known.

• When a slice is extracted from the bottom of the column, one cycle of the mixing process is

initiated and all other slices must move down the column. Below the “Mixing Horizon” at height MH, each slice is mixed with the two slices immediately above it. There is no mixing above MH. The MH value is similar to the HIZ but it relates to a column of broken rock rather than to a column of in situ rock as in the calculation of HIZ.

• The bulking or swell factor for the material determines how far slices may move down in

response to draw. Slices are initially at the in situ unit weight but this value is reduced on caving and bulking as discussed in Section 7.4.1.

• When the slices are being mixed, some material redistribution takes place in the columns to

ensure that slice volumes are reasonably well maintained and that the overall mass balance is maintained.

• The value of MH is calculated by first calculating HIZ as described in Section 7.4.1 using

fracture frequency per metre rather than RMRL as the horizontal axis of the left hand diagram in Figure 7.3. Laubscher’s draw control factor, C, must also be estimated. The height of the mixing horizon is then calculated as

MH = (HU + HIZ/ C) x SF (7.5) where HU = height of the undercut level above the extraction level and SF = swell factor.


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• Different mixing fractions are assigned to fine and coarse components of the material so that the fines move down the draw column quicker than the coarse component.

• The first dilution entry may be calculated using the terms defined in Equations 7.4 and 7.5

as D = (( SF – 1) + ( H – HIZ) / A )) x C x 100 (7.6)

where H = height of the ore.

It should be noted that this equation is based on, but differs from, Equation 7.4 given by Laubscher (1994).

Horizontal Mixing

PC-BC models the horizontal movement of material above the drawpoints up to the mixing horizon. Figure 7.14 shows a schematic cross-section of two draw cones and the zone in which mixing is assumed to occur. In modelling horizontal or cross-drawpoint mixing, PC-BC aims to move material between drawpoints in a non-reversible process. For each drawpoint, the following steps are involved: • Identify the neighbouring drawpoints for the current drawpoint including the one that is in

the same draw cone. • Compute the average tonnage extracted from the neighbours and compare this with the

tonnage extracted from the current drawpoint. • If the ratio of the current drawpoint tonnes to the average for the neighbours exceeds a

user-defined threshold (suggested as 3.0), then isolated draw is assumed with no cross-drawpoint mixing.

• A target tonnage is set for cross-drawpoint movement based on the comparison of the

current drawpoint tonnage and the average for the neighbours. The target tonnage is then adjusted and apportioned taking into account the fact that more movement can be expected to take place between drawpoints in the same drawbell. Only shared material and material below MH is considered.

Diering (2000) makes the point that the modelling of cross-drawpoint mixing in this way is relatively new and that there is not enough detailed information available on draw patterns to calibrate the approach. It is possible that the discrete particle numerical modelling approach to be introduced in Appendix C will be able to advance understanding in this regard.


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Figure 7.14: Schematic cross-section of two draw cones showing the assumed zone

of cross-drawpoint mixing (Diering 2000)

Depletion of Material

As material is extracted from the draw column, material moves down and is mixed vertically and horizontally as discussed above. During this process, use is made of the height to volume curve constructed early in the process to update the slice file inventory and to locate slices reasonably accurately in space. The overall simulated mining process involves a number of steps concerned with the material to be removed over a range of periods (of days, weeks, months or years). After draw and mixing have taken place over some time, the tonnage available to individual drawpoints will become depleted. This is represented in PC-BC by the progressive removal of slices from the slice file as illustrated in Figure 7.15. The process of depleting or exhausting slices at a drawpoint begins by determining whether or not the bottom slice contains enough material to meet the tonnage requirement for the period being considered. If it does, the material may be removed and the simulation continued. If not, all material in the slice must be extracted and the slice depleted. All of the remaining slices will then move down by one and vertical mixing will occur. As the slices move down, they expand from their in situ unit weights to their caved unit weights and the new height of each slice is computed from the height-volume curve for the draw cone. Not all slices move initially because only the swell should be removed from the bottom of the column. As material moves down the column, the proportion of shared and unique material increases. As a result of these various processes, the thicknesses of slices will vary as they move down the column. This is one of the reasons for the difference in the results obtained by modelling vertical mixing before (Option A in Figure 7.12) or during draw down and depletion (Option B being discussed here).


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Figure 7.15: Depletion of slices from the slice file during simulated mining in PC-BC

(Diering 2000)

Production Schedules

The preparation of production schedules involves consideration of the range of factors outlined in Section 7.4.2, not all of which are accounted for explicitly in PC-BC. However, because PC-BC simulates the extraction from each active drawpoint period by period subject to a range of constraints and inputs, it can calculate and display a range of parameters required in production scheduling. These include: • achievable tonnages for each period; • available tonnage and grade estimates for each drawpoint for each period; • periods when each drawpoint will be active or closed; • the number of new drawpoints required to be commissioned during each period; and • the total number of drawpoints available and the active caving area during any period. Diering (2000) lists the following features and capabilities of PC-BC as being relevant to short-term draw control procedures: • Ready storage and retrieval of planned or actual tonnages. A useful feature is the two-way

link with Excel spread-sheets which permits ready manipulation of the data in the production planning process.

• Graphical display of planned or actual tonnages in various formats including two-

dimensional plan view plots and three-dimensional isometric or histogram plots.


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• The ability to produce visual representations of draw columns showing the current state of

draw or depletion and the predicted position of dilution material. • The ability to schedule production from different production blocks to ensure that draw

down is even or accords with maximum allowable daily rates of draw. Diering (2000) also notes that there is scope for further improvement of PC-BC in this area including the incorporation of improved draw control rules and constraints, improved interfaces with supervisory control and data acquisition (SCADA) systems linked to LHDs, and improved database interfaces and storage. 7.5.2 De Beers’ Linear Programming Based System

Guest et al (2000) provide an instructive summary of the historical development of draw control procedures at De Beers’ Kimberley diamond mines. The methods used have developed from no control initially, to manual recording of daily production, to proactive manual control systems, to successive computerised draw control systems and finally to the development of a system based on linear and mixed integer programming. The first version of the program PC-BC described in Section 7.5.1 was used on Premier Mine (Owen and Guest 1994). Figure 7.16 shows the complete draw control system flow chart proposed by Guest et al (2000). The overall system includes a data collection system, a mineral resource database, a cave simulator, a short-term scheduler, and a long-term scheduler which makes use of linear and mixed integer programming. A range of geotechnical, mining and metallurgical constraints such as those listed in Section 7.4.2 are accounted for in this system. Linear programming is a powerful mathematical technique used to optimise (minimise or maximise) a linear objective or utility function subject to linear boundary or constraint functions. It has been used in a wide range of engineering, transportation, economic and mining (eg Graham-Taylor 1992) applications since it was first developed for military applications in the 1940s (Vajda 1961). Where the functions to be optimised or the constraints are non-linear, an extension of linear programming known as mixed integer programming may be used, as in the draw control system described by Guest et al (2000).


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Figure 7.16: A complete draw control system flow chart (Guest et al 2000) Figure 7.17 shows the iterative process used in the linear programming based long-term scheduler developed by Guest et al (2000). The purpose of the present account of this approach is not to explore the sophisticated mathematical and computational techniques involved, but to illustrate the importance and treatment of a number of the factors introduced earlier in this chapter, and some that were not. Indeed, at the time of writing, the system is undergoing further development at the JKMRC as part of the International Caving Study Stage II.


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Figure 7.17: The iterative process within the linear programming based long-term

scheduler (Guest et al 2000)

In their initial examination of the use of linear and mixed integer programming techniques to develop an optimal depletion strategy for a block cave, Guest et al (2000) found that geotechnical rules and constraints would constitute the over-riding controls and limit mining flexibility. The kimberlite orebodies with which they were concerned are highly variable both geotechnically and in terms of the grades that can occur within a pipe. This makes mining to a constant grade difficult, the more so if geotechnical constraints are taken into account. It was considered necessary to model the entire cycle of ore extraction and processing, taking into account a range of geotechnical and other constraints such as: • the ore flow process in which the material moves from node to node. The nodes modelled

are drifts, ore passes, haulage systems, underground accumulation areas, shaft systems and the treatment plant;

• metallurgical constraints that include plant capacities and the blends of rock types required; • economic constraints including cost and price changes, discounting, exchange rates and

profit to revenue ratios; and • geological constraints such as grade and stone size and density. Guest et al (2000) considered that by using the combined constraint set, it should be possible to produce a better or optimised production mix resulting in better utilisation and efficiency of production resources, even though only small areas of flexibility existed because of the strict geotechnical rules and constraints imposed. The objective function chosen for the linear and mixed integer programming was to maximise the Net Present Value (NPV) over the mine life and over each of several user-defined time periods. The maximisation of NPV is closely related


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to maximising ore production. Thus the model should aim to minimise waste or dilution tonnage which generates no revenue but imposes costs. With the use of discounting factors in user-defined periods, the model will always seek to maximise production sooner rather than later because this will optimise the NPV as indicated by the objective function. Although, as has been indicated, this system is still undergoing further development, Guest et al (2000) concluded that it offers the following advantages over the earlier draw control systems used on De Beers’ block caving mines: • Draw control is now an integral part of the business cycle using inputs from geology,

mining, metallurgy and finance. • The approach optimises returns over the mine life as well as over multiple time periods

within the mine life. The model always takes into account the long-term implications when solving particular short-term problems.

• The system is flexible and generic and provides fast turn around enabling planners and

managers to explore a range of options and have confidence in the outcomes. • All historical production data are stored in the system. This allows re-planning and re-

depletion exercises to be carried out and the results to be compared with the current or other depletion strategies. The linear and mixed integer programming approach has produced 20% improvements over the former manual spread-sheet systems.

• The ability to store the desired plan with its constraint set makes the management, control

and auditing tasks more feasible. • For feasible outcomes, the optimum solution for a given constraint set is always presented. • With the advent of new hardware and software and the incorporation of new algorithms,

run times will decrease rapidly.

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CHAPTER 8

GEOTECHNICAL MONITORING 8.1 THE PURPOSES OF MONITORING

onitoring is the surveillance of engineering structures either visually or with the aid of instruments. Throughout mining history, miners have monitored the performance of their excavations using their eyes and ears. They have made visual

observations of rock falls, of the opening of cracks in the rock and of the effects of convergence on timber supports. They have also traditionally listened to the rock “talking” – an early form of seismic monitoring. Geotechnical monitoring is an integral part of the modern mine rock mechanics programs that have developed over the last 40 – 50 years. Geotechnical monitoring practice is now so well developed that it may appear that, on occasion, a monitoring program has been undertaken simply because it had to be seen that some action was being taken and other ideas were in short supply. Dunnicliff (1988) makes the point well in the frontispiece of his comprehensive book on geotechnical instrumentation for field monitoring: “Every instrument on a project should be selected and placed to assist with answering a specific question: if there is no question, there should be no instrumentation.” Following Franklin (1977), Brady and Brown (1993) have suggested that, in a generic sense, geotechnical monitoring may be carried out for four main reasons: 1. to record the natural values of, and variations in, geotechnical parameters such as water

table level, ground levels and seismic events before the initiation of an engineering project; 2. to ensure safety during construction and operation by giving warning of the development of

excess ground deformations, groundwater pressures and loads in support elements, for example;

M

Chapter 8: Geotechnical Monitoring

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3. to check the validity of assumptions, conceptual models and values of soil or rock mass properties used in design calculations; and

4. to control the implementation of ground treatment and remedial works. In mining geomechanics, most monitoring is carried out for the second and third of these generic purposes, although the first and last are not unknown. Monitoring the safety of the mine structure is a clear responsibility of the mining engineer as is the monitoring of its impact on the local environment. For the second of these purposes, establishing a baseline for measurements is essential (the first of the generic purposes of geotechnical monitoring). Monitoring to check the rock mass response and, as a consequence, adjust the overall mine design or take remedial action, is equally important. The use of cave induction techniques and the monitoring of the results provides an example of this in block and panel caving mines (eg van As and Jeffrey 2000). Rock masses are extremely complex media whose engineering properties are difficult, if not impossible, to predetermine accurately ahead of excavation. The models used to predict the various aspects of rock mass response to mining are based on idealisations, assumptions and simplifications. It is vitally necessary, therefore, to obtain checks on the accuracy of predictions made in design calculations. The use of monitoring, particularly for this purpose, is part of the observational method which is a corner-stone of modern geotechnical engineering practice (Peck 1969). In cave mine engineering, geotechnical monitoring is carried out for three main purposes: 1. to check on the initiation and development of caving; 2. to check the stability of the extraction level installations and other items of mine

infrastructure; and 3. to monitor the development of surface subsidence. In each of these instances, there are elements of safety assurance and of checking on design assumptions and predictions. Monitoring for these purposes will be discussed in turn in this chapter. There is a fourth main type of monitoring in block and panel caving mines - monitoring the effects of the draw control strategies and the resulting fragmentation. It may be argued that this is not strictly geotechnical monitoring but it is an important part of ensuring the overall cost-effective operation of a caving mine. The issues of draw control and fragmentation have been the subjects of previous chapters of this book. The measurement of fragmentation was discussed in Chapter 4 and so that discussion will not be repeated here.


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8.2 GEOTECHNICAL MONITORING SYSTEMS

8.2.1 General Considerations

Having decided why we are monitoring, a number of other questions arise: What do we monitor? How do we monitor? Where and when do we monitor? What accuracy and precision (or reliability) are required in the measurements? How do we validate and interpret the results? These questions can only be answered in detail on a case by case basis, especially in terms of their application to caving. However, some useful general guidance is available. A detailed account of geotechnical instrumentation is given by Dunnicliff (1988). Brady and Brown (1993) discuss geotechnical monitoring in underground mining generally while Szwedzicki (1993) provides a useful range of case history and more general papers on geotechnical monitoring in open pit and underground mining. 8.2.2 What is Monitored?

Not all physical quantities or derived parameters such as stress, can be measured directly. Although it may appear that a wide range of geotechnical variables may be monitored, only two basic physical responses, displacement and pressure can be measured relatively directly with current technology. The third important parameter measured in geotechnical monitoring systems is time. The earliest systematic mine monitoring was probably of the surface subsidence associated with mining operations. The problem of surface subsidence was first studied in Belgium in the 1830s. The studies of Fayol (1885) were among the first systematic subsidence surveys to be made. Young and Stoek (1916) report measurements of surface subsidence associated with the underground mining of coal, iron ore and salt in Europe and the USA in the nineteenth century. Surface subsidence associated with the block caving method of mining has been monitored since at least the 1920s (eg Fletcher 1960). Following Terzaghi’s formulation of the principles of modern soil mechanics (Terzaghi 1925) and his development of the observational method in geotechnical engineering, the use of geotechnical instrumentation to assist with field observations (eg of the performance of dams) developed in the 1930s and 1940s. Simple mechanical and hydraulic instruments were used to


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measure deformations and earth and water pressures (Dunnicliff 1988). Early monitoring instruments purpose built for underground mining applications were for measuring convergence (eg Greenwald et al 1937) and later stress components. The development of borehole extensometers (Merrill 1954, Potts 1957) and of stressmeters (Potts 1954) “took off” in the 1950s and these instruments were applied to caving mines (eg Merrill 1962). Microseismic monitoring in underground mines (involving essentially the measurement of displacement and time) was also being investigated at that time (Obert and Duvall 1957). Brown (1993) listed a number of the items that may be monitored in a mine geotechnical monitoring program: • fracture or slip of the rock on an excavation boundary (observed visually); • movement along or across a single joint or fracture (either monitored by a simple

mechanical “tell-tale” or measured more accurately); • the relative displacement or convergence of two points on the boundary of an excavation; • displacements occurring within the rock mass away from the excavation boundary; • surface displacements or subsidence; • changes in the inclination of a borehole along its length; • groundwater levels, pressures and flows; • changes in the normal stress at a point in the rock mass; • changes in the loads in support or reinforcing elements such as steel sets, props, rock bolts,

cables and concrete; • normal stresses and water pressures generated in fill; • settlements in fill; • seismic and microseismic emissions; and • wave propagation velocities. As was noted earlier and will become apparent from this list, displacement is one of the primary quantities measured in mine geotechnical monitoring programs. Measurements can be made of the absolute displacements of a series of points on the boundaries of an excavation or, with


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more difficulty, within the rock mass. The relative displacement or convergence of two points on the boundary of an excavation is easier to measure than the absolute displacement. An example of a convergence monitoring system in an extraction level drift is shown in Figure 8.1. Because the relative displacement of two points can usually be measured, a measurement of normal strain can be obtained by assuming that the strain is uniform over the base length of the measurement.

Figure 8.1: Drawpoint drift convergence monitoring station, Henderson Mine, USA (Brumleve and Maier 1981)

It is important to recognise that the “measurement” of most other variables of interest, notably forces and stresses, requires the use of a mathematical model and material properties (eg elastic constants) to calculate the required values from measured displacements, strains or pressures. Indeed, as Burland (1967) notes, “stress is a philosophical concept – deformation is the physical reality.” As a general rule, it is preferable to use directly measurable parameters for purposes of comparison and decision making rather than parameters calculated from mathematical models using measured parameters as input. 8.2.3 How is it Monitored?

The modern instrumentation systems used to monitor given geotechnical variables will usually have three components. A sensor or detector responds to changes in the variable being monitored. A transmission system which may use rods, electrical cables, hydraulic lines or


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radiotelemetry devices, transmits the sensor output to the read-out location. A read-out and/or recording unit such as a dial gauge, pressure gauge, digital display, magnetic tape recorder or computer, converts the data to a usable form and presents them to the engineer. Usually, the data will then be stored in a database and be available for processing and presentation in various ways. The monitoring systems used in modern mining operations can be very sophisticated and expensive. However, it should be remembered that valuable conclusions can often be reached from visual observations and observations made using very simple measuring devices. Laubscher (1981, 2000) has emphasised the value of “keeping it simple” in monitoring in caving mines. In order that the monitoring system should fulfil its intended purpose economically and reliably, it should satisfy a number of criteria: • easy installation, if necessary under adverse conditions; • adequate sensitivity, accuracy and reproducibility of measurements; • robustness and suitable protection to ensure durability for the required period of operation; • ease of reading and immediate availability of the data to the engineer; and • negligible mutual interference with mining operations. The terms accuracy, error, precision, sensitivity and resolution as applied to measuring devices require careful definition. Brady and Brown (1993) give the following definitions. The stated accuracy of an instrument indicates the deviation of the output or reading from a known input. Accuracy is usually expressed as a percentage of the full-scale reading. The error is the difference between an observed or calculated value and the true value. Errors may be either systematic or random. Precision is a measure of the ability of the instrument to reproduce a certain reading. It may be defined as the closeness of approach of each of a number of similar measurements to the arithmetic mean. Precision and accuracy are different concepts. Accuracy requires precision and an absence of bias, whereas precision implies a close grouping of readings whether they are accurate or not. The sensitivity of an instrument is variously defined as the ratio of the movement on the read-out unit to the change in the measured variable, the input to output ratio (the inverse of the


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previous measure) or the smallest measurement detectable by the instrument. Sensitivity defined in the third way may depend on the readability of the read-out unit (the closeness with which the scale may be read), and the least count (the smallest difference between two indications that can be detected by the read-out unit). The modes of operation of the sensing, transmission and read-out systems used in monitoring devices may be mechanical, optical, hydraulic or electrical. Although simple and valuable mechanical and hydraulic devices are still in use, the emphasis in modern systems is on electronic systems with digital or analog data transmission and computer-based recording and processing systems. It is beyond the scope of this book to provide a detailed account of the range of monitoring techniques and instruments available. As an illustration, Table 8.1 gives a classification and summary of the methods available for measuring displacements of various types. 8.2.4 Where and When is it Monitored?

It is an all too common experience for engineers attempting to interpret the results of a monitoring program to be disappointed that measurements weren’t taken earlier (when?), that they hadn’t been taken in different locations (where?) and that some additional piece of information is not available (what?). In other words, the planning of monitoring programs is often deficient. In mining, operational constraints may mean that the where? and the when? are addressed inadequately from a monitoring perspective. Of course, the answers to these questions and the optimal design of monitoring systems and layouts will vary with the purpose of the monitoring program. The questions of where? and when? become especially challenging when the monitoring data are to be used in formal back analysis procedures to estimate parameters such as the pre-excavation stresses or the in situ rock mass properties (see, for example, Akutagawa et al 1991a,b). Intuitively, it appears that a number of factors could influence the efficiency of the parameter estimation process and the accuracy of the back-analysed solutions assuming, of course, that the basic model used for the calculations is valid. Among the factors relating to the monitoring system itself are likely to be (Brown 1993): the quantities measured; the numbers of measurements; the locations of the measurements; the accuracy of the measurements; the directions of the measurements or the components measured; the distances between measuring points; and the timing of the measurements with respect to the excavation process.


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Table 8.1: Displacement monitoring methods (Windsor 1993)

Deformation Measurement Technique

Measurement Access

Measurement Method

Measurement Sensitivity

OBSERVATIONAL TECHNIQUES

GPS

Terrestrial Surveying

EDM & Automatic Surveillance

E

E

E

M

M

A

M

M

M

INSTRUMENT TECHNIQUES

Movement Indicators

Axial

Shear

B/F

B/F

O

O

L

L

Convergence Indicators

Wire/Tape

Rod

F

F

M

M/A

M

M

Strain Meter

Resistance Strain Gauges

Vibrating Wire Strain Gauges

B

B

A

A

H

H

Joint Meters

Glass Plates

Pin Arrays

Strain Gauges

Proximity Transducer

Fibre Optic

Potentiometers

F

F

B

F

B

B/F

M

M/A

A

A

A

A

M

M

H

H

H

H

Extensometers

Fixed Extensometer

Wire/Rod

Reference Point Sensing

Strain Sensing

Portable Extensometer

Magnetic Anchor

Magneto-strictive

Sliding Micrometer

B

B

B

B

B

B

M/A

M/A

A

M

M

M

H

H

H

H

H

H

Inclinometers

Fixed Inclinometer

Portable Inclinometer

B

B

A

M

H

H

Deflectometers B A H

Extensometer-Inclinometer B M H

Extensometer-Deflectometer B M H

F = Rock Face, B = Borehole E = Exposure, M = Manual, A = Automatic, O = Observational, L = Low, M = Medium, H = High

In addition to these considerations, of course, it is necessary that the important practical questions of the instrumentation and installation costs, of access and of potential disruption to mining operations be taken into account in the planning process. To the best of the writer’s knowledge, there is no general technique available for determining an optimal measurement


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pattern for a given problem. Here again, planning must proceed on a case by case basis bearing in mind the purpose of the measurements and the factors identified above. One further factor must be borne in mind in planning and interpreting monitoring results in rock mechanics. As Hudson (1992) has pointed out, many of the responses and parameters monitored may, in fact, be coupled and interact. Hudson has developed the interaction on matrix as a method of identifying and studying how one factor affects another. The matrix can be presented at different levels of detail. Figure 8.2 illustrates the point that the monitoring of parameters and interactions in the finest level matrix will be the easiest to understand and use. In the coarse resolution left hand matrix, R, S and P are basic parameters (eg rock mass structure, in situ stresses). RS represents the influence of R on S and SR represents the influence of S on R, etc. The right hand matrix shows the detail of cell P in the left hand matrix. Hudson (1992) argues that the decodability of any monitored composite variables must be considered very carefully and that his systems approach should be used in deciding what to measure for any back analysis requirements.

Figure 8.2: Rock interaction matrix illustrating the ease or difficulty of interpreting

monitoring data for use in back analysis (Hudson 1992)


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8.3 MONITORING THE INITIATION AND DEVELOPMENT OF CAVING

8.3.1 Why?

It is highly desirable, if not essential, that the initiation and progress of caving be monitored, especially in a new block caving operation or location. The monitoring data may be required to: • indicate that a critical undercut dimension has been achieved and that caving has been

initiated; • provide a measure of the rate of caving for use in establishing an appropriate production

rate and developing a draw control strategy; • ensure that an open void or air gap does not develop above the caved mass of ore; • ensure that exploration, access or instrumentation drifts overlying the cave can be

abandoned in advance of the cave propagating to that level; and • assist in the interpretation of subsidence and other monitoring data. 8.3.2 What and How?

There are four general types of measurement made in monitoring cave initiation and propagation – manual distance measurements in open holes, time domain reflectometry (TDR), microseismic monitoring and cavity measurement or surveying. Each of these methods will be discussed below. In practice, these methods are often used in combination. In addition, the use of extensometer arrays is sometimes also attempted, although such attempts have not always been entirely successful (eg La Rosa and Chen 1997). Manual methods. Some form of depth measurement is made manually in one or more open holes drilled from the surface or from a drift overlying the cave. Measurements of this type are of most use when there is an air gap above the caved ore. The application of ingenuity has led to a number of variants of this method being developed locally at mine sites. For example, a weight (eg a used mill ball) may be attached to the end of a measuring cable or rope which is lowered down a hole onto the top of the caved ore pile and a length measurement taken from the hole collar. The weight is then pulled up until it contacts the cave back and another measurement taken. From these two measurements, the height of the air gap and the elevation of the cave back may be determined (Chen 2000). A folded, hinged and/or spring loaded device may be attached to the end of the measuring cable such that it opens and can be drawn back to be in contact with the cave back once the cable intersects the void. An example of such a device is shown in Figure 8.3.


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Figure 8.3: Cable and plug method of measuring the position of the cave back (Julin 1992)

This method has been developed further at the Tongkuangyu Copper Mine in China (Chen 2000, Zhang 1997). The measuring device used consists of a geophone-type detector with “whiskers”, a wire with measuring marks and a receiver unit. As the detector travels along a borehole above the cave with the whiskers in contact with the borehole wall, the noise generated is detected by the receiver. When the detector reaches the void and the whiskers lose contact with the borehole walls, the sound disappears and the length of wire down the hole gives a measurement of the location of the cave back. The detector is then lowered onto the caved ore pile. When the whiskers touch the broken ore, noise is detected again and the difference between this and the first measurement gives the height of the air gap. A series of monitoring holes may be used to determine the shapes of the cave back and the top of the ore pile and the variation in air gap thickness may be determined. Time Domain Reflectometry (TDR). TDR is an electrical pulse testing technique originally developed to locate faults in coaxial power transmission cables. Subsequently, it was adapted with considerable success for monitoring the deformation of cables grouted into rock masses. Movements in the rock mass deform the grouted cable which changes the cable capacitance


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locally and thereby the reflected wave form of the voltage pulse. By monitoring changes in these reflection signatures, it is possible to monitor both local extension and local shearing (Dowding et al 1988). TDR methods have been the main methods used to monitor cave initiation and propagation for the last 20 years. Examples of this application are given by Brumleve and Maier (1981), Stewart et al (1984), Rech and Watson (1994), La Rosa and Chen (1997), de Nicola Escoba and Fishwick Tapia (2000), Rojas et al (2000b) and Rachmad and Sulaeman (2002). When the cave back or possibly the zone of loosening or the seismogenic zone illustrated in Figure 1.8, reaches the grouted cable, the cable will extend and may shear. It will eventually break and the pulse will be reflected back up the cable. Although this produces a very clear change in signal as in the example shown in Figure 8.4, the location of the break does not necessarily coincide with the location of the cave back surface. Distances can be calculated from pulse travel times. As in the manual open hole methods, measurements made in a series of holes can be used to estimate the shape of the cave back over the caving region (Chen 2000).

Figure 8.4: TDR trace showing the onset of caving, Northparkes E26 block cave, Lift 1

(La Rosa and Chen 1997)

Microseismic methods. Microseismic monitoring has been carried out in caving mines primarily to assist in resolving the problem of rock bursts that have occurred in high stress


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environments (eg Dunlop and Gaete 1997) although a trial of a microseismic system for monitoring cave growth was undertaken at the Henderson mine in the 1970s (Leighton 1978). The results obtained from more recent microseismic monitoring for rock burst management and developments in hardware and software, suggested that microseismic monitoring could be used to study the development and mechanics of caving (Chen 1998, Dunlop and Gaete 1997). Microseismic events result from rock fracture or from slip on pre-existing discontinuities. Since these mechanisms are likely to be involved in cave initiation and propagation following undercutting (see Section 1.2.2), it follows that microseismic systems should be able to monitor their development (eg Duplancic and Brady 1999, Trifu et al 2002). Brady and Brown (1993) give the following description of seismic or microseismic monitoring systems that have been used in underground mining generally. The wave propagating from the source of a seismic or microseismic event is detected by a geophone (a velocity gauge suitable for detecting frequencies in the range 1-100 Hz) or an accelerometer which converts the mechanical vibration into an electrical analogue. The electrical signal is then amplified and transmitted to a receiving station. After signal conditioning, the signal goes to the computer interface which determines if it is strong enough to trigger the timing and control component of the system. If so, the arrival times of the wave at each of the geophones or accelerometers are determined. These data are used to locate the source of the event in three-dimensional space using geometric and seismic velocity relations established by calibration. The amplitude and duration of the signal may be used to determine the relative magnitude of the event. The best documented example of the use of microseismic monitoring to study the development of caving and the effect of a cave inducement technique is that at the Northparkes E26 block cave, Lift 1 (Chen 1998, Duplancic and Brady 1999, van As and Jeffrey 2000). Figure 8.5 shows a schematic of the system which is described by Duplancic and Brady (2000) in the following terms. The system is made up of seismic sensors, processing seismometers, a multiplexer and a central computer. Triaxial accelerometers were chosen as the seismic sensors because of their wide acceleration magnitude range. They allow definition of seismic events occurring close to the transducer. The 12 accelerometers were grouted into horizontal holes that were core drilled to depths of 10 m. A dedicated processing seismometer calibrates and monitors each accelerometer. It also stores records of motion from the accelerometer for comparison with those from other seismometers via a central computer. The intelligent multiplexer combines messages from several seismometers onto one communication line to the central computer. The multiplexer also rebroadcasts signals from the central computer for the synchronisation of recording at the stations. The central computer runs the controlling software for the system. It compares the records from individual seismometers and if the parameters set to define a seismic event are satisfied, that event is registered as a seismic event. For initial processing of seismic data, a number of physical parameters of the


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rock mass (eg shear and compressional wave velocities) are required. These parameters are determined through a calibration process and the analysis of subsequent monitoring data.

Figure 8.5: Schematic diagram of the microseismic monitoring system, Northparkes E26 block cave, Lift 1 (Duplancic and Brady 1999)

Four types of seismic or microseismic event were identified at Northparkes (Chen 1998): 1. caving events occurring in or above the cave back; 2. events arising from crushing and the associated machines; 3. events generated by blasting; and 4. events occurring within or under the broken ore as a result of production activities. Crushing and blasting events had distinctive wave forms and could be distinguished. Seismic source locations and seismic moments could also be used to identify caving-related events (Duplancic and Brady 1999). Figure 8.6 shows a summary of the results obtained in the months of February and March 1997. These results show that the immediate cave back is a destressed zone with few seismic events. Above this is a high energy release zone in which multiple seismic events occur as a result of slip on pre-existing discontinuities in the seismogenic zone identified in Figure 1.8 by Duplancic and Brady (1999). The results of the microseismic monitoring agreed well with the results of TDR and open hole measurements. The microseismic system also allowed hazardous areas to be identified and avoided by personnel and equipment (Chen 1998). Van As and


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Jeffrey (2000) also used the system to locate the fracturing events associated with a campaign of hydraulic fracturing to induce caving.

Figure 8.6: Caving-related microseismic events, Northparkes E26 block cave, Lift 1

(Chen 2000) Trifu et al (2002) give a more recent example of the use of microseismic monitoring to characterise the cave front at the Ridgeway Gold Mine, New South Wales, Australia. Although Ridgeway uses sublevel rather than block or panel caving, the techniques used and the results obtained by Trifu et al (2002) are relevant to present considerations. The Ridgeway microseismic array consisted of 11 triaxial accelerometers installed in four vertical boreholes drilled from surface with up to three sensors per hole, grouted in a staggered array at depths of approximately 90, 200 and 500 m. Using event locations (which were accurate to within 2-4 m) and source mechanisms, this system provided continuous, three-dimensional tracking of the cave front as it progressed to surface from an initial mining depth of approximately 500 m. Trifu et al (2002) report that initially a microseismically active zone of about 20 m high developed in the Peripheral Host Rock domain immediately above the orebody. Source mechanism analysis showed that both shear and volumetric failure occurred in this good quality rock mass (RMR = 70). The microseismic sources delineated a well-developed arch-shaped caving zone in this material. As the cave developed into the less stiff, overlying Caprock domain (RMR = 61), microseismicity became more scattered and less intense. Finally

+ N S 5 3 5 0 0 .0 + N S 5 3 4 0 0 .0 + N S 5 3 3 0 0 .0 + N S 5 3 2 0 0 .0+ E W 1 1 00 0 .0+ E W 1 0 90 0 .0+ E W 1 0 80 0 .0+ E W 1 0 70 0 .0

0 1 0 0 mS C A L E

O p e n c u t In f lu e n c e dZ o n e

L o w e r E n e rg y

D e s tre s s e d Z o n e

H ig h e r E n e rg y

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3 .0

2 .5

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1 .5

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microseismic activity decreased even further as the cave progressed into the more highly fractured and significantly less stiff Near Surface Caprock domain (RMR = 49). Cavity Monitoring Systems. If an open void or air gap is allowed to develop above the ore pile, and if line-of-sight access is available into the cave, laser-based cavity monitoring or measurement systems of the type used in surveying open stopes may be used to obtain three-dimensional measurements of the shape of the cave back and of the top of the ore pile. Stewart et al (1984) refer to the use of a surveying laser and a hand held rangefinder to measure the profile of an open cave from inspection drifts at the Henderson mine when there was concern that a stable arch was forming over part of the cave. Laser-based systems may be used to measure the cave back profile when access is available (eg de Nicola Escobar and Fishwick Tapia 2000) and have also been used to examine undercuts for any unwanted remnant pillars. 8.4 EXTRACTION LEVEL AND INFRASTRUCTURE MONITORING

8.4.1 Why?

The monitoring of the extraction level installations and of other items of mine infrastructure is carried out for similar reasons as much of the geotechnical monitoring of other types of underground mines. Because of the high development costs and relative lack of flexibility in terms of the mining method and the locations of centres of production in a caving mine, it is essential that the extraction or production level remains stable and available for production operations. Many examples of the distress and failure of extraction level installations have been experienced in the history of block and panel cave mining. Monitoring results may be required to: • check the rates of deformation and the stability of production drifts as undercutting and

caving progress; • check on the efficacy of the support and reinforcement systems installed; • provide early warning of the development of excessive deformations or other forms of

distress in mine accesses, haulageways and other items of mine infrastructure; • assess the effectiveness of undercut and extraction level development strategies; • ensure safe access to the centres of production for personnel and machines; • manage a rock burst hazard in high stress environments; • check on assumptions made in design;


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• ensure the continuing integrity of the overall mine structure; and • assess the effectiveness of any remedial or repair work. 8.4.2 What and How?

In pursuit of these objectives, a wide range of measurements and observations may be made: • convergence monitoring in extraction level drifts, drawpoints and other items of

infrastructure; • extensometer measurements to identify the depth to which displacement is occurring in the

rock mass around infrastructure excavations; • the measurement of stress components in concrete and shotcrete linings usually by

hydraulic pressure cells; • mine-wide and local seismic and microseismic monitoring systems; • damage mapping in concrete and shotcrete linings and in exposed rock on the extraction

level and in drawpoints; • monitoring the extension of, and movement on, cracks in linings; • stress measurements on and around the extraction level to monitor stress changes

associated with undercutting and the development of the cave; • monitoring loads in support and reinforcing elements such as steel supports, rock bolts and

cable bolts; • seismic tomography to measure pillar damage; and • borehole camera surveys to measure the degree of rock mass damage. 8.4.3 Examples

Convergence measurements. Figure 8.1 shows a simple example of a convergence monitoring station in a drawpoint drift at the Henderson mine. In this and many similar cases, the convergence is measured by a tape or rod extensometer. Not only the total convergence, but also the convergence rate and any acceleration of that rate, are important indicators in such cases (eg Rachmad and Widijanto 2002).


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Figure 8.7 shows some convergence measurements made by Agapito and Shoemaker (1987) at two stations in a haulage drift in the Questa Mine, New Mexico, USA. Station 1 located in the haulage near an intersection, showed high horizonal convergence and an initially high convergence rate attributed to the influence of stresses induced by the passage of the cave. The vertical convergence was much lower. Station 2, located 10 m to the east of Station 1 away from the intersection and in better ground, gave low convergences, indicating stable conditions. The flattening of the curves indicates a long-term stability trend. However, the high horizontal convergences and convergence rates at some locations such as Station 1, suggested that the initial Split Set bolts and shotcrete support was inadequate. The installation of 6 m long fully grouted rock bolts and re-shotcreting made this area stable. This example illustrates how simple convergence measurements can warn of potential instabilities, the need for corrective measures and the effectiveness of those measures.

Figure 8.7: Horizontal and vertical convergences at two monitoring stations in a haulage drift, Questa Mine, New Mexico, USA (Agapito and Shoemaker 1987)

Stress changes. Early examples of the monitoring of stress changes resulting from undercutting and caving are given by Merrill (1962), Merrill and Johnson (1964) and Utter and Tesch (1965). Brumleve and Maier (1981) give an excellent example of the use of monitoring to study the changes in the stress field and their effects on and around the extraction level of the Henderson mine over a three year period. The pre-mining stress field was determined from overcoring measurements made at three sites. The changes in the stress field accompanying

Con

verg

ence

(mm

)


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cave growth, and in the local stress field associated with the cave abutment zone, were measured by a borehole deformation gauge, vibrating wire gauges and hydraulic stress cells. Table 8.2 shows the changes in three stress components measured ahead of the advancing undercut in a pillar on the 8100 (production) level. The hydraulic pressure cells were grouted into horizontal holes in the pillar in three orientations – flat to measure vertical stress, 45o away from the cave to measure the tangential stress and 45o towards the cave to measure the radial stress. The results show that the minimum stress increment is generally radial and that it is tensile close to the line of the undercut. The increments of tangential and vertical stress just ahead of the undercut can be very high. These results are in accord with the results of elastic stress analyses of the type discussed in Chapter 5.

Table 8.2: Monitored stress changes due to undercut, 8100 level, Henderson Mine (Brumleve and Maier 1981)

Distance to

Cave m

Vertical Stress kPa

Radial Stress kPa

Tangential Stress

kPa 0 -3,792 40,680

7.6 9,136 -2,068 38,610

15.2 17,100 552 33,780

22.9 19,820 6,550 19,310

30.5 13,440 5,171 12,760

38.1 9,032 3,309 8,963

45.7 6,274 1,931 5,861

53.3 4,482 1,034 3,447

60.9 3,309 621 2,068

68.6 2,413 345 1,310

76.2 1,586 345 690

83.8 827 621 552

91.4 759 345

Negative sign indicates relative tension

Reinforcement loads. A further element of the monitoring program reported by Brumleve and Maier (1981) is the measurement of loads in the cable bolts used to reinforce the production level ahead of undercutting and caving. Figure 8.8 illustrates the 13 m long instrumented cable bolts installed between production level and undercut drifts. Figure 8.9 shows the axial tension in one of these bolts monitored over a period of 6 months as the undercut approached and advanced past the bolt. In period A, the rock mass between the undercut and production levels and the cable bolt tension were stable. The load cell was outside the zone of influence of the abutment zone of the


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cave. In period B the undercut was advanced to the location of the load cell. Caving stresses in the abutment zone compressed the rock mass reducing the cable tension. In period C, the undercut was advanced past the load cell reducing the abutment zone stress on the rock mass which was able to expand into the overlying undercut area, tensioning the cable to a higher load. In period D, the undercut was 33.5 to 42.7 m past the load cell and the cable tension stabilised at a high value. Subsequently, as caving continued, the rock mass in which the instrumented cable was located was reloaded by the broken ore in the cave and the cable tension decreased. Visual inspection of the production drift over the monitoring period showed that the loading and unloading associated with cave advance caused spalling and initial cracking of the drift and drawpoint linings.

Figure 8.8: Instrumented cable bolt, 8100 level, Henderson Mine (Brumleve and Maier

1981)


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Figure 8.9: Monitored cable bolt tension showing passage of cave front, Henderson Mine (Brumleve and Maier 1981)

Pillar damage. Rojas et al (2000b) provide an example of the use of a wide range of monitoring techniques to confirm the effectiveness of the pre-undercut method introduced in the Esmeralda sector of Codelco-Chile’s El Teniente mine. Figure 8.10 shows the results of seismic tomography and borehole camera surveys carried out on extraction level pillars under conventional (post-undercut) and pre-undercut conditions. The improvements in observed pillar rock mass quality are quite marked. 8.5 MONITORING SUBSIDENCE AND GROUND MOVEMENT

8.5.1 Why?

The objective in block and panel caving is to cause the ore and any overlying rock to cave progressively in a controlled manner. A consequence of this is that the cave will usually extend to the surface producing surface subsidence. In many cases, the cave propagates close to vertically over the undercut area (eg Stewart et al 1984, Duffield 2000). Any soils or weathered rocks at the surface may not be able to sustain a near-vertical face and can be expected to fail back to a shallower stable slope angle. Experience has also shown that the direction of propagation of the cave to surface can be influenced by the topography and by the presence of

Cab

le lo

ad (k

N)


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structures such as faults or dominant inclined bedding, schistocity or jointing (Laubscher 2000). The expression of surface subsidence may also be influenced by a pre-existing open cut and may, in turn, cause instability of those slopes (Owen 1981). The mechanisms of cave propagation to surface will be discussed and illustrated in Chapter 9.

Conventional Pre-undercut

a) Poor rockmass condition pillar (3000 to 4300 m/s) and good condition pillar (5500 to 7000 m/s) P wave velocities

Tom

ogra

phie

s

b) Video images showing a poor and a good rockmass condition

Bor

ehol

e ca

mer

a

Figure 8.10: (a) Seismic tomography, and (b) borehole camera surveys of extraction level pillars for conventional or post-undercut and pre-undercut strategies, El Teniente

mine, Chile (Rojas et al 2000b) Because of these features of the subsidence resulting from block and panel caving, and because it is difficult to predict them in advance of mining, it is necessary that the development of surface subsidence and of ground movements around the cave be monitored. This monitoring information is required to, for example:


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• ensure safety by helping define exclusion zones above and around the cave both on the surface and underground;

• give warning of the potential impact of subsidence on surface installations; • monitor the stability of overlying or nearby open pit slopes; and • give warning of any otherwise unanticipated impact of caving on underground installations

adjacent to the cave and on nearby current and future mining blocks. 8.5.2 What and How?

Depending on the geometry of the orebody, the surrounding geology and the local topography, subsidence and ground movement measurements may take several forms. The most common types and methods of measurement are: • surface measurement of the horizontal and vertical movements of a grid of surface survey

points. Base-line measurements should be made before mining commences and a survey control point should be established outside the zone of influence of the cave. Precise levelling and electronic distance measurement techniques are usually used;

• aerial photogrammetric surveys of a series of surface survey points. Aerial photography in

its own right is an invaluable method of assessing surface subsidence, particularly when the cave has broken through to surface;

• surveys and measurement of the opening and movement on surface or underground cracks;

and • underground surveying and extensometer and inclinometer measurements to monitor

ground movements underground around the cave. 8.5.3 Examples

Several examples of the surface subsidence resulting from block and panel caving mines will be given in Chapter 9. In fact, details of the techniques used and the results obtained from surface subsidence monitoring programs are not published in the open literature as commonly as might be supposed. Indeed, most of the published case histories are now some years old. Fletcher (1960) described the history of subsidence associated with block caving at the Miami Mine, Arizona, USA over a 30 year period. Some details of this example are given in Section 9.3. Panek (1981a) and Panek and Tesch (1981) reported the results of an extensive program of ground movement monitoring carried out both on the surface and underground at the San


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Manuel Mine, Arizona, USA. One of the most complete examples published is the development of subsidence associated with early panel caving at the Henderson Mine, Colorado, USA (Stewart et al 1984). In this case, aerial photography, a surface survey grid and time domain reflectometry were use to monitor the progression of caving and the development of surface movement and subsidence on Red Mountain overlying the mine. Details of this example will be given in Section 9.3. Finally, the surface crater produced by a sudden collapse at the Salvador Mine, Chile, described by de Nicola Escobar and Fishwick Tapia (2000) will be illustrated in Chapter 10 where major collapses and air blasts will be discussed. In this case, the surface crater was measured by an aerial photogrammetric survey.

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CHAPTER 9

SURFACE SUBSIDENCE 9.1 INTRODUCTION

n general terms, subsidence is the lowering of the ground surface following the underground extraction of ore or another resource. It results to a greater or lesser extent from all forms of underground mining but is particularly pronounced in caving methods of

mining. As the orebody caves and is extracted progressively, the overlying cap rock will also cave and move downwards with the remaining ore producing a characteristic surface depression. This form of subsidence is often referred to as discontinuous subsidence to distinguish it from the continuous subsidence that is characteristic of longwall coal mining. Discontinuous subsidence is characterised by large vertical subsidence over restricted surface areas and the formation of steps or discontinuities in the surface profile. As an example, Figure 9.1 shows the surface expression of the caves produced at the Henderson and Urad mines, Colorado, USA.

Figure 9.1: Surface craters produced by caving at the Henderson and Urad mines,

Colorado, USA (Laubscher 2000)

I

Chapter 9: Surface Subsidence

347

If, as is often the case in block and panel caving mines, the orebody is vertical with a well-defined geological cut-off between it and the surrounding country rock, the cave will propagate vertically to surface with inclined surface slopes forming only in any weak or weathered surface layers. However, a number of features of the orebody and the local geology and topography can influence the overall angle made by the cave boundary with the vertical or the horizontal, sometimes called the angle of draw and the angle of break or subsidence, respectively. Some of these factors are: the dip of the orebody; the shape of the orebody in plan; the depth of mining and the associated in situ stress field; the strengths of both the caving rock mass and the rocks and soils closer to the surface; the slope of the ground surface; major geological features such as faults and dykes intersecting the orebody and cap rock; prior surface mining; the placement of fill in a pre-existing or the newly produced crater; and nearby underground excavations.

The influences of some of these factors will be illustrated by the examples to be given in Section 9.3 and the analyses to be developed in Sections 9.4 and 9.5. For practical mine design, it is also important to define a zone of influence around the cave as well as the caving or subsidence zone itself. Within the zone of influence there may be deformations of the rock mass which, although small in comparison to those occurring within the caving zone, will be large enough to damage excavations and mine infrastructure located within the zone. The rock within this zone could be highly or over-stressed by the redistribution of stresses that accompanies the development and upwards progression of the cave, and then destressed (at least in the lateral direction) as the cave develops fully. The limits of the zone of influence of the cave are best determined by a combination of monitoring, local experience and numerical modelling (Butcher 2002b). An example of the calculation of the zone of influence for the Ten 4 Sector of the El Teniente mine (Karzulovic et al 1999) will be given in Section 9.6. 9.2 TYPES AND MECHANISMS OF DISCONTINUOUS SUBSIDENCE

9.2.1 Types of Discontinuous Subsidence

The classic form of subsidence associated with block caving outlined above is not the only form of discontinuous subsidence that has resulted from underground mining or from block caving. This brief account of the types of discontinuous subsidence illustrated in Figure 9.2 is based on that given by Brady and Brown (1993).


348

Figure 9.2: Types of discontinuous subsidence: (a) crown hole; (b) chimney caving; (c)

plug subsidence; (d) solution cavity; (e) block caving; (f) progressive hangingwall caving (Brady and Brown 1993)

Crown holes (Figure 9.2a) result from the collapse of the roofs of generally abandoned shallow open workings. They are not associated with caving methods of mining and are listed here only for completeness. Chimney caving, piping or funnelling involves the progressive migration of an unsupported mining cavity through the overlying material to the surface (Figure 9.2b). Chimney caves may form in weak overburden materials, in previously caved material or in well jointed rock which unravels progressively. If chimney caving is sudden rather than progressive, it is sometimes known as plug subsidence (Figure 9.2c). The various mechanisms involved will be discussed further in Section 9.2.2. Chimney caving can occur in block and panel caving mines as a result of excessive drawing from an isolated drawpoint as discussed in Chapter 7 and illustrated in Figures 7.1b and 7.2b. Chimney caves are sometimes referred to as sinkholes (eg Goel and Page 1982), but this term is also used specifically to describe the subsidence features associated with pre-existing solution cavities in limestones and dolomites (Figure 9.2d). Szwedzicki (1999) adopts a more general definition of a sinkhole as a cylindrical or conical depression caused by the discontinuous subsidence of a rock mass. The large scale discontinuous surface subsidence resulting from block caving which is the major concern here, is illustrated in Figure 9.2e, in Figure 9.1 and in the examples given in Section 9.3. When the caving orebody is inclined as it is in some caving mines, progressive


349

hangingwall caving may result as illustrated in Figure 9.2f. An analysis of this form of caving to surface will be given in Section 9.5. 9.2.2 Chimney Caving Mechanisms

As indicated above, three distinct chimney caving mechanisms may be identified, each associated with different geological environments. The first mechanism occurs in weathered, weak or previously caved rock. It is a progressive mechanism which starts with failure of the cave back, stope roof or inclined hangingwall. If a stable, self-supporting arch cannot be formed in the material, generally because of its low shear strength, the disintegration or failure may propagate progressively towards the surface as illustrated in Figure 9.3. As material falls it will bulk and tend to fill the stope or drawpoint void. Unless the stope is initially large and open, or unless sufficient material is drawn progressively, the void may eventually become filled with caved material which will provide support for the upper surface and so arrest development of the chimney. It is for this reason that the development of chimney caving is so closely associated with draw control. Figures 7.1a and 7.2a show examples of the flow patterns associated with chimney caving, piping or funnelling to surface in caved materials having particles of different shapes, sizes and gradings.

Figure 9.3: Progressive vertical chimney cave development in homogeneous material

(Bétourney et al 1994)


350

The progressive internal shear failure mechanism illustrated in Figure 9.3 has been well documented in model studies of the failure of shallow tunnels in sand and clay (Atkinson et al 1975) and in model studies of the mining of steeply dipping tabular orebodies. Once initiated, propagation of the failure to surface can be very rapid. This can give the impression that the cave reaches the surface instantaneously and that a plug subsidence rather than a progressive failure mechanism is involved. It has been found that limiting equilibrium analyses of the final collapse configurations give good approximations to the ultimate collapse conditions observed in a range of applications to which this mechanism applies (Atkinson et al 1975, Bétourney et al 1994, Brady and Brown 1993, Egger 1983). This use of limiting equilibrium analysis is developed in Section 9.4. The second chimney caving mechanism is also progressive but results from the unravelling of a discontinuous rock mass. It is similar to the caving of a well-jointed orebody under low stress conditions outlined in Section 1.2 2. The rock material itself may be quite strong and may not fail except by induced tension in bending or point loading. The mechanism is controlled by the pre-existing discontinuities in the rock mass. In the distinct element model of an idealised case of a crown pillar shown in Figure 9.4, a self-supporting arch is developed by the induced stresses in the first case (despite the orientation of the vertical set of discontinuities), whereas unravelling of the crown occurs in the weaker rock mass in the second case. The mechanism illustrated in Figure 7.2b is an example of a similar mechanism operating in a coarse blocky caved material under draw.

Figure 9.4: Distinct element model prediction of crown pillar response in (a) good, and

(b) fair rock conditions (Carter and Miller 1995)


351

In the third mechanism, plug subsidence, failure is controlled by one or more major structural features which provide low strength surfaces on which the plug of undercut rock may slide under the influence of gravity. In this case, the rock will undergo essentially rigid body displacement without breaking up or dilating if the vertical distance through which it falls is restricted. Clearly, this mechanism is amenable to analysis by limiting equilibrium methods. In practice, these three mechanisms of chimney caving are not as clearly distinguished as the preceding accounts might imply. For example, Allen (1934) and Boyum (1961) describe an example of chimney caving that was controlled by geological planes of weakness, but which developed progressively as in the first mechanism. After three years of mining by a caving method at the Athens Mine in northern Michigan, USA, a chimney cave bounded by sub-vertical dykes developed and progressed rapidly to surface through 600 m of cover (Figure 9.5).

Figure 9.5: North-south section, Athens Mine, Michigan, USA, showing plug

subsidence controlled by dykes (after Boyum 1961)


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9.3 EXAMPLES OF SURFACE SUBSIDENCE ARISING FROM BLOCK AND PANEL CAVING

9.3.1 Miami Mine, Arizona, USA

Fletcher (1960) describes the development of the subsidence profile over a period of three decades at the Miami Mine, Arizona, USA, where a mixed copper ore had been mined by block caving since 1926. Initially, the cave had vertical sides but the angle of break (measured with respect to the horizontal) decreased as the depth of mining increased. Tension cracks defined the extremities of the subsidence area. Toppling of blocks isolated by tension cracks occurred in some areas. A major fault, dipping at 45o to the east, governed cave development on the western side of the mine. Figure 9.6 shows the surface profile and the location of tension cracks on the E800 section of the Miami Mine in February 1958. In this case, localised chimney caving as well as mass subsidence occurred.

Figure 9.6: E800 section, Miami Mine, Arizona, USA, showing mined areas and surface subsidence, February, 1958. Coordinates are in feet (after Fletcher 1960)

9.3.2 San Manuel Mine, Arizona, USA

Hatheway (1968) gives a detailed account of the progressive development of subsidence during the first 10 years of block caving at the San Manuel Mine, Pinal County, Arizona, USA, from 1956 to 1965. The subsidence associated with three block caves is described, but only the largest of the three, that associated with the South orebody, will be considered here. At the time of interest, mining was by a block caving method with gravity draw through a grizzly level.

Figure 9.7 shows a plan of the South orebody block cave and subsidence zone, and a vertical section along the centre line of the subsidence zone showing the main features of the pre-mining geology and the subsidence profile at August 1964. The predominant structural feature


353

of the area is the San Manuel fault which has resulted in the placement of a thrust sheet of conglomerates against the quartz monzonite and granodiorite porphyry orebody. As shown in Figure 9.7b, both the massive orebody rocks and the stronger overlying conglomerates have been cut by three major north-west trending faults, the Mammoth and Cholla faults and the less steeply dipping San Manuel fault. Associated with these major faults are several near-vertical normal faults having opposing dips to the north-east and south-west. This faulting exerted a major influence on the development of subsidence.

(a) (b)

Figure 9.7: (a) Plan and subsidence zone, and (b) vertical section along subsidence zone centre line, South orebody, San Manuel Mine, Arizona, USA (after Hatheway

1968)

The progressive development of the subsidence profile with continuing draw is illustrated schematically in the generalised vertical sections shown in Figure 9.8. Subsidence in the initial and the more advanced stages of caving is illustrated. Hatheway (1968) found that surface subsidence developed in the following sequence:

• Chimneying or piping propagated vertically over the initial drawpoints to the orebody –

conglomerate interface at the San Manuel fault (Figure 9.8a). It is assumed that this was the result of uneven initial draw.

• The vertical progression of the chimney was halted and deflected at the San Manuel fault.

There is some suggestion that the stronger conglomerate did not cave immediately, but eventually suffered flexural failure and caved into the void created by continued drawing of the weaker ore.

• Vertical tension cracks formed at the surface marking the initial boundary of the

discontinuous subsidence zone.


354

• As mining progressed, individual chimneys coalesced into a wider subsidence zone as illustrated in Figure 9.8b. At the same time, new tension cracks formed on the surface and progressively became deeper.

• With continuing draw, the caved mass of rock inside the subsidence crater moved

downwards under gravity on steeply inclined shear surfaces as shown in Figure 9.8b. • As mining continued, new tension cracks and shear surfaces developed expanding the

subsidence zone. As shown in Figure 9.7a, by 1965 the South orebody crater was approximately 900 m long, 600 m wide and 150 m deep.

• The north-west trending faults exerted a major influence on the orientation of the tension

cracks and the development of the subsidence crater. In particular, the Cholla fault (see Figure 9.7b) arrested development of the crater in the north-easterly direction for several years.

(a) (b)

Figure 9.8: Development of surface subsidence at the San Manuel Mine, USA, in (a) the initial, and (b) the more advanced stages of mining (after Hatheway 1968)


355

9.3.3 Henderson Mine, Colorado, USA

An overview of panel caving which began at the Henderson Mine in August 1976, was given in Section 1.3.4. As shown in the generalised vertical section in Figure 1.14 and the photograph in Figure 9.1, a subsidence crater, described locally as a glory hole, has been produced on the western side of Red Mountain. Stewart et al (1984) provide an account of the development of the subsidence crater from its first appearance in September, 1980, to late 1983. Cave development and surface subsidence were monitored using the time domain reflectometry, aerial photography and surface surveying techniques discussed in Chapter 8. Caving at Henderson was initiated on the 8155 level in August, 1976 in the west-central portion of Panel 1 (Figure 9.9) at a production rate of 4,540 tonnes per day. This rate had increased to 14, 510 tonnes per day by the end of 1978, and to 25,400 tonnes per day by December, 1979. It remained at this rate until September, 1980. During this period, draw was generally uniform over the expanding undercut area which, by September, 1980, had reached the plan dimensions shown in Figure 9.9. As mining of Panel 1 developed over this initial 50 month period, the cave propagated vertically through more than 1000 m of igneous rock and broke through to surface on 10 September, 1980 as a steep-walled cavity. This represented an overall average vertical growth rate of 0.7 m per day. As shown in Figure 9.9, the cave broke through to surface directly above the initial production area. Stewart et al (1984) suggest that the use of 67 m high boundary cut-off drifts and the steeply dipping fractures at surface probably contributed to the vertical growth of the cave. As opposed to the San Manuel Mine example described above, geological features such as rock unit contacts and alteration zone boundaries appeared to have little influence on the development of the initial subsidence crater.

From 1980, undercutting continued in Panel 1 primarily north of the initial cave area. Panel 2 undercutting took place between March and August, 1981. Production and development were halted temporarily in October, 1982 and resumed in early 1984. During this period, subsidence crater growth did not follow the direction of undercut advance. Rather, under the influence of topography and faulting, crater growth was primarily east-west or up- and down-slope. As in the other examples described here, local faulting (in this case intersecting faults) also influenced the shape of the crater perimeter. Interestingly, during the shut-down in 1982 and 1983, the crater continued to expand in volume, possibly as a result of the compaction of caved material under continued static loading (Stewart et al 1984).


356

Figure 9.9: Plan showing the position of the initial surface subsidence crater relative to

the Panel 1 undercut area, Henderson Mine, USA (after Stewart et al 1984) 9.4 ANALYSIS OF CHIMNEY CAVING AND PLUG SUBSIDENCE

9.4.1 Limiting Equilibrium Analysis

As discussed in Section 9.2 and by analogy with other areas of geomechanics, limiting equilibrium calculations might be expected to be helpful in estimating the ultimate collapse conditions when chimney caving develops by the first of the mechanisms discussed in Section 9.2. They should also be useful in analysing plug subsidence, although in this case an assumption of zero dilation on the vertical or near vertical slip surfaces should be made. However, limiting equilibrium analysis cannot be expected to produce reliable results when the second or unravelling chimney caving mechanism develops in strong discontinuous rock. The limiting equilibrium analysis presented here is that developed by Brady and Brown (1993). Consider the general case illustrated in Figure 9.10. The base and top of a vertically sided block coincide with the back of a cave and with the ground surface, respectively. It is assumed that rigid body motion of the block occurs by sliding vertically downwards under the influence of gravity when the shear resistance developed on the vertical block boundary is exceeded. This shear resistance will depend on the effective normal stress on the block boundary. This requires that assumptions be made about the distributions of normal stress and groundwater pressure with depth. It will be assumed that the vertical in situ stress is given by σzz = γz where z is the depth below surface and γ is the unit weight of the overlying material, and that there is an equal all-round horizontal normal stress of σxx = σyy = kγz where k is a constant.


357

Figure 9.10: General block geometry for limiting equilibrium analysis (Brady and Brown 1993)

If τ is the limiting vertical shear stress that can be developed on an element on the boundary surface of dimensions δp x δz to which an effective normal stress of σn' is applied, then the available shear resistance for the entire surface will be

∫∫ τ=z

0

p

0 p Q ddz (9.1)

where z and p are such that all surface elements are summed.

If W is the total weight of the block, then the factor of safety against shear failure on the vertical block boundaries is simply

WQ F = (9.2)

The complexity of the manipulations involved in obtaining solutions for particular cases will vary with the geometry of the block, the groundwater pressure distribution and the shear strength criterion used. If u(z,p) is the groundwater pressure at an element and an effective stress Coulomb shear strength law is used, Equation 9.1 becomes


358

( )[ ]{ } pd zd tanpz,u - zk c Qz

0

p

0φ′γ+′= ∫∫ (9.3)

where c' and φ' are the effective cohesion and effective angle of friction, respectively.

If an effective stress form of the Hoek-Brown non-linear strength criterion (Hoek and Brown 1980, 1997) with the parameter a = 0.5 is used, Equation 9.1 becomes

( )( ){ } pz - pz,u- Q tm

p

0ddA B

nB1

ci

z

0σσσ= −∫∫ (9.4)

where the peak strength of the rock mass, or of the shear surface, is given by

( )0.5' ' 21 3 ss b c cm σ σσ σ= + + (9.5)

or τN = A(σN – σtN)B (9.6)

in which σ1S' is the major principal effective stress at peak strength, σ3' is the minor principal effective stress, mb and s are constants that depend on the properties of the rock and the rock mass quality as measured by the RMR, σc is the uniaxial compressive strength of the intact rock material, A and B are constants depending on the value of mb, and τN = τ/σc

σN = σn/σc

( )21

2b

btN 4sm -

2m σ +=

An example of general interest is shown in Figure 9.11a. A block of width a and base length b has one pair of sides oriented in the direction of the strike of an orebody and the other pair in the dip direction. The dip is α, the maximum height of the block is h, and the water table is at a distance d below the ground surface. The groundwater pressure is assumed to be zero at the


359

water table, to have an initially hydrostatic rate of increase with depth, to be zero at the stope hangingwall, and to have an infinite rate of decrease with increasing depth at this point. The skewed parabolic water pressure distribution shown in Figure 9.11b satisfies these boundary conditions. For this distribution, the maximum water pressure is z'γw/2, and the total water pressure force generated over a depth z' = z – d, is z'2γw/3.

Figure 9.11: (a) Rectangular block geometry, and (b) assumed water pressure distribution for limiting equilibrium analysis (Brady and Brown 1993)

In this case, it is necessary to evaluate the shear resistance developed on each of the four vertical faces. The total shear resistance is then given by

Q = 2QBCGF + QDCGH + QABFE (9.7)

Brady and Brown (1993) give the complete solutions for the cases in which the groundwater level intersects the up-dip face DCGH, intersects the stope hangingwall, and is below the block ABCDEFGH (ie d ≥ h). This last case is of particular interest here because it can generally be expected that in an operating block or panel caving mine, the groundwater level will have been drawn down to below the crown pillar or the caved material. In this case, substitution for the weight of the block, W, and the total shear resistance, Q, in Equation 9.2 gives the value of the factor of safety against shear failure of the block for a Coulomb shear strength criterion as


360

( )

( )( )

⎩⎨⎧

αα+

αφ′

+αγα+′

=cosb

sin b -h h sin b -2h tank

cos ab cos b ac2 F

22

1

( )⎭⎬⎫⎥⎦

⎤⎢⎣

⎡ α+α+

3 sin b sin b -h h

a2

22

(9.8)

For the case of an excavation with a horizontal back, α = 0 and Equation 9.8 reduces to

⎟⎟⎠

⎞⎜⎜⎝

⎛φ′+

γ′

⎟⎠⎞

⎜⎝⎛ +

= tankh c2 ab

b a F (9.9)

If the non-linear Hoek-Brown strength criterion is used, the evaluation of the integrals in Equation 9.4 can be cumbersome. However, solutions can be found readily for simplified cases. Brady and Brown (1993) show that for the simple case of a rectangular opening with a horizontal back and no groundwater pressures, the factor of safety against shear failure of the block is

( )

( ) ( )[ ]B1tm

B1tm2

B1c - -h k B 1abhk

A b a2 F ++

−

σσγ+γσ+

= (9.10)

When there is no water involved, normal stresses and shear strength parameters can be expressed in total stress terms as in Equation 9.10. A further simplification can be made by noting that in weak or caved rock masses, the tensile strength of the rock mass, σtm, is likely to be close to zero in which case Equation 9.10 reduces to

( )

( )B 1abA hkb a2

F B1

B-1c

BB

+γσ+

= − (9.11)

This result shows that the factor of safety decreases as the opening dimensions increase, the depth of the block decreases, k decreases, the rock mass quality decreases, the uniaxial compressive strength of the intact rock decreases and as the unit weight of the rock increases. Consider the hypothetical example of a block cave with approximately equal plan dimensions of 180 m having progressed to within 120 m of the surface. Take the unit weight of the weak rock mass to be 0.027 MN m-3 and the value of k in weathered rock near the surface to be 0.75.


361

It must be assumed that there is space available into which the caved rock or plug can fall if chimney caving or plug subsidence is to occur. Assuming zero water pressures and total stress conditions, substituting the assumed values of the parameters into Equation 9.9 gives

φ+= F tan c 823.0 (9.12)

At limiting equilibrium, F = 1.0. It is instructive to investigate the combinations of values of the Coulomb shear strength parameters, c and φ, for which equilibrium can be maintained, or failure can be expected to occur. Assuming that φ = 35o, tan φ = 0.70 and so for limiting equilibrium conditions (F = 1.0), the cohesion, c, must be 0.364 MPa which is a reasonable value for a weathered or a caved rock mass. If the friction angle is reduced to 30o, the value of the cohesion required to maintain equilibrium increases to 0.608 MPa. These illustrative calculations do not allow for a factor of safety of, say 1.3 or 1.5, usually used when limiting equilibrium calculations are carried out for mine design purposes.

Hoek (1989) carried out an analysis of the stability of a rectangular crown pillar which is similar to the analysis presented here. He allowed for the presence of a water table within the crown pillar and for the ratios of horizontal to vertical in situ stress to be different for the two pairs of vertical surfaces of the rectangular block. Hoek (1989) used the Hoek-Brown non-linear strength criterion and allowed the parameters involved in the solution, other than the plan dimensions of the pillar, to vary according to a normal distribution. He then carried out a Monte Carlo simulation to investigate the influence on the calculated factor of safety of the pillar height, the Rock Mass Rating or RMR, the material constant mi in the Hoek-Brown criterion, the intact rock strength σc, the water table depth and the two k factors.

It must be emphasised that Hoek’s analysis and that presented here will not apply to all mechanisms involved in chimney caving or the progression of a block cave to surface. However, the analyses are applicable to the ultimate collapse conditions in some mechanisms and provide a simple check on overall crown pillar stability.

Bétourney (1994) and Bétourney et al (1994) presented a limiting equilibrium analysis and some numerical modelling results for the progressive chimney caving mechanism illustrated in Figure 9.3 which they refer to as chimneying disintegration failure. They also analysed the case of up-dip chimneying in an inclined orebody. The approach used is to approximate successive internal failure surfaces by circular arcs which represent lines of active pressure. In the vertical caving case shown in Figure 9.3, each rupture line is composed of symmetric circular arcs intersecting at an apex which is a point of passive earth pressure. The equilibrium of each successive volume of rock is assessed using the method of slices developed for slope stability studies.


362

Following Bétourney (1994), Bétourney et al (1994) reported the results of their analyses as a relationship between stope span, rock mass cohesion and factor of safety (for a particular value of rock mass unit weight) as shown in Figure 9.12. They found that for the chimneying disintegration mechanism to occur, the block size and degree of interlocking, and hence the cohesion, of the rock mass needed to be low. Interestingly, Bétourney (1994) applied this approach in back-analysing the well known failures at the Brier Hill Mine (Rice 1934) and the Athens Mine as illustrated in Figure 9.5 (Allen 1934, Boyum 1961).

Figure 9.12: Relationship between stope span, L, rock mass cohesion, cm and factor

of safety, F, for chimneying disintegration stability analysis (Bétourney 1994).

9.4.2 Empirical Methods

In another study of crown pillar stability which is of interest in the current context, Carter (1992), Carter and Miller (1995) and Golder Associates (1990) plotted the rock mass quality as measured by the Rock Mass Rating (RMR) or the NGI Q index against the scaled crown span for a database of caved and stable crown pillars as shown in Figure 9.13. The heavy critical span line shown in Figure 9.13 divides stable from unstable cases and is represented by the equation

SC = 3.3 Q 0.43 x sinh 0.0016 (Q) (9.13)

where SC is the maximum scaled span for a given pillar beyond which failure may occur. The values of the scaled crown spans, CS , for the data plotted in Figure 9.13 were calculated from the expression


363

( )( )

5.0

Rs cos4.011 ⎪⎭

⎪⎬⎫

⎪⎩

⎪⎨⎧

θ−+ρ

=ST

S C (9.14)

where S is the span of the crown pillar, ρ is the specific gravity of the rock, T is the thickness of the crown pillar, θ is the dip of the orebody and SR is the span ratio or ratio of the span, S, to strike length, L, of the crown pillar. For values of Q < 50 or RMR < 80, the straight line portion of the curve shown in Figure 9.13 may be approximated by SC = 3.3 Q 0.43 (9.15)

Carter and Miller (1995) applied a probabilistic approach to the assessment of crown pillar stability based on the empirical chart. Although the empirical and probabilistic approaches are of interest and probably have some application to studies of caving to surface in block and panel caving mines, it is suspected that the database on which Figure 9.13 is based does not include crown pillars of the sizes usually encountered in caving operations. Consider the hypothetical example analysed by the limiting equilibrium method in Section 9.4.1 above. The scaled span for this case is calculated from Equation 9.14 as 19.1 m for T == 120 m, S = L = 180 m, θ = 90o and ρ= 2.7. From Equation 9.15, the value of Q required to maintain stability is calculated as 59 (or RMR = 81). Intuitively this may appear to be reasonable, but it represents a much stronger rock mass than the strength parameters satisfying limiting equilibrium conditions for the same case.

Figure 9.13: Crown pillar case records plotted as scaled spans against rock mass

quality (after Carter and Miller 1995)

Scal

ed c

row

n sp

an C

s (m

)


364

9.5 LIMITING EQUILIBRIUM ANALYSIS OF PROGRESSIVE HANGINGWALL CAVING

When the orebody is not massive and relatively steeply dipping, caving of only the hangingwall need be considered unless the footwall caves under the influence of a major discontinuity or shear zone (eg Henry and Dahner-Lindqvist 2000). In such cases, progressive caving on the hangingwall may result as mining progresses down-dip using caving methods of mining. Classic examples of this form of discontinuous subsidence are those at the Grängesberg and Kiirunavaara mines, Sweden, and Gath’s mine, Zimbabwe. This account of the limiting equilibrium analysis of progressive hangingwall caving is based on that of Brady and Brown (1993). (a) (b) (c)

Figure 9.14: Progressive hangingwall failure sequence with increasing depth of mining: (a) mining from outcrop; (b) failure of overhanging wedge; (c) formation of steep face; (d) development of tension crack and failure surface; (e) development of second tension crack and failure surface; (f) initial open pit; (g) development of tension crack and failure surface; (h) development of second tension crack and

failure surface; and (i) progressive failure with increasing mining depth (after Hoek 1974)

This form of subsidence often occurs when the near surface portions of the inclined orebody have been extracted by open pit methods. Alternatively, underground mining operations may commence in relatively weak, previously unmined ground at shallow depths and proceed down-dip in stages. Figure 9.14 shows the sequence of progressive hangingwall failure postulated by Hoek (1974) for the cases of mining from an outcrop (a – e) and from an open pit (f – i). It is


365

assumed that at each new stage of mining, a tension crack and a failure surface form in the hangingwall rock mass at a critical location determined by the strength of the rock mass and the imposed stresses. In some cases, mechanisms other than this may occur. For example, major discontinuities such as faults may provide preferential shear planes. Alternatively, if a major set of persistent discontinuities dips steeply in a similar direction to the orebody, toppling of the hangingwall rock mass may occur as illustrated in Figure 9.15.

Figure 9.15: Toppling of steeply dipping hangingwall strata (after Heslop and Laubscher 1981)

Hoek (1974) developed a limiting equilibrium analysis for predicting the progress, from a known initial position, of hangingwall failure with increasing mining depth. This analysis assumed a flat ground surface and full drainage throughout the caving mass. Brown and Ferguson (1979) extended Hoek’s analysis to account for a sloping surface and groundwater pressures in the tension crack and on the shear plane. The idealised model used for the extended analysis is shown in Figure 9.16.


366

Figure 9.16: Idealised model used in limiting equilibrium analysis of progressive hangingwall caving (after Brown and Ferguson 1979)

The variables involved in the analysis are:

A = base area of wedge of sliding rock mass c′ = effective cohesion of rock mass H1 = mining depth at which initial failure occurs H2 = mining depth at which subsequent failure occurs Hc = depth of caved material S = width of orebody T = thrust on failure plane due to caved material Tc = thrust on foot wall due to caved material U = water-pressure force on failure surface V = water-pressure force in tension crack W = weight of wedge of sliding rock Wc = weight of caved material Z1 = depth of initial tension crack Z2 = depth of subsequent tension crack Zw = depth of water in tension crack

α = dip of upper ground surface (shown positive in Figure 9.16 but may be negative)

γ = unit weight of undisturbed rock mass γc = unit weight of caved material


367

γw = unit weight of water θ = inclination of T to normal to failure surface '

nσ = effective normal stress on failure plane

τ = shear stress on failure plane φ′ = effective angle of friction of rock mass φw = friction angle between caved and undisturbed rock ψ0 = dip of orebody ψb = angle of break ψP1 = dip of initial failure plane ψP2 = dip of subsequent failure plane.

The analysis is based on the following assumptions: (a) Mining and caving occur for a large distance along strike compared with the cross-

sectional dimensions shown in Figure 9.16. As a consequence, the problem may be reduced to one of two dimensions. Calculations are carried out for unit thickness perpendicular to the plane of the cross section.

(b) The initial position of the hangingwall face is defined by known values of the

geometrical parameters H1, Hc, Z1, ψp1.

(c) The extent of caving at the new mining depth, H2, is defined by a tension crack which

forms to a critical depth and strikes parallel to the orebody. (d) Failure of the hangingwall rock mass occurs along a critical, planar, shear surface

whose location is determined by the strength properties of the rock mass and the imposed effective stresses.

(e) The hangingwall rock mass has homogeneous and isotropic mechanical properties. Its

shear strength can be defined by an effective stress form of Coulomb's criterion. (f) Water may enter the tension crack and seep along the potential failure surface into the

underground excavations, producing a triangular distribution of excess water pressure along the shear surface.

(g) In carrying out the limiting equilibrium calculations, simplified distributions of stress

within the caved and caving masses are used. In particular, the effective normal stresses and shear stresses acting on the failure plane are averaged using the methods of statics. Lupo (1997) subsequently modified this approach to include estimates of the tractions applied to the shear surface by the movement of broken rock during draw.


368

The development of the equations used in the solution is given in Appendix D which also sets out the stepwise sequence of calculations required to obtain numerical solutions. A simple iterative scheme is required to determine values of ψP2 and ψb for a given stage of mining. Example

Brown and Ferguson (1979) used the limiting equilibrium analysis to predict the progress of hangingwall caving at Gath's Mine, Zimbabwe. At several sections, the progress of caving had been monitored as mining had progressed down-dip from the 99 level to the 158 level (Figure 9.17). Using the dimensions and problem idealisation shown on Figure 9.17 with γ = 28 kN m-3, γc = 25 kN m-3, α = 0, U = V = 0, φw = 35°, and c′ = 200 kPa and φ′ = 40° determined from Bieniawski's geomechanics classification scheme, values of Z1 = 31.3 m, ψP2 = 61.4° and ψb = 66° were calculated for mining in increments from the 99 level to the 158 level. These values agreed remarkably well with field data, and provided some confidence in the applicability of the method in this case.

Using the 158 level as the starting point, successive calculations were then made of the locations of the tension crack and shear surface as mining advanced down-dip at this section. The results obtained are shown in Figure 9.17. The results were found to be sensitive to the value of Hc at the beginning of each mining lift. The results shown in Figure 9.17 were obtained using values of Hc that were determined from calculations of the volumes of caved materials and estimates of the volumes of material dawn historically and likely to be drawn in the future.

In practice, the geometry of the problem is rarely as simple as that used in the model. Care has to be taken in assigning values of geometrical parameters such as orebody dip and width and depth of caved material, and of rock mass properties. In the case of Gath's Mine, the real problem also involved the use of backfill in the cave (which served to steepen the angle of break), a three-dimensional effect at the end of the mined section of orebody, the influence of a hill, the presence of a pillar designed to protect the mill area, and the influence of a major set of geological discontinuities which controlled cave development at one location. The effects of some of these factors could be assessed by sensitivity studies using the analysis, but assessing the influence of others had to remain a matter of judgement and experience.


369

Figure 9.17: Idealisation of progressive hangingwall caving, Gath's Mine, Zimbabwe

(after Brown and Ferguson 1979) 9.6 SUBSIDENCE PREDICTION IN PRACTICE

9.6.1 General Approach

In order to apply the concepts and techniques outlined earlier in this chapter to predict surface subsidence, the extent of the subsidence crater, and the zone of influence of the cave in a given case, several steps are necessary. Generally, two different classes of ground deformation are involved – the large deformations within the caved or subsided zone, and the smaller deformations around the cave and within its zone of influence both underground and on the surface. It is also usually necessary to determine the times at which the various ground movements will occur with respect to the progress of mining. Based on experience in a number of block caving mines in South Africa, Butcher (2002b) has outlined the steps required in making these predictions. Experience suggests that, in a block or panel cave, all of the ground above the cave footprint will cave. There will usually then be a zone outside the approximately vertical orebody contact that will break back to produce the subsidence zone as defined by the angle of break or subsidence (see Figure 9.19). The general steps suggested by Butcher (2002b) to make these predictions are: 1. Project the footprint contact perimeters on all levels to surface to establish the area that

clearly will be destroyed by caving. 2. Make a preliminary estimate of the angle of break using Laubscher’s empirical method

based on MRMR values (Laubscher 1994). 3. Calibrate this estimate against observed angles of break in this or similar mines.


370

4. Check the estimated angle of break using other methods of analysis appropriate to the likely failure mechanism such as that developed by Hoek (1974) and extended by Brown and Ferguson (1979).

5. Modify the current estimate of the angle of break to take account of local geological features such as faults or discontinuities that are likely to control step-like slope failure modes.

6. Use numerical modelling (using the FLAC code, for example) to check the angle of break and to estimate stresses and displacements induced in the rock mass around the caved zone.

These steps are similar to those used in making subsidence predictions at CODELCO-Chile’s block and panel caving copper mines (Flores and Karzulovic 2002a, Karzulovic et al 1999). The application of this approach will be illustrated in the following section. 9.6.2 Prediction of Caving Induced Subsidence at Rio Blanco and El Teniente Mines,

Chile

In order to improve the prediction of the inevitable surface subsidence or crater associated with block caving at the Rio Blanco Mine, in 1991 the Andina Division of CODELCO used a limiting equilibrium analysis similar to that developed by Brown and Ferguson (1979). The “theoretical” value of the angle of break so calculated was adjusted to take account of local factors such as the presence of faults and the amount of broken material in the crater, and to obtain improved agreement with a database of observed values for rock masses of similar quality. The results were presented as a chart of the angle of break against height above the crater floor for rock of a given type or rock mass quality. This general method was then extended and applied to predict the evolution of the horse-shoe shaped subsidence crater at the larger El Teniente Mine (Figure 9.18). This extension of the Rio Blanco analysis introduced a FLAC analysis to assist in defining the zone of influence of the cave adjacent to the crater walls. The horizontal displacements predicted by the FLAC model were calibrated against the available observations made in drifts which had shown subsidence-induced damage. In this way, it was possible to define threshold values of horizontal displacement likely to be associated with damage. The variables involved in this approach are defined in Figure 9.19. The chart of angle of break, α, against height above the crater floor, H, for the Ten 4 – Regimiento Sector at El Teniente is shown in Figure 9.20. It will be noted that the angle of break is higher for the Braden breccia than for the other rocks and that it decreases markedly with increasing height above the crater floor. The chart developed to estimate the zone of influence of the cave on the highwall side of three production sectors on the El Teniente Ten 4 level, is shown in Figure 9.21. It will be noted that for heights above the crater floor of 500 m or more, the lateral extents of the zones of influence, ti(z), of the three production sectors represented in Figure 9.21, are in the range 50-70 m. Rojas et al (2001) give a similar chart to Figure 9.20 for the El Teniente Sur and Esmeralda sectors.


371

PIPAPIPABRADENBRADEN

QuebradaQuebradaTenienteTeniente

TenienteTeniente4 4 FortunaFortuna

Teniente 4Teniente 4RegimientoRegimiento

Teniente 5Teniente 5PilaresPilares

TenienteTeniente3 3 IslaIsla

TenienteTenienteSub 6Sub 6

TenienteTeniente4 Sur4 Sur

NN

Figure 9.18: Plan view of the subsidence crater at the El Teniente Mine, Chile, in 1998 showing the Braden Pipe and the locations of the production sectors (Karzulovic et al

1999)


372

Figure 9.19: Geometrical parameters used in the analysis of subsidence, El Teniente

Mine, Chile (Karzulovic et al 1999)

H

α βh

AB

ASDC

tC

zti

SURFACE TERRAIN

CRATER WALL

BROKEN MATERIAL

DRIFT AFFECTED BY SUBSIDENCE

H

α βh

AB

ASDC

tC

zti

SURFACE TERRAIN

CRATER WALL

BROKEN MATERIAL

DRIFT AFFECTED BY SUBSIDENCE

H Crater depth (crater wall height)

h Average height of the column of broken rock

A b Area of the base of the crater

A S Area of the crater at surface

α Angle of break

β Angle of influence (NOT a good definition of the zone of influence)

ti Extent of the zone of influence at an elevation z above the base of

the crater

tc Extent of the cracked zone at surface

Dc Distance to the main cracks


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Figure 9.20: Curves developed to estimate the angle of break, Ten 4 – Regimiento Sector, El Teniente Mine, Chile (Karzulovic et al 1999)


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Figure 9.21: Plots developed to estimate the zone of influence of the cave, highwall side, Ten 4 - Regimiento Sector, El Teniente Mine, Chile (Karzulovic et al 1999)

( ) C z zt xxi δ=

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CHAPTER 10

MAJOR OPERATIONAL HAZARDS 10.1 SCOPE

he planning and operation of a block or panel caving mine involves a number of major risks or hazards. As has been noted previously, caving methods of mining require the investment of significant amounts of capital in mine development before production

can commence and a cash flow can be generated. Caving methods of mining are relatively inflexible so that, if a mistake is made, it is not usually either a straight forward or an inexpensive matter to change the mining method. In the International Caving Study Stage I, a risk analysis approach known as CaveRisk was developed. An account of CaveRisk will be given in Chapter 11. CaveRisk has been developed in a number of modules dealing with the major areas of risk in caving operations – caveability, fragmentation, excavation stability and caving performance. These topics have been the subject of earlier chapters in this book. Within each of these major areas of risk, there are a number of specific issues which, in themselves, constitute risks or hazards that must be addressed in mine planning and operation. In a wide-ranging paper presented to MassMin 2000, Heslop (2000) discussed the risks involved in block caving under several headings: • geological risks; • geotechnical risks; • valuation risks; • operating risks; • environmental risks; • financial risks; and • political risks.

T

Chapter 10: Major Operational Hazards

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Heslop (2000) divides the area of operating risk into four further categories: 1. Operational hazards – rock bursts, air blasts, mud rushes and water and slurry inrushes that

could result in loss of life and/or premature mine closure. 2. Design risks – those risks that have an economic impact and are the result of incorrect

assessments of the ground conditions or effects of stress etc based largely on geotechnical data collected years before.

3. Draw risks – those risks that have an impact on the current and future ore and grades that

will be recovered and are the result of incorrect assessments of the issues that influence draw.

4. Automated equipment risks – risks arising from an over-reliance on advanced technology to

achieve critical levels of performance from LHDs and drills. The areas of design and draw risk were discussed in Chapters 2 – 7 while Chapter 8 was concerned largely with the management of those risks through the use of monitoring. The areas of operational hazards and automated equipment risks were not part of the program of research carried out under the International Caving Study Stage I. Because they have geotechnical causes, and because preventing their occurrence and/or minimising their impact is of vital importance to the safe and efficient operation of caving mines, the four operational hazards defined by Heslop (2000) will be the subject of this chapter. Because they are not the only possible cause of air blasts, and because they can cause loss of life and put production at risk in their own right, a fifth form of operational hazard, major collapses, will be added to Heslop’s list for present purposes. Figure 10.1 shows a classification and some inter-relationships of these five major operational hazards. This figure is similar to, but differs from, Figure 11.8 which shows the same hazards in the form of a logic tree in which main and secondary processes are identified.

Figure 10.1: Classification of major operational hazards


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A significant difficulty encountered in assembling this less than comprehensive account of these five types of operational hazard was one of terminology. It appears that local terminology has developed in a number of countries, companies and mining districts. This terminology is not always clearly defined or transparent to the outsider. Heslop’s (2000) terminology will be used here, and the various terms will be defined as they are introduced. Accounts will be given of the nature, causes, effects and, where possible, of the measures used to avoid or ameliorate the effects of, each of the major operational hazards to be discussed. In practice, these hazards should be addressed using a risk assessment and management approach of the type to be discussed in Chapter 11. 10.2 MAJOR COLLAPSES

10.2.1 Terminology

In the present context, major collapses will be taken to include occurrences such as: • the uncontrolled collapse of crown or sill pillars to surface or to a mined-out overlying void

(type 1); • uncontrolled falls of large blocks or volumes of rock from the back of the undercut or,

more usually, the cave (type 2); and • the collapse, progressive or otherwise, of excavations on and above the extraction level

(type 3). An important corollary of this definition of a major collapse is that the event must cause damage of some consequence to the operation. In extreme cases, it could result in loss of life, loss of production from a panel or block, or extensive infrastructure damage. In less extreme cases, it could cause expensive delays in production or involve the need for remedial work. The categories of major collapse listed above all relate to the undercut and extraction levels or to the cave itself. It is possible, but not likely, that major collapses having life or production threatening consequences could also occur in mine accesses and other items of infrastructure. Flores (1993) and Diaz and Tobar (2000) report a more specific use of the term “collapse” at the El Teniente mine, Chile. In this sense, a collapse is defined as any undesired fall or slip of rock into an underground excavation. A collapse may have total, partial or no structural control and may occur at a range of speeds from very gradually to very rapidly (Flores and Karzulovic 2002a). A collapse involves failure of the rock mass over a significant area, usually on the extraction or undercut levels, resulting in the worst case in the total closure of drifts. The use of the term outlined above includes, but is not restricted to, collapses of this type. The more specific use of the term at El Teniente is intended to distinguish this type of collapse from those resulting from rock bursts as discussed in Section 10.3 below.


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10.2.2 Causes

Collapses of the first type listed in Section 10.2.1 and illustrated schematically in Figure 10.2, are the most dramatic of the major collapses to be considered here. They result when the height of the crown or sill pillar reduces to such an extent that failure can occur by one of a number of mechanisms including progressive unravelling, chimneying, buckling or snap through instability and, most commonly in block and panel caves, shear failure on the vertical or sub-vertical boundaries of the crown or sill pillar. The analysis of the latter type of failure by the limiting equilibrium method was discussed in Section 9.4.

Figure 10.2: Major collapse, type 1, involving collapse of the entire crown pillar If occurrences of these types are to result in the deleterious consequences associated with the definition of major collapses being used here, it follows that there must be a void into which the collapsing mass can fall under the influence of gravity. In a block or panel cave, such a void will exist only if it has been created by the inappropriate drawing of caved material. This can result from a mismatch between draw and caving rates. Such circumstances have arisen in a number of cases when the cave has not propagated in the desired manner and caved ore has been drawn to maintain production. The massive collapses occurring in these cases have usually produced air blasts as will be discussed in Section 10.5. Examples of major collapses of this type are discussed by Bell (1999), de Nicola Escobar and Fishwick Tapia (2000), Laubscher (2000) and Kendrick (1970). Major collapses of the second type occur when a block or large volume of rock is isolated by the surface of the undercut or cave back and discontinuities such as faults and/or fractures caused by the induced stresses, and falls or slides under the influence of gravity. Figure 10.3 illustrates a possible occurrence of this type. As shown in Figure 10.3, these failures and


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collapses of the rock mass are more likely to occur when the back of the undercut or cave is convex downwards than when it is concave. Production and rock mechanics engineers are accustomed to dealing with major failures of this general type in other forms of underground mining involving large excavations. A difficulty experienced in block and panel caving mines is lack of the access to the rock mass and the opportunity to inspect it and take remedial measures that exists in other forms of mining.

Figure 10.3: Major collapse, type 2, involving falls of large blocks from the cave back

Collapses of the third type on and around the extraction level are stress-induced but, as illustrated in Figure 10.4, they may be exacerbated by the presence of faults or other major and persistent discontinuities. Collapses of this type have long been experienced in caving mines (Julin 1964, Flores 1993), usually ahead of the cave as a result of the stress abutment produced ahead of the undercut. They are sometimes referred to as “weight” problems. As Flores (1993) and Diaz and Tobar (2000) point out, such collapses are usually progressive rather than sudden and violent. As shown in Figure 10.4, they almost invariably involve failure of the pillars left on and above the extraction level. They may take several weeks or months to develop fully and once initiated are difficult to arrest by the provision of support or reinforcement. In the worst cases they may result in the closure of drifts and drawpoints. The levels of stress induced in the undercut and extraction levels under a range on mining conditions were discussed in Section 5.5.


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Figure 10.4: Major collapse, type 3, involving failures on and around the extraction level (after Flores 1993)

Karzulovic et al (1991) give an instructive example of a slow, structurally controlled collapse on the extraction level in Panel II of the Rio Blanco Mine, Chile. The orebody is in andesite in what is known locally as secondary ore having an RMRL value of about 40. A large pyramidal key block having an approximately rectangular base with dimensions of 30 m by 30 m and an estimated height of 130 m was formed by three major geological structures, the cave front and the floor of the undercut level. Before undercutting, the block was stable under the influence of the confining stresses acting on its inclined faces. However, the process of undercutting truncated the block to a height of 18 m and relieved the confining stresses making the base of the block unstable. This applied excessive dead-weight load to the pillars around the extraction level, some of which had been weakened by the presence of an old exploration drift. The unstable key block failed slowly, producing complete collapse of the production and drawpoint drifts in the affected zone. The problem was back analysed successfully using block theory (Goodman and Shi 1985). Block theory is now used routinely by CODELCO-Chile’s Andina Division to check for the presence of potentially unstable key blocks and to plan the caving process in such a way that block stability problems are minimised (Flores and Karzulovic 2002a).


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10.2.3 Effects

The effects of major collapses of the types discussed above are obvious, varied and sometimes disastrous. Many of them have been described by Laubscher (2000) and include: • uncontrolled caving to surface or to overlying mined out areas (type 1); • provision of a means of entry of water, waste, mud or tailings into the mine (type 1); • air blasts associated with the collapse (types 1 and 2), potentially resulting in the loss of

life; • arrest of cave propagation (type 2); • the imposition of excess loads on the extraction level pillars as a result of large blocks of

rock “sitting down” on the major apex (types 2 and 3); • hangups in or above drawpoints as illustrated in Figure 10.5, leading to the need for

secondary blasting and the loss of production (type 2); • damage to, and possible closure of, extraction level drifts and drawpoints, leading to the

loss of production (types 2 and 3); and • the need for expensive and time-consuming remedial work and/or a change to the mining

plan or sequence (type 3).

Figure 10.5: Hangups in and above drawpoints (after Laubscher 2000)

10.2.4 Prevention and Amelioration

Just as the effects of these major collapses are many and varied, so are the means of preventing them and of ameliorating their effects. Generally, but not invariably, they can be prevented by careful planning, design and operation of the mine. The most important preventive and ameliorative measures include:


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• strictly controlling the amounts and rates of draw according to the principles discussed in Chapter 7 (types 1 and 2);

• maintaining an adequate cover of caved ore over the drawpoints according to the principles to be discussed in Section 10.5.4 (types 1 and 2);

• ensuring that any openings that can be identified as having significant risk of exposure to inrush or air blast are safeguarded appropriately;

• suitably shape the undercut and major apices as discussed in Chapter 5 (type 2); • use an undercutting strategy (see Section 5.2) that will reduce extraction and undercut level

stresses to manageable levels under the prevailing circumstances (type 3); and • protect the extraction level excavations by appropriate pre-reinforcement as discussed in

Section 6.5 (types 2 and 3). 10.3 ROCK BURSTS

10.3.1 Terminology

Rock bursts have been defined classically as the uncontrolled disruption of rock associated with a violent release of energy additional to that derived from falling rock fragments (Cook et al 1964). More recently, a rock burst has been defined more simply as “a seismic event which causes violent and significant damage to tunnels and other excavations in the mine” (Ortlepp 1997). As this definition indicates, rock bursts are a sub-set of a broader range of seismic events which arise from conditions of unstable equilibrium within the rock mass and involve the release of stored strain energy and the propagation of elastic waves through the rock mass (Brady and Brown 1993). They may or may not produce damage to mining excavations. Seismic events in mines are commonly characterised by parameters originally developed in earthquake seismology. Source parameters such as energy output, seismic moment, source radius and radiated energy may be related to the damage sustained by mining excavations. A commonly used measure of event magnitude in earthquake seismology is the Richter magnitude which is a linear function of the logarithm of the seismic moment which is obtained from the spectral analysis of the body waves radiated from the source (Brady and Brown 1993). 10.3.2 Causes

The release of stored strain energy which produces rock bursts may have one of two causes – unstable slip on a pre-existing plane of weakness, usually a fault, and unstable brittle fracture of intact rock. In either case, energy which causes damage to the mine excavations is released dynamically into the rock mass. The resulting damage to the mine structure may occur some distance from the initial site of slip or fracturing. Table 10.1 summarises the main types of seismic event observed in underground mines and the associated source mechanisms. It also


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summarises the factors determining the intensity of the seismic impulse, the factors influencing the site response to the event, and the nature of any resulting rock burst (Ortlepp 1997). Table 10.1: Types of seismic event and damage mechanisms in underground mines

(Ortlepp 1997)


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Two separate conditions must be satisfied in order to generate rock burst conditions: • the induced stresses must be high enough to overcome the strength of the fault or the rock

mass (in whatever mode of failure applies); and • the resulting slip or fracture must be mechanically unstable, releasing energy that cannot be

absorbed in the processes of slip or fracture themselves. This implies a particular relationship between the stiffnesses of the loading system and the fault or rock mass whose failure gives rise to the rock burst. It means that rock bursts will usually occur only in strong, brittle rocks.

The mechanics of unstable slip or fracture in an underground mining environment have been analysed and discussed by Jaeger and Cook (1979), Salamon (1983), Brady and Brown (1993) and Ortlepp (1997), for example. A number of indicators of rock burst potential have been used. One of the first was the Energy Release Rate (ERR) or the spatial rate at which energy is released by the extension of a mining excavation (Cook et al 1964). If the ERR exceeds the rate at which the released energy can be absorbed non-violently, then a rock burst hazard might be expected to exist. Good correlations between ERR and rock burst incidence were observed in the deep level gold mines of South Africa from the 1960s and so this approach found application in combating the rock burst hazard (eg Dempster et al 1983). Subsequently, an alternative criterion known as the Excess Shear Stress (ESS) was introduced by Ryder (1987) for the case in which fault slip is the initiating mechanism. Figure 10.6 illustrates the phenomenon of unstable slip and the concept of ESS which is the difference between the limiting static and dynamic strengths of the fault at the prevailing normal stress. It may be calculated from the results of stress analyses, although account must be taken of the effects of progressive fault displacement as mining progresses if accurate results are to be obtained (Brady and Brown 1993).


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Figure 10.6: Unstable slip on a fault

10.3.3 Effects

The best known occurrences of rock bursts are those experienced in strong, brittle rock in the deep level gold mines of South Africa and in the Kolar Gold Fields of India. However, rock bursts can also occur in a range of other mining environments. Their effects include violent fracture and possible dynamic expulsion of the rock from and near the boundaries of excavations, and the full or partial closure of mining excavations including accesses. The damage can be localised or cover a significant area and can lead to a loss of life. In principle, there is no reason to expect that rock bursts will not occur in caving mines if the conditions for their occurrence are satisfied. As was noted in Chapter 1, there is now an increasing tendency for block and panel caving methods of mining to be used in rocks that are stronger and more brittle, and under higher stresses at greater depth, than those in which caving methods have been applied traditionally. The high percentages of extraction existing on and around extraction levels, and the stress abutments associated with undercut and caving fronts, suggest that the high stress conditions required to initiate bursting could occur in caving mines. The major experience of rock bursts in block and panel caving mines has been in the strong primary ore at the El Teniente mine, Chile. As indicated in the overview of this mine given in Section 1.3.2, the extremely high lateral in situ stresses associated with the nearby subduction zone, have had a major influence on the mining of this orebody. Rock bursts were first experienced on the El Teniente 4 level in 1976 (Alvial 1992) and have been a continuing problem since 1982 when mechanised panel caving began in the primary ore which is stronger and stiffer than the secondary ore that had been mined to that date. The rock bursts experienced at El Teniente have had a range of intensities and effects which can be localised or cover large areas on a given level. They occur in the main in the pillars on and


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around the extraction level. The damage can range from localised spalling in a drift, to closure of one or more drifts, and to the complete destruction and closure of a part of the mine. Clearly, rock bursts causing such damage can also destroy equipment and installations and lead to a loss of life. Figure 10.7 illustrates the range of damage caused by rock bursts at El Teniente. One of the first massive rock bursts destroyed an area of about 18 000 m2 on an extraction level in Teniente 4 developed 12 m below a former extraction level which had been severely damaged previously (Kvapil et al 1989).

Figure 10.7: Rock burst damage, El Teniente Mine, Chile

(a) category 1, (b) category 2, (c) category 3, and (d) category 4 (Kvapil et al 1989) The history of mining and combating the rock burst hazard in the El Teniente Sub 6 sector (the location of which is shown in Figure 1.10), provides a most instructive case history and will be discussed in some detail here and in Section 10.3.4 to follow. Production from Sub 6 started in mid-1989 using a conventional (post-undercutting) panel caving method. Six months later, the Sub 6 sector experienced induced seismicity and associated rock bursts. Rojas et al (2000a) provide the following history of subsequent rock bursts and the associated damage: • On 18 January 1990, a 3.6 Richter magnitude event generated wide-spread damage,

notably to a temporary transportation drive. Operations on Sub 6 had to be stopped as a consequence.


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• On 2 July 1990, a 3.2 Richter magnitude event caused so much damage that all mining activities on Sub 6 were stopped for nine months. Systematic repair of all damaged areas was carried out during the stoppage.

• On 23 May 1991, the area experienced a sequence of seismic events initiated by a 4.0

Richter magnitude event. The previously reinforced area was damaged to varying degrees. Production was stopped for a further five months. During this time some of the strategies used to successfully combat the rock burst hazard were developed. Monitoring using a global digital seismic network began in January 1990.

• On 25 March 1992, a 3.7 Richter magnitude event severely damaged the main access to the

Sub 6 production area. 10.3.4 Prevention and Amelioration

Experience in over a century in South Africa has shown that the occurrence of specific seismic and rock burst events are impossible to predict, and almost impossible to prevent, in an area known to be rock burst prone. This has also been the experience in block and panel caving mines that are susceptible to rock bursts. However, it has proven possible to reduce the frequency and severity of rock bursts, and to limit their effects, by a prudent combination of mining strategies (most notably undercutting strategies in caving mines), spatial and temporal rates of mining, monitoring and the provision of support and reinforcement (see, for example, Ortlepp 1997, Rojas et al 2000a, Salamon 1983, Stacey and Ortlepp 2000). Automation of equipment can also reduce the risk of exposure of operators to the hazard (Moyano and Vienne 1993). The examples given here will again relate to the El Teniente Sub 6 sector as reported by Rojas et al (2000a). The experience at El Teniente has been that rock bursts are most likely to occur in the early stages of mining a new block, panel or sector before the cave has connected to surface or to an overlying caved zone. During this period, the incidence and effects of rock bursts can be controlled by measures to be outlined. These measures may not always be effective in preventing all local scale seismicity associated with the interaction of excavation geometry, induced stresses and rock mass properties. In the case of El Teniente Sub 6, rock bursts appear to have been associated with rock fracture rather than with unstable slip on pre-existing discontinuities. However, in other cases rock bursts have been associated with slip on pre-existing discontinuities (Ortlepp 1997). An experimental production plan aimed at ameliorating the rock burst hazard was implemented in the Sub 6 sector from the beginning of 1994 and had the following features: • the production area was limited to a previously undercut area and to an area with a shorter

column of ore (240 m) to the overlying caved zone on Teniente 4 than in other areas;


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• the spatial and temporal rates of production were kept as uniform as possible to minimise stress concentrations and the sizes of rock fracture zones;

• extremely low production rates were used initially and then gradually increased; • ore was extracted using only remotely controlled equipment; and • the extraction rate was limited according to a criterion based on a time-weighted average of

seismic activity over the preceding two weeks. As shown in Figure 10.8, fewer rock bursts were experienced in this period of experimental mining than had been the case in the initial period of mining (see Section 10.3.3). At the end of the experimental mining stage, there was a transitional stage which included an undercutting trial before full production was resumed early in 1998. Since 1997, only a few minor rock bursts producing minor damage have been experienced (Rojas et al 2000a).

Figure 10.8: Progress in reducing rock burst incidence, El Teniente Sub 6 (Rojas et al 2000a)

10.4 MUD RUSHES

10.4.1 Terminology

Mud rushes are sudden inflows of saturated fines from drawpoints or other underground openings. Mud rushes may be sub-divided into two groups depending on the source of the mud. External mud rushes are those in which the mud is produced externally to the underground cave mining environment. External mud rushes are produced from three main


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sources – tailings or slimes, back fill placed underground and open pit slope failures. Internal mud rushes are those in which the mud is produced by the comminution of shale or other clay-forming country rocks and clay mineral-rich ores within the cave muck pile. There may also be a category of mixed mud rushes in which the mud arises from a combination of internal and external sources (Butcher et al 2000). The term mud rush as defined above is used particularly in South Africa. In some other mining areas such as Indonesia and The Philippines, the term wet muck flow is used to describe a similar but not identical phenomenon in which there is a sudden collapse and rapid run-out of wet granular material following some disturbance. Such flows come from the drawpoint and may occur during active draw or at other times (Hubert et al 2000). Key features of both mud rushes and wet muck flows as defined here, are that their occurrence is sudden and that the muck or mud is liquefied by some means and flows rapidly. 10.4.2 Causes

Butcher et al (2000) note that four conditions are required for a mud rush to occur: • potentially mud-forming minerals must be present in the cave; • water must be present; • there must be a disturbing or triggering mechanism such as draw, blasting or seismicity;

and • there must be discharge points through which the mud can enter the mine workings. In Freeport Indonesia’s Intermediate Ore Zone block cave, wet muck flows can occur when there is more than 20% sand sized material in the ore, the water content is greater than 8.5% and the ratio of the field dry density to the maximum dry density is less than 0.9 (Hubert et al 2000). Butcher et al (2000) distinguish primary internal mud rushes from two secondary mud rush mechanisms, rapid muck pile compaction (involving the development of high pore water pressures) and reduced muck pile/waste capping drainage. The reduced drainage mechanism relates to the second requirement for the occurrence of mud rushes, the presence of water. In this case, the presence of mud reduces drainage through the muck pile so that more water is retained in the muck pile and becomes available to feed mud rushes. Water balance calculations will provide a good indication of whether or not water is accumulating within the cave. Figure 10.9 shows a hypothetical example of the pumping history of a block caving operation from block development through full production to the draw down of reserves. In this case, assuming a constant supply of water into the block, the retention of water in the muck pile is indicated.


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Figure 10.9: Hypothetical mine pumping history showing muck pile water retention, assuming a constant quantity of water reporting to underground pumps (Butcher et al

2000)

Water may enter the draw column from a number of sources including: • surface run-off; • the direct entry of rain water through the subsidence zone or open pit; • melting snow collected in the open pit or subsidence zone (Torres et al 1981); • groundwater drained through an overlying void, the country rock or major geological

features such as faults (Hubert et al 2000); • surface reservoirs or lakes; • tailings or slimes placed near or in the open pit or subsidence zone; • surface or groundwater entry through unsealed boreholes; • leakage from underground water reticulation or storages; • hydraulic fill placed elsewhere underground; and • inrushes from water-filled abandoned workings. Obviously, one of the major means of avoiding mud rushes is to control the supply of water as discussed in Section 10.4.4 below. The third requirement for the occurrence of mud rushes is some form of disturbance to act as a trigger mechanism. The most obvious form of disturbance is the drawing of ore which will disturb the pre-existing conditions of equilibrium and allow the mud within the muck pile to flow. A high rate of extraction from a single drawpoint may increase the possibility of a mud rush occurring because any mud-bearing waste cap enters the draw column sooner and because larger quantities of water are likely to enter a more highly drawn zone. In essence, the more heavily drawn drawpoints provide preferred mud pocket pathways through the ore column.


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A different form of disturbance or trigger is provided by the vibrations arising from blasting, rock bursts, earthquakes, major falls of ground and the nearby operation of large pieces of equipment. The vibration, repeated or dynamic loading of saturated or almost saturated mud in the draw column or a drawpoint, may generate high pore water pressures which will reduce the shear strength of the mud to almost zero and permit it to flow. This is the classic mechanism of liquefaction in soils which may also occur in hydraulic fill (Mitchell 1983). It may also be caused by increased static or quasi-static loads as in the rapid muck pile compaction mechanism described above. The fourth requirement for the occurrence of mud rushes is a discharge point through which mud can enter the mine workings. In block and panel caving mines, drawpoints provide the obvious discharge points. The wet muck flows reported at the Freeport Indonesia mines (Barber et al 2000, Hubert et al 2000) are said to originate entirely from within drawpoints. However, if mud is liquified, it may be able to flow into other excavations such as exploration drifts, undercut level development or open drill holes connecting with the cave. While not resulting from exactly the same mechanisms as mud rushes and wet muck flows, clay or mud in the drawn ore can cause problems during loading and transportation and in the drawpoints. Torres et al (1981) report that at the Rio Blanco mine, Chile, ore containing clayey fines with water contents of between 4% and 8% caused hangups in the ore passes. When the water content was above 8%, the material became fluid introducing flow control problems. These problems, particularly those occurring at the lower water contents, were more easily controlled with LHD loading that when a grizzly system had been used. 10.4.3 Effects

Depending on the volume of mud involved, a mud rush or wet muck flow can cause severe damage as well as loss of life. The rapidity of mud rushes means that there is little warning and little chance of escape for personnel or equipment in their direct paths. Many fatalities have occurred historically as a result of mud rushes (Hunt and Daniel 1952-53, Butcher et al 2000). As well as the direct damage to equipment and infrastructure and the potential fatalities caused by mud rushes, they can cause loss of production and impose significant clean-up costs. But perhaps the most severe “indirect” effects of mud rushes are the air blasts that can result from the rapid compression of the air in the drifts into which the mud suddenly flows. The damage and loss of life associated with air blasts can be extreme (see Section 10.5). 10.4.4 Prevention and amelioration

It was noted in Section 10.4.2 that four conditions have to be satisfied if mud rushes are to occur in a given mine. They include the mud forming potential of the orebody and the country rocks and the availability of water. Mud rush potential should be evaluated at an early stage in


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the risk assessment of a potential block or panel caving operation using the techniques to be discussed in Chapter 11. On the basis of this risk assessment or of mining history in this or similar nearby mines, it should be possible to establish whether or not an operation is likely to be mud rush prone (Butcher et al 2000). The risk of mud rushes actually occurring in circumstances identified as mud rush prone can be reduced by a series of design and operational measures as discussed by Butcher et al (2000) and Laubscher (2000). The first steps to take are to ensure that there is no avoidable ingress of mud forming materials or water into the cave. This can be done through a range of design and operational measures. Design and siting of tailings dams

The incorrect design and siting of tailings dams has been a significant contributing factor to mud rushes. Obviously, tailings, slimes or other wastes that could behave as a fluid or provide a source of mud, should not be disposed of into open pits or subsidence zones above caving operations, or in areas likely to be affected by developing subsidence. Nor should they be disposed of where there is an available flow path into the underground workings. Tailings and slimes dams and their foundations should be designed and sited according to the best geotechnical practice to avoid failure and flooding of the mine. Tailings dam failures have occurred in apparently unlikely circumstances and their design should not be taken lightly (eg Morgenstern 1996, Olalla and Cuellar 2001). Open pit slopes

The design and excavation of overlying open pit bench and overall slopes in potentially mud forming soils or weak rocks such as mudstones and shales is critical to the minimisation of mud rush potential. Here again, the best geotechnical practice should be used with the effects of rainfall and groundwater, mining sequence and blast damage being taken into account (Hoek et al 2000). Backfilling of underground stopes

Where backfill is used in underground stopes that may have a hydraulic connection to the cave, it is of paramount importance that decanted water is not permitted to escape from the filled area, that the filled stopes do not become sumps for the accumulation of groundwater, and that stope bulkheads are carefully designed to avoid failure, especially in circumstances in which the fill may become liquefied. Bulkhead failures have been more common than they should be in underground metalliferous mines in recent years. The liquefaction potential of lightly cemented paste fill should not be overlooked.


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Drainage

The provision and maintenance of adequate drainage both inside and outside the mine is one of the most important measures that can be taken to avoid or minimise the impact of mud rushes. Butcher et al (2000) recommend that the following measures be taken in mud rush prone mines: • the hydrological and hydrogeological regimes should be defined and a mine water balance

determined; • surface and underground drainage systems must be designed to prevent rain and

groundwater from entering slopes, open pits, subsidence zones, muck piles and filled stopes;

• surface drains must be kept free of obstruction; • underground water reticulation systems must be designed and maintained to prevent

leakage; • a groundwater monitoring system should be established to detect variations in the phreatic

surface surrounding the mine and measure the quantities of groundwater entering the mine; • records should be kept of the quantities of water pumped from the mine and used in water

balance calculations; • similarly, records should be kept of the quantities pumped from boreholes used for

dewatering; • drainage tunnels, dewatering holes and pumps must be maintained, kept clear and protected

from theft or vandalism; and • an annual audit of mine drainage measures should be conducted by a hydrogeologist. Draw control

The potential influence of uneven draw in enhancing mud rush potential was discussed briefly in Section 10.4.2. Experience suggests that over drawing and isolated draw conditions may be triggers for mud rushes in mud rush prone mines. Butcher et al (2000) recommend that in mines having a history of mud rushes, a maximum drawpoint extraction percentage (eg 120% of the allocated drawpoint reserve) should be set as a shut off limit to prevent mud ingress. 10.5 AIR BLASTS

10.5.1 Terminology

An air blast is the rapid flow of air through an underground opening following compression of the air in a confined space, most frequently by the sudden fall of a large volume of rock. In underground coal mines, such phenomena are sometimes known as wind blasts.


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10.5.2 Causes

The most common causes of air blasts, and the causes of the most damaging air blasts experienced in block and panel caving mines, are major collapses of the first and second types discussed in Section 10.2 above. In the worst case illustrated in Figure 10.2, the sudden collapse of the crown pillar into the underlying void has a piston-like effect, rapidly compressing the air below it as it falls under the influence of gravity. The compressed air, now under high pressure, will escape through whatever opening is available to it. In caving mines, this is most obviously through the ore pile, but it can also be through any other opening intersecting the previously open void. The displaced air will be expelled through the drawpoints and then travel through the extraction level drifts, usually to collect in perimeter drives and make its way out of the mine via accesses and ventilation and hoisting shafts having direct connections to the extraction horizon. Less severe but still violent and damaging air blasts may be caused by falls of ground of the second type (see Figure 10.3), especially those with large basal areas that can compress the air ahead of them as they fall. Heslop (2000) suggests that air blasts of this type are more common in caving mines than might be supposed. Examples of their occurrence have been reported, for example, at Shabanie in 1964 (Laubscher 2000), Urad in 1968 (Kendrick 1970), Northparkes in 1999 (Bell 1999) and Salvador in 1999 (de Nicola Escobar and Fishwick Tapia 2000). In the latter case, illustrated in Figure 10.10, arching had caused cave propagation to be arrested. Cave induction procedures using drilling and blasting had been implemented. On 5 December 1999, a plan area of approximately 4000 m2 collapsed in the north-west zone of the mine, daylighting in the subsidence crater as shown in Figure 10.10. It was estimated that the air velocity generated in the main access was in excess of 500 km/hr. Air blasts may arise from a number of other causes all involving the sudden compression of air in a confined space. They may be a damaging result of mud rushes or wet muck flows (Butcher et al 2000), especially when the liquid mud or muck is present in such quantities that it can completely fill the drifts into which it flows. McPherson and Pearson (1997) describe an air blast problem arising from a completely different cause at PT Freeport Indonesia’s Grasberg mine. Ore mined in the open pit passes down through a number of 2-4 m diameter and 500 m long ore passes into a series of 10 m diameter and 100 m deep ore bins below. At the time concerned, this system of ore passes and bins handled about 10,000 tonnes of ore per hour. The falling rock in any given ore pass generated such high air pressures in the corresponding ore bin that when the ore feed ceased, an air blast travelled up the ore pass to the open pit, blowing rock fragments and copious amounts of grit and dust back into the conveyor feed level. McPherson and Pearson (1997) carried out a theoretical analysis which showed that the air pressure generated was the result of aerodynamic drag on the falling particles of rock. It is clear that similar air blasts (but probably of lesser


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magnitude because the fall heights are likely to be less) could be generated in ore handling systems and by falls of rock into open caved voids in block and panel caving mines.

Figure 10.10: Sequence of collapse events and air blast, Salvador mine, Chile (de Nicola Escobar and Fishwick Tapia 2000)

10.5.3 Effects

Air blasts, particularly those arising from major collapses of the first type, can be extremely violent and damaging events. The damage caused by air being displaced at high pressures and velocities through underground openings of relatively small cross-sectional areas (20-25 m2 in typical cases) can be dramatic. Large pieces of equipment, including vehicles, can be overturned or picked up by the blast and destroyed. Support elements and services can be ripped from the walls of excavations, safety doors can be blasted away, shaft installations can be stripped and, of course, people in the path of the blast can be seriously injured or even killed


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instantaneously. Rock particles entrained in the air stream travelling at speeds of up to several hundred kilometres per hour can cause severe impact damage. A less horrific, but nevertheless important, consequence of air blasts of this type, is that they generate large amounts of dust, particularly through the drawpoints, producing “white out” conditions in which visibility on the extraction level is largely or totally lost. The dust and particles of larger grain size can cause eye injuries to personnel in the vicinity as reported at Salvador (de Nicola Escobar and Fishwick Tapia 2000). The air blasts transmitted to the extraction level through the drawpoints can have sufficient force to knock down and injure people (Kendrick 1970). Where rapid caving to surface is involved, high dust plumes can be produced on the surface by some of the compressed air escaping up through the falling or fallen ground. The questions then arise as to how air blast velocities can be estimated, what categories of damage are likely to be caused by air blasts of particular velocities, and what specific response triggers should be used to control air blast risk? Although published experience of air blasts in block and panel caving mines is limited, much information of value is available from studies of wind blasts in Australian underground coal mines (eg Fowler et al 1996, Fowler and Sharma 2001, Hayes 2000, Mills et al 2000). These studies have involved field monitoring, physical and analytical modelling, the development of wind blast management plans and alleviating the effects of the wind blast hazard by hydraulic fracturing. The approach adopted at the Ridgeway Gold Mine, New South Wales, Australia, in developing an Inrush Major Hazard Management Plan (Logan 2001, Dunstan and Power 2003) provides an excellent illustration of best practice in this area. Four air blast velocity or wind gust categories were defined based on the Saffir-Simpson Hurricane scale which measures wind gust velocities at ground level as is appropriate in the underground mining application. These categories and some of their effects are listed in Table 10.2. A simple model was developed to predict the air blast velocities in drives that would be potentially associated with a major cave back failure. The model is based on a “leaky piston” which generates instantaneous overpressures in the muck pile. Air velocities are linked to overpressures using Boyle’s Law. More complex models may be developed for this purpose. The fragmentation of the muck pile and variability of input data were taken into account in developing the nomogram shown in Figure 10.11. The nomogram provides a relationship between potential air exit velocity in drives (for two, three or four exit pathways), expansion volume or air gap height, and height of muck pile cover above the nearest air exit point. The nomogram shown in Figure 10.11 applies for a particular set of parameters and medium fragmentation in the sublevel caving operations at the Ridgeway Mine and is presented here for purposes of illustration only.


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Table 10.2: Example of the classification and potential impacts of wind gusts or air blasts (after Logan 2001, Dunstan and Power 2003)

Air or wind gust velocity Classification (refs 1-4) Potential impact (refs 1-4)

<15 m/s –

(55km/hr)

Class green

Moderate gale Beaufort Scale No. 7

Difficult walking into wind

15 – 35 m/s

(55 – 125 km/hr)

Class yellow

Cyclone/hurricane category 1

Gale (F0) tornado

Some damage to windows,

signs, vent bags, air and water

pipes.

Unprepared person may be

knocked over.

35 – 45 m/s

(125 – 170 km/hr)

Class orange


Moderate (F1) tornado

Significant damage to signs, vent

bags and pipes. Some cars

overturned. Small projectiles

(sand).

45 – 60 m/s

(170 – 225 km/hr)

Class red


Significant (F2) tornado

Major damage to installations.

Light objects become projectiles

>60 m/s

(225 km/hr)

Class red

Cyclone categories 4 and 5

Severe to inconceivable

(F3-F6) tornado

Severe damage.

Heavy cars lifted off ground and

thrown.

Metal buildings collapsed or

severely damaged.

References: 1. Beaufort Scale

2. Saffin-Simpson Hurricane Scale

3. Tropical Cyclone Intensity Scale

4. Fujita Tornado Scale


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Moonie

Air velocity in drive (s), m/s (multiply by 3.6 for km/h) Expansion volume height, expressed in height (m)

Figure 10.11: Upper quartile potential air velocities in drives resulting from large scale back failure for medium fragmentation – an illustrative example from Ridgeway Gold

Mine, Australia (after Logan 2001, Dunstan and Power 2003) 10.5.4 Prevention and Amelioration

Damaging and life threatening air blasts are consequences of other types of event, namely major collapses and mud rushes. Clearly, the most effective means of preventing air blasts and of ameliorating their effects is to prevent or minimise the frequencies and magnitudes of their causes. As has been noted, the most severe air blasts experienced in block and panel caving mines arise from large falls of ground into open voids or air gaps above the muck pile as illustrated in Figures 10.2 and 10.3. The most obvious and effective means of preventing these air blasts is to ensure that excessive air gaps are not permitted to develop. This can be done quite simply by adopting strict draw control strategies of the type discussed in Chapter 7. If cave propagation is slowed significantly, or is arrested and a stable cave back develops, there may be a temptation to continue to draw caved ore in order to maintain production and cash flow. Experience makes it clear that this temptation should be resisted at all cost. If the development of a large void or air gap above the caved ore pile is unavoidable, the impact of any air blast produced by a major collapse can be reduced by ensuring that the height of the air gap is restricted and that an adequate cover of caved ore is maintained above the drawpoints. Laubscher (1994, 2000) has developed guidelines on these issues based on decades of accumulated experience. He points out that, where a stress caving mechanism operates, it is not always possible to avoid the development of some air gap above the caved ore. He suggests that no significant air blast should result from falls from a sub-horizontal cave back if the height of the air gap is restricted to no more than 10 m. When there is an inclined cave back, Laubscher (2000) suggests that there should be no air gap (other than an expansion gap) to avoid migration of ore or dilution down the muck pile slope. These guidelines imply the


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establishment of a monitoring system to monitor the rate of caving and the development of any air gap. Obtaining the necessary measurements may not be a straightforward matter, especially in the early stages of caving. Laubscher (2000) points out that finely fragmented and well graded ore is more likely to provide effective resistance to the passage of an air blast than more coarsely fragmented and poorly graded ore. He suggests that a cover of 60 m of well graded ore or 90 m of poorly graded coarse material might be required to minimise the effects of air blasts produced by compression of the air in a void above the caved ore pile. Of course, the efficacy of such measures will depend on the volumes of the void and of the rock fall producing the air blast. Laubscher’s parallel guideline on the height of the air gap should not be ignored. Indeed, it would appear logical to link the permissible absolute height of the air gap to the height of the column of caved ore. Except for low ore column heights (of less than, say, 30-40 m) for which the application of such a rule-of-thumb is likely to be unworkable, restricting the air gap to 10 - 20% of the height of the caved ore column should generally provide protection against the air blast hazard. It has been noted previously that the drawpoints are not the only openings that may be connected to the cave and through which the compressed air may escape. Other such openings include exploration drives, undercut level development and open boreholes. It is essential that potential entries into the void be safeguarded in advance of their being intersected by the cave. Methods that have been used include solid concrete plugs, engineered bulkheads and backfilling drives with broken material. 10.6 WATER AND SLURRY INRUSHES

Water and slurry inrushes are those occurrences in which water and/or slurry enter the mining zone from some external source such as a water storage dam, a tailings dam or a backfilled stope (Heslop 2000). Their likely causes, effects and appropriate preventive and ameliorative measures are similar to those discussed in Section 10.4 for some aspects of mud rushes and so will not be repeated here. Strictly speaking, water and slurry inrushes can represent a different type of operational hazard from mud rushes although they can be contributing factors to mud rushes. Accordingly, they have been listed separately for completeness.

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CHAPTER 11

RISK ASSESSMENT FOR BLOCK CAVING 11.1 INTRODUCTION TO RISK ASSESSMENT

n Chapter 1, the concept of risk in cave mining was introduced, the risk factors associated with a block or panel cave mining project or operation outlined, and a number of terms associated with risk assessment were defined. The purpose of this chapter is to expand on

that introduction and to describe an approach to risk analysis for cave mining known as CaveRisk, developed by Summers (2000a) as part of the International Caving Study Stage I. For completeness and ease of reference, some of the introductory material presented in Section 1.4.2 will be repeated here. Techniques known as risk analysis, risk assessment and risk management are now applied to a range of engineering and other undertakings. The terminology used may vary with the discipline or sphere of application. The concern here is with the assessment and management of the engineering and technical risks associated with the adoption and operation of a particular mining method. For this purpose, a useful generalized definition of risk assessment is that given by the UK Engineering Council:

Risk assessment is a structured process which identifies both the likelihood and extent of adverse consequences arising from a given activity.

Engineering decisions of the type being considered here are subject to a number of uncertainties, the manifestation of which often results in the failure of a project to meet its objectives in full or in part. These uncertainties can be considered to be of two general types:

• what we know we don’t know or parameter uncertainty; and

• what we don’t know we don’t know or conceptual uncertainty.

I

Chapter 11: Risk Assessment for Block Caving

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Parameter uncertainty is the easier of these two types of uncertainty to account for in engineering procedures. The long-established concept of a factor of safety is a commonly used method of attempting to account for parameter uncertainty. The strength testing of any natural material such as rock will yield a range of results but a single value is usually used in deterministic calculations. The selected value may be an average, a percentile value, or some other estimate. The selected value is often termed the “representative” value, yet we know that it may not be representative in all cases. There is always a risk that the actual strength mobilised in a specific location could be at the low end of the range of the test results, and that the load applied is at the upper end of the expected range. A factor of safety is often used to compensate for this uncertainty. The factor of safety can be viewed as a crude risk management tool. Probabilistic methods are also used as an alternative approach to addressing the same issue in geotechnical engineering (eg Christian et al 1994, Pine 1992). Modern risk analysis techniques examine risk in a more explicit manner than does applying a series of safety factors (or partial safety factors) specified in a code of practice. Conceptual uncertainty or uncertainty about how particular sets of conditions will develop and what their eventual outcomes will be, is usually of greater concern and more difficult to assess. A risk assessment and management approach seeks to understand the sources of risk associated with a given project or design, to evaluate their consequences and to put in place procedures to manage those risks. Risk assessment approaches have been used in the international minerals industry since at least the late 1980s. Most major companies have developed procedures for assessing and managing the risks associated with a range of aspects of their operations. In Australia, for example, there is a national standard on risk management (Standards Australia 1999), specialist procedures, tools and services are available (eg Joy 1994), and State-based mining regulations often specifically prescribe risk assessment methods to meet duty of care responsibilities (eg the New South Wales Mines Inspection General Rule 2000). National Minerals Industry Risk assessment guidelines are being developed through the Minerals Council of Australia (Joy 2002). Risk assessment and management approaches are used to develop core or major hazard management plans such as that referred to in Section 10.5.3 concerning air blasts (Booth and Hume 2002, Joy 2001, Logan 2001). 11.2 DEFINITIONS

Summers (2000a,b) has compiled a set of definitions of terms suitable for use in the consideration of risk in the present context. This is not intended to be a comprehensive glossary of terms used in risk assessment but is based on definitions published elsewhere, modified to suit the mining engineering environment. Some of these definitions were introduced in Chapter 1. It should be noted that they differ from those given by other authorities (eg Standards Australia 1999, Joy 1994).


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Hazard - a potential occurrence or condition that could lead to injury, damage to the environment, delay, or economic loss. Risk analysis - a structured process that identifies both the likelihood and the consequences of the hazards arising from a given activity or facility. Risk evaluation – the appraisal of the significance of a given quantitative (or, when acceptable, qualitative) measure of risk. Risk tolerability - a range between the maximum tolerability, above which the project must be abandoned, and the minimum tolerability, below which a risk is so small that its occurrence will not affect the project. Risk assessment - comparison of the results of a risk analysis with risk acceptance criteria or other decision parameters. Risk reduction measure - an action that could be adopted to control a risk by either reducing the likelihood of occurrence, or by mitigating the consequences of an occurrence. Risk management - the process by which decisions are made to accept known risks, or the implementation of actions to reduce unacceptable risks to acceptable levels. 11.3 PROJECT DEVELOPMENT

The UK Association for Project Management suggests that, in general, implementing project risk analysis and management is best done in the early stages of a project’s life cycle when it is likely to be most effective and useful in guiding the development of the project. Introducing risk analysis late in the life cycle is difficult and with few compensating benefits, because contracts have been placed, equipment purchased, commitments made, reputations are on the line, and managing change is difficult and unrewarding. The stages in the life cycle of a mining operation are shown in Figure 11.1. The route by which a mining project develops, from the earliest scoping study to the construction stage, reflects an ever increasing level of complexity with more and more detail being added. Most mining projects go through three or four stages of analysis prior to a decision being made to finalise the investment strategy and commence construction. These stages may be referred to as a scoping or pre-feasibility study, a feasibility study, an investment or bankable study and a detailed design study. The improvement in detail and understanding implicit in this progress is illustrated in Figure 11.2.


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Concept

Design

Construction

Operation

Modification

Disposal &Closure

Improvement

Figure 11.1: Life cycle stages of a mining project

Time

Scoping study

Feasibility study

Investment study

Detailed design

Figure 11.2: Stages in the development of a mining project

To be of the greatest benefit, risk analysis should be undertaken at the earliest stage of the study and updated regularly as the project reaches the various milestones.


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11.4 RISK ANALYSIS TOOLS AND CONCEPTS

11.4.1 Risk Analysis Tools

Qualitative and quantitative engineering risk analysis employs a variety of tools, some of which have been developed specifically for the purpose with others being adapted from other disciplines. The risk analysis literature contains numerous examples of the tools that can be applied, especially to safety risk investigations (eg Joy 1994, Summers 2000b). Some computer-based systems are also available. Many of the concepts underlying these tools are applicable to engineering risk assessment, but some modification is required to ensure that the analysis is not driven by the capabilities of the techniques available.

Some of the tools that are considered relevant to gaining an overall understanding of the components and structure of a block cave study are:

Concept Map - a map of the main stages, phases or controls of a block cave project

Influence Diagram - an unstructured representation of the links and dependencies between properties, processes and design controls in the project

Logic Tree - a hierarchical presentation of the properties, processes and design controls in the project

Of these, the logic tree and the concept map are the two tools that will be used most extensively in this chapter. Generally, influence diagrams are used to collect thoughts and ideas (eg through brain-storming) which are then transformed into the more structured representation of a concept map and a logic tree. Such an approach was used to establish influence diagrams in the development of the CaveRisk system to be described in Section 11.5.

11.4.2 Sources of Risk

All forms of human activity, or inactivity, contain elements of risk. It is the role of a risk analysis to evaluate risks from all sources. In the mining environment, the sources of risk include some or all of the following: • optimistic or unsupported assumptions, • limited or poor quality data, • unjustified extrapolation of past experience, • use of inappropriate computer models, • models and parameters that have not been properly validated,


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• unsuitable or misdirected “expert” advice, • unwillingness to acknowledge previous failures, • unexpected changes in conditions, • natural variability, • aggregation of risks, and • external hazards.

11.4.3 Uncertainty

The definition and discussion of risk given in the previous sections, referred to uncertainty as the primary source of all risk, uncertainty in the actual value of a parameter or property of interest, and uncertainty as to how a particular set of conditions will develop and produce an outcome. These were referred to as parameter uncertainty and conceptual uncertainty, respectively. As an example from block and panel caving, consider the approach to caveability assessment discussed in Chapter 3 in which the objective is to predict the minimum hydraulic radius or shape factor of the undercut required to induce caving of a given rock mass. In this approach, if a rock mass parameter such as Q', RMR or MRMR is determined and plotted in a particular way on a caveability chart, then the minimum hydraulic radius required may be predicted. For this method to be reliable, the conceptual relationship between MRMR or other parameter and hydraulic radius must be robust. Because the relationship is purely empirical (ie it does not follow some law of physics but is based on human observations and assumptions), there is a conceptual uncertainty (the relationship might not be valid for all circumstances of geology or geometry as was suggested in Chapter 3) and this constitutes a risk. Furthermore, the characteristics of the rock mass which are measured are often indices (eg RQD), or cover a range of natural variability (eg UCS), yet a unitary value of the rock mass parameter is often derived and used. Clearly, a single value is highly unlikely to be universally applicable across the whole site, because there will be some regions in which the rock properties fall in the tails of the distributions that are used to determine the average or “most likely” values. This can be described as parameter uncertainty; we can never be entirely certain that the "representative" value chosen will be correct for each location. As has been noted in this context in Chapter 3, engineers sometimes introduce or impute false precision to an uncertain parameter. Reporting the average uniaxial compressive strength of a rock sample to a number of decimal places (or even to one), does not alter the range of values that might exist in the sample results. In the approach to open stope stability and caveability assessment discussed in Section 3.3, this and other aspects of uncertainty were allowed for by incorporating probabilities into the charts following Mawdesley et al (2001).


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Two types of uncertainty are involved in making subjective assessments, that which is inherent in the variable itself, and the judgement and level of knowledge of the person making the assessment. There are several reasons why experts may make inaccurate estimates of the uncertainty distributions intended to describe unknown variables. These reasons may be associated with bias or other aspects of human nature as well as with flaws in the processes used to make the assessments. 11.5 CAVERISK

11.5.1 Purpose

An approach to risk analysis for block and panel cave design known as CaveRisk was developed as part of the International Caving Study Stage I (Summers 2000a). CaveRisk provides a structure to guide the designer through the evaluation of a prospective caving operation, allowing the project team to determine the risk in each of the defined areas. CaveRisk does not tell the designer what he or she should or should not do. Rather, the system guides the designer through the steps required to identify where the areas of risk exist and the approximate severity of each risk. The risks associated with five major design areas are determined in a structured manner to be described below. The methodology encourages designers to adopt risk assessment as an integral part of the investigation and design of their block or panel caving project. By using CaveRisk and ranking the risks across the project, the Project Manager should be able to direct scarce resources to address the most important issues and should reduce the chance that excessive effort is expended on non-critical items. Project Managers should not attempt to compare risk scores from other projects to try to determine if one is more risky than another. Similarly, there is no absolute threshold score below which a project can be considered to have acceptable risk. 11.5.2 Outline of CaveRisk

CaveRisk provides a structure to guide the Project Manager through the process of investigation and evaluation of a prospective caving operation, and will allow the project team to arrive at judgements about the risk in each of the defined areas. The judgements and the assessments will be those of the project team, and the Project Manager will determine the levels of risk that will be accepted. At no point will CaveRisk tell the Project Manager that a particular course of action should or should not be followed; rather, the system will provide a guide through the steps required to identify where risks exist and the approximate severity (or magnitude) of those risks.


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The Project Manager should not work alone, nor should the project team operate in isolation. Inevitably, there will be a need to use experts in fields relevant to the investigation and design. For example, the collection and interpretation of geotechnical data should be managed and supervised by a competent geotechnical engineer who is experienced in underground mine evaluation and design. By understanding the issues, the Project Manager and the project design team will be able to take advice from the most appropriate specialists and consultants. Risk identification

Risk identification is perhaps the most important part of the risk assessment process and relies heavily on “expert judgement”. In the preceding discussion, conceptual uncertainty was shown to be a major source of risk in any mine design. In arriving at a successful block or panel cave design, the Project Manager will need to answer five specific questions:

• Have we identified all the issues relevant to this site? • How well does the design team understand each of these issues? • Does the design adequately address each of the issues? • How serious will it be if we have failed in one or more of these issues? • How easily can we manage inadequacies in the design?

These questions must also be addressed in the formal risk assessment. In fact, the risk assessment method presented here sets out to answer all five questions, covering each issue of relevance, and from the answers to assess the risk to the project.

Issues of relevance

The issues of relevance in addressing risk in block and panel caving projects were identified from the pooled experience of a group of those associated with the International Caving Study Stage I and from a comprehensive search of the literature including the Block Caving Manual (Laubscher 2000). Each issue has been positioned within a logic tree as discussed in Section 11.5.3 below. Risk to success

Threats to the success of the project can be summarised in two of the questions posed above: (i) how well does the design team understand each issue, and (ii) does the design adequately address each of the issues? Answering these questions involves considerations of: • the quality and reliability of the basic investigation; • the validity and reproducibility of the modelling;


408

• the value and relevance of external advice; and • the appropriateness of the design.

In this context, modelling is taken to include both conceptual models (such as the distribution of RMR zones and resource modelling) and any numerical or deterministic models used to predict the expected behaviour using a set of input parameters. Parameter uncertainty was discussed previously as a major source of risk in mining projects.

For each of the headings, a graduated assessment scheme has been devised ranging from a poor understanding up to a comprehensive investigation and solution. These assessments are captured in the form of a likelihood that there are inadequacies in the overall understanding of each issue. Seriousness of inadequacies

The seriousness of a failure to understand and address each issue is judged in terms of a consequence for the overall project. A number of economic and non-economic consequences are allowed for. Ease of management

The final question that the Project Manager needs to ask is the degree to which a failure to understand the issue can be managed. To some extent this is linked to the consequences of a failure, but different issues can be managed to differing degrees. Perhaps the more difficult issues to manage are those relating to the caveability of the rock mass. Figure 11.4 shows clearly that the issues governing the success of caving are dominated by naturally occurring properties and factors. In general terms, these natural properties and factors are fairly difficult to manage, although cave inducement techniques can exert some modifying influence.

11.5.3 Topics and Focus Issues

The identification of the issues to be examined during the evaluation process was carried out by members of the International Caving Study team, and appear to users of CaveRisk as a series of logic trees. The issues have been grouped into five major topics which have been discussed in detail in earlier chapters of this book:


409

• Caveability

• Fragmentation

• Caving Performance - including the caving process and cave related design

• Excavation Stability - including mine layout and excavation design

• Major Hazards - items that require special attention during design

Each of the five topics has been further divided into a number of focus issues. These focus issues include items that may be classed as processes or properties, and which influence the measurable performance of the individual topics as indicated by the following list:

Topic Focus Issues

Caveability Ore strength, waste strength, undercut geometry and cave inducement.

Fragmentation Block geometry, detachment mechanism, impact breakage, autogenous comminution and draw strategy.

Caving Performance Caving rate, cave shape, undercut strategy, fragmentation, mine layout and economics.

Excavation Stability Ore strength, waste strength, layout geometry and natural events.

Major Hazards Inrushes, uncontrolled collapses, rock bursts and air blasts.

The five major topics and their related focus issues form the “concept map” referred to in Section 11.4.1. It will be noted that, not surprisingly, some of the focus issues occur under more than one topic. For example, it is apparent that the strength properties of the waste rock should influence both the caveability of the waste and the stability of the mining excavations. Based on the definitions of the topics and the focus issues, it now becomes necessary to assemble a set of "logic trees" to indicate the connections between the focus issues and to identify other, less important sub-issues, that are also relevant to the design of a caving mine. These logic trees are presented in Figures 11.4 to 11.8, with Figure 11.3 providing a legend. In these figures, each focus issue is defined as either:

• a property, • a process, or • a design control,


410

and each is classed as being either “main” or “secondary” (see Figure 11.3).

SecondaryControl

MainControl

SecondaryProcess

MainProcess

SecondaryProperty

MainProperty

Duplicated

Details elsewhere

Figure 11.3: Symbols used in Figures 11.4 to 11.8

CaveabilityCaveability

HydraulicRadius

CaveInducement

HydraulicFracturing

BoundaryWeakening

BulkBlasting

CraterBlasting

OreMRMR

WasteMRMR

UndercutGeometry

OreRMR / Q

OreM Rating

Lab Strength

G/Water

Fabric

Structures

Other Caves

Total Stress

GradeDistribution

WasteM Rating

Other Caves

Total Stress

Regional σ

Induced σ

Regional σ

Induced σ

WasteRMR / Q

Lab Strength

G/Water

Fabric

Structures

Mathew’sStability

JKMRCCaveability

Figure 11.4: Caveability


411

FragmentationFragmentationOre

FragmentationWaste

Fragmentation

PrimaryFragmentation

SecondaryFragmentation

IncipientBlock

Detachment

Time Lag Structure

Fabric

σ Fracturing

σ Level

RockStrength

ImpactBreakage

AutogenousComminution

DrawStrategy

Air GapBlockStrength

Cave Rate

Swell Factor

Draw RateLab Strength

Fabric

Defects

DifferentialDraw

SizeDist’n

Cave Rate

Draw Rate

ColumnGeometry

BlockStrength

AspectRatio

Draw Rate

Cushioning

FinesMigration

Defects

Fabric

Figure 11.5: Fragmentation

SecondaryBreakage

Caving PerformanceCaving Performance

Economics MineLayout

UndercutStrategyFragmentation Cave

RateCaveShape

Cut-off GradeOptimisation

U/C LevelSelection

MineCapacity

ExtractionSequence

Hang-upTreatment

Direction

Geometry

Type

U/CRate

Caving σ

Seismicity

Lock-up

Regionalσ

Inducedσ

Natural

Induced

Draw Rate

Lock-up

OreRMR/Q

Caving σ

UndercutLevel

ExtractionLevel

OreFragment’n

OreRMR/Q

SurfaceEffects Subsidence

InfluenceAngle

Figure 11.6: Caving performance


412

Excavation StabilityExcavation Stability

InfrastructureWaste

Stability

Crusher

Ore Passes

Ventilation

Pumping

NumberLocation

NumberLocation

Size

Air loss

Operability

MRMR

Undercut LevelOre

Stability

Fragmentation

Seismicity

Geometry

Ht of U/CU/C Profile

Blastdamage

Remnants

Rock type

MinorApex

MajorApex

Extraction LevelOre

Stability

Water

Operability

Cave Shape

Tonnage call

Total σ

MRMR

G/W

Rain

Longholes

Inrushes

Productivity

Fragmentation

Muckingmethod

Safety

Abrasiveness

Seismicity

Extractionratio

Figure 11.7 Excavation stability

Major HazardsMajor Hazards

UncontrolledCollapseRockburst

Air Blast

Inrushes

Mud InrushWater Inrush

Figure 11.8: Major hazards


413

A property, as the term implies, is a physically measurable parameter which describes the physical characteristics of the rock mass, the caved material, or the geological environment. A process is either a direct action (such as blasting) that is applied to the rock mass during mining, or a naturally occurring event associated with mining, such as impact breakage. Design controls are constraints that are built into the system of extraction (such as undercut direction) but are not directly related to the behaviour of the rock mass. For the purposes of the risk model, it is assumed that each topic, focus issue and sub-issue is independent of any other topic, focus issue or sub-issue. This assumption is considered to be valid for present purposes because of the manner in which issues have been established in the hierarchy, although it may not always be valid in a more general sense. Where the same issue occurs in more than one location in the logic tree, it is defined only once, and then referred to in all other instances. 11.5.4 Likelihood and Consequences

Risk is defined quantitatively as the product of the probability of an event occurring and the consequences of that event when it does occur. In the design of a block cave operation, the risk of concern is that associated with inadequacies in our understanding of the important aspects of the design. To establish the risk, the likelihood and consequences of an inadequacy must be assessed. Classification schemes have been devised for making these assessments for each of the focus issues. The classification schemes use four likelihood and consequence categories with a fifth category being added for “don’t know”. Once the user has selected the most appropriate likelihood and consequence classification, using one or more of the descriptions or the semi-quantified measures, a score is assigned from which the risk can be determined. The assessments will change as the project evolves, as more knowledge becomes available and as confidence is gained in understanding each issue. In essence, the aim of further data gathering exercises and design studies is to reduce the levels of risk to acceptable values. Each issue is examined and the designer specifies the likelihood of inadequacies existing in understanding based on: • reading of the knowledge base; • input from members of the project team; • advice from consultants; and • guidance from other specialists.

Likelihood

Likelihood of occurrence is assessed using the categories shown in Table 11.1, according to the four main criteria, investigation, modelling, external advice, and design, or the general descriptive criterion if none of the specific criteria are applicable. The approximate probability included in Table 11.1 is based on a log scale and can be used as an additional quantification of


414

the likelihood of occurrence and hence of the risk. The probability relates to the likelihood of a single event occurring and will often not be applicable to multiple events which should be expressed as a frequency of occurrence. If the user makes different assessments against different criteria, the most severe assessment should be used to determine the risk.

Table 11.1: Likelihood classes Likelihood of Inadequacies and Score

Very Unlikely Unlikely Probable Highly Likely

Unknown

SCORE ASSIGNED

1 2 5 10 10

General Descriptive

Issue well understood.

Issue mostly understood but some gaps.

Some fundamentals of the issue are understood.

Poor understanding of the issue.

-

Investigation

Full data set plus interpretation report.

Some data and some interpretation complete.

Some data collected but not interpreted.

Little or no relevant data.

-

Modelling

Set of reproducible results using a proven modelling tool.

Good modelling results but some questions remain.

Limited or ambiguous modelling results. Results from a poorly validated modelling tool.

Little or no modelling.

-

External Advice

Successfully reviewed by renowned expert.

Some expert advice but not fully documented.

Internal reviews only, or some expert consultation.

Little formal external expert advice.

-

Design Robust design takes account of all issues.

Partial design with outstanding issues or inadequacies.

Only an outline design that does not address all issues.

No design exists.

Approximate Probability

< 0.1% 0.1% to 1% 1% to 10% > 10% -

Consequences

Consequences are assessed as the non-economic or the economic impact of the investigation, the modelling, the expert advice or the design being less than optimal. The assessments are made using Tables 11.2 and 11.3 for non-economic and economic consequences, respectively. Non-economic consequences include issues such as personnel safety, environmental impact and the reputation of the owner. These non-economic consequences may assume greater importance in the early stages of a project, especially when the overall economics of the project are not clear.


415

Economic consequences are assessed in a semi-quantitative manner in terms of capital, operating cost, critical path delays, and delays in production. In all cases, the assessments are normalised against either the asset value of the project, the annualised operating cost, or the planned monthly production. To improve any quantification of delay, the economic value of one month’s delay is required. The value of delay will probably vary between the delay to the commencement of initial production and the delay to routine production once mining has begun. Delays in bringing the project into operation are described as critical path extensions and will include financial costs such as loan re-payments and delays in generating revenue. As with the likelihood of occurrence, the most serious consequence is selected for the determination of the risk.

Table 11.2: Non-economic consequence classes

Non-Economic Consequences and Score

Very Low Low Moderate High Unknown

SCORE ASSIGNED 1 2 5 10 10

Environmental Impact Localised

degradation Widespread

degradation Severe

degradation Likely

prosecution -

Community Impact Negligible Slight Moderate Severe -

Personnel Safety No injuries Minor

injuries Serious

injuries Fatalities -

Lost Time (shifts) 0 0 to 500 500 to 5000 > 5000 -

Owner’s Reputation Negligible Slight Moderate Severe -

Table 11.3: Economic consequence classes

Economic Consequences and Score

Very Low Low Moderate High Unknown

SCORE ASSIGNED 1 2 5 10 10

Capital Impact

(AV = Asset Value)

< 0.1% AV 0.1% AV

to % AV

1% AV

to 10% AV

> 10% AV -

Annualised Operating

Cost (OC)

< 2% OC 2% OC

to 5% OC

5% OC

to 10% OC

> 10% OC -

Project Interruption

(critical path)

< 1 month 1 month

to 3 months

3 months

to 6 months

> 6 months -

Delayed Monthly

Production (MP)

< 25% MP 25% MP

to 50% MP

50% MP

to 100% MP

> 100% MP -

NB: It will be necessary to apply a value to a 1 month critical path delay


416

11.5.5 Risk Determination and Risk Acceptance

As indicated above, risk is determined for each item from the relationship:

Risk = Likelihood × Consequences

The likelihood and consequence assessments are made by the user following a thorough investigation of the potential hazards and advice from suitable advisors. Based on these assessments, the risk for each item is determined from the matrix shown in Table 11.4. As discussed above, the most serious consequence is used to determine the level of risk. The severity of the risk is a seven-fold measure grading from Level 1 (lowest) to Level 7 (highest). The colouring of the cells indicates the approximate significance of the risk levels. In general, the risk significance is taken to be as follows:

• Levels 1 to 3 are low risk (green) and probably acceptable; • Levels 4 and 5 are moderate risk (amber) and not acceptable

other than in the early (eg pre-feasibility) stages of a project; and

• Levels 6 and 7 are high risk (red) and unacceptable at any stage in a project.

This approach is a simplification because the level of acceptable risk will actually depend upon the nature of the consequence.

Table 11.4: Risk determination matrix

Most serious consequence Very Low Low Moderate High Very Unlikely Level 1 Level 2 Level 3 Level 4 Unlikely Level 2 Level 3 Level 4 Level 5 Probable Level 3 Level 4 Level 5 Level 6 Highly Likely Level 4 Level 5 Level 6 Level 7

Risk acceptance is a complex matter and is influenced by several factors many of which are difficult to quantify. The level of risk acceptance depends on the type of investment under consideration (eg expansion or green field site) and the nature of the risk itself. Furthermore, a mining company may be quite averse to risk in the safety, health and environmental areas, for example, but be prepared to accept financial risks more readily. For these reasons, it was not considered possible or advisable, to include any risk acceptance criteria in CaveRisk.


417

In general, Project Managers would be well advised to adopt the so-called ALARP principle of risk acceptance (As Low As Reasonably Practicable). The ALARP principle suggests that risk reduction measures will be adopted where they are not disproportionately costly compared to the size of the risk. Thus, the ALARP principle suggests that a risk reduction measure costing $10 million should not be adopted for a risk that has a 1% likelihood of occurrence and a most likely loss of $200 million. In other words, spending $10 million to address a risk worth $2 million (1% x $200 million) would be considered disproportionately costly. However, where safety is the concern, such a calculation may be too simplistic, and other risk aversion considerations may have to be taken into account.

11.5.6 Risk Manageability

In the determination of the risk, it is necessary to consider the level of manageability of each issue compared to all others. Manageability should not be confused with the seriousness of the consequences associated with each issue. Rather, it is a measure of the ability to control the risk if it develops. Part of risk manageability is also the ability to identify when a risk is about to eventuate, using one or more “risk indicators”. For example, the inability to initiate a cave as intended is likely to be difficult to manage in the early stages of undercutting, whereas it may be possible to influence cave development in the later stages of the operation. To account for these variations in manageability, a “manageability multiplier” has been devised which will effectively apply a weighting to the raw risk level. Thus, the “risk manageability” for each issue will be determined as:

Risk Manageability = Risk Score × Multiplier

The level of manageability for each Main Item is proposed within CaveRisk but the designer may over-ride this if he or she has sufficient evidence or reason to support such an approach. Clearly, it is not in the interest of the project to reduce the manageability multiplier simply in order to reduce the risk score. The unmanageability classes and multipliers are defined in Table 11.5.

Table 11.5: Risk unmanageability classes

Unmanageability Multiplier Very Low Low Moderate High MULTIPLIER 1 10 20 50 Manageability Easily

managed Manageable Manageable

with difficulty Unmanageable


418

The levels of unmanageability have been preset in the CaveRisk system for all of the main controls, processes and properties shown in Figures 11.4 to 11.8. The preset levels are set out in Table 11.6. The system will apply the risk manageability multiplier at the level of the main controls, processes and properties.

Table 11.6: Unmanageability ratings for focus items

Controls Processes Properties

Hydraulic Radius

H 50 Cave Inducement M 20 Waste RMR M 20

Draw Strategy H 50 Waste Fragmentation M 20 Ore RMR H 50 Economics L 10 Ore Fragmentation H 50 Cave Shape M 20 Mine Layout M 20 Primary Fragmentation M 20 Undercut Strategy

M 20 Secondary Fragmentation

L 10 Ore Stability H 50

Autogenous Comminution

M 20 Waste Stability

M 20

Impact Breakage M 20 Block Strength H 50 Air Gap L 10 Fragmentation M 20 Cave Rate L 10

Based on the risk score and the manageability assessment, the Risk Manageability number is determined, within the following ranges:

Risk Manageability Calculation Result

Risk Manageability Level

Risk Manageability Flag

1 to 9 RM Level 1 Green



50 to 99 RM Level 4 Amber



500 to 999 RM Level 7 Red




419

The coloured flag is set within CaveRisk to signify the level of risk manageability for Main Items.

11.5.7 Risk Presentation

The presentation of information within CaveRisk is based on the hierarchical structures shown in Figures 11.4 to 11.8. The user drills down through the various focus issues, to ever finer points of detail, and makes assessments to the lowest level. The system then “rolls up” the results, and colours each focus issue or topic according to the most serious risk assessed for all its subordinate issues. This is illustrated in Figure 11.9. The level of risk is indicated by the colour, and the user can interrogate the system by drilling down to determine which issue(s) control that risk.

Figure 11.9: Hierarchy of risks throughout focus issues

11.5.8 Rules Operating in CaveRisk

The hierarchy of presentation of the risks is described in the previous section but other rules that the user should understand are applied during the risk determination process. These rules cover the determination of risk level and the calculation of the risk manageability index.

TOPIC Risk Level = 5

Sub-Issue Risk Level = 2


Sub-Issue Risk Level= 1


Focus Issue Risk Level = 5



Sub-Issue Risk level = 2


Sub-Issue Risk level = 2




420

Risk level

The risk level can be determined at any point within the structure of topics, focus issues and sub-issues from the likelihood and consequence assessments. However, there are a number of rules that are followed governing the hierarchy. For the purposes of this description, the term “node” refers to items, focus issues, and sub-issues without differentiation.

1. If a node has daughter nodes, then its risk will be equal to the worst case risk of all its daughters.

2. If a node has no daughters, it cannot inherit risk from any other node and must be assessed

directly. 3. Where an assessment has not yet been made and no inheritance is appropriate, the risk is

considered to be null. Risk levels are not initiated in the program to high or to low values. 4. If a node has daughter nodes, but the daughters have not yet been assessed (presumably

because of lack of data in the early days of a project) then it can be assessed directly. However, once at least one daughter node is assessed, the node inherits the risk level from its daughter. The program contains warnings to make users aware that they are about to over-write assessments.

5. The most severe assessment, for likelihood and consequences, is used to determine the risk. 6. The risk level is determined between Level 1 and Level 7, using the likelihood and

consequence assessments from Tables 11.1, 11.2 and 11.3, and the matrix in Table 11.4. 7. A node can be defined as “not applicable” and the node, plus all daughter nodes, is

excluded from calculation and all of the reporting functions. Nodes defined as “not applicable” are greyed within the program. The user may reverse the definition at any time.

Risk manageability index

The risk manageability index is determined from the risk level for the node and the pre-set unmanageability ratings given in Table 11.6. The unmanageability ratings have been defined at focus issue level. The following rules are used to govern the hierarchy of the risk manageability index.


421

1. The risk manageability index is assumed to apply to all the daughter nodes of a node where the index is provided.

2. The risk manageability index is not calculated for any node where the risk has not yet been

determined. 3. The calculated risk manageability index is not inherited from the daughter node. In all

cases the index is calculated from the assigned unmanageability rating and the risk determined according to the rules noted above.

4. The risk manageability index is determined from the score equivalent to the likelihood and

consequence levels assessed using Tables 11.1, 11.2 and 11.3; and the unmanageability rating specified in Table 11.6. The risk manageability index is the product of the three numbers.

5. Where the risk at the node is null, the risk manageability index is not calculated. 6. Where the node is defined as “not applicable” the risk manageability index is not

calculated. 11.6 CONCLUSION

This chapter has described the application and implementation of the CaveRisk system. Risk assessed using this method is not an absolute measure but provides an internal ranking of the relative risk levels of most of the issues that require examination in a block or panel caving project. Project Managers should not attempt to compare the risk scores of different projects in an attempt to determine if one is more risky than another. Similarly, there is no threshold risk score below which a project can be considered to have an acceptable risk. The methodology proposed is intended to encourage managers and designers to adopt risk assessment as an integral part of the evaluation and design of their caving project. The CaveRisk system has the following advantages: • it provides a structured approach to analysis and design; • it is evolutionary and updateable from experience; • it provides a ranking of the risks to various aspects of a design; and • it encourages wide-ranging input to each project from experts and specialists.


422

By using CaveRisk and ranking the risks across the project, the Project Manager will be able to direct scarce resources to address the most important issues and will reduce the chances of excessive effort being expended on non-critical issues.

The description of CaveRisk given in this chapter indicates that the system is intended to be used to determine the risk to the mining design and its implementation. In most projects three other main areas will contribute to the final decision as to whether or not the project is viable, processing, economics, and health, safety and environment. To some extent, health and safety issues are implicit in the mining design and are addressed within CaveRisk, but there are other, broader, issues that must be examined prior to any decision being taken to proceed. These issues lie outside the scope of this book.

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465

``

APPENDIX A

GLOSSARY

his glossary does not attempt to define every mining or geomechanics term used in this book. Rather, it lists and defines only those terms that have particular relevance and application to block and panel caving methods of mining.

Advance or advanced undercut – an undercutting strategy in which the undercut mining face is advanced slightly ahead of a partially developed extraction level. Air blast – the rapid flow of air through an underground opening following compression of the air in a confined space. Air gap – the clear vertical distance between the top of a pile of caved ore and the in situ cave back. Angle of break or angle of subsidence – the angle made by the cave boundary (usually by its intersection with the surface) and the horizontal. Angle of draw – the complement of the angle of break, ie the angle made by the cave boundary with the vertical. Arching – the formation of an at least temporarily stable arch or beam of in situ or broken rock by the transmission of lateral inter-block or inter-particle lateral forces. Block Cave Fragmentation (BCF) – an expert system program developed by Dr. G S Esterhuizen for predicting the degree of fragmentation likely to be produced in a block cave. Block caving – a mass mining method in which a block of ore is undercut and, following the drawing of some of the broken ore, caves progressively.

T

Appendix A - Glossary

466

Brow – the overhanging, often unsupported, near vertical face at the intersection of vertical or sub-vertical and horizontal excavated surfaces as at the interface between a drawpoint and a drawbell. Bulking or swelling – the increase in volume of a mass of rock when it has caved or been otherwise broken and removed from its in situ state. Bulking factor – the proportional increase in the volume of a mass of in situ rock when it has caved or bulked, usually represented by the symbol, B, so that an in situ volume, V, becomes a volume of V(1+B) on bulking. Bull nose – a convex pillar or surface of less than 90o in plan at the entry to a drawpoint. Camel back – a convex pillar or surface of more than 90o in plan at the entry to a drawpoint. Caveability – a measure (often non-quantitative) of the ability of an orebody to cave under particular circumstances. Cave back or crown – the under-surface of the in situ, but possibly disturbed, rock above a cave or pile of caved ore. Cave inducement – the process of inducing caving by some technique in addition to, or other than, undercutting (eg drilling and blasting or hydraulic fracturing). Cave initiation – the process of the initiation of “natural” caving by undercutting and drawing some of the broken ore. Cave propagation – the process of propagation of an initiated cave by the progressive drawing of broken ore in a planned and controlled manner. Caving – the process of the detachment of in situ rock from the cave back. Caving rate – the average rate, usually expressed in mm per day, at which the cave naturally propagates upwards. Chimney caving or chimneying – the progressive migration of an underground cavity through the overlying material to the surface, usually over a restricted plan area. Continuous subsidence – the smooth or even downwards movement of the surface over a mined area.


467

Crown pillar – a horizontal pillar of unmined rock left above a caved or mined-out area. Dilution – the entry and mixing with the target ore of cap or country rock or unwanted lower grade material. Discontinuity - a collective term used for mechanical breaks or discontinuities in rock masses having low or zero tensile strengths. Discontinuous subsidence – the downwards movement of the surface above a mined area producing a surface profile that is not smooth or even but contains breaks or steps involving vertical or sub-vertical surfaces. Draw – the process of extracting caved or broken ore from a cave or stope. Draw column – the vertical, approximately cylindrical, column of caved or broken ore, drawn through a single drawpoint or a group of adjacent drawpoints. Draw cone – the shape of the surface defining the limits of movement of particles on an initially horizontal line through the caved or broken ore as draw occurs progressively through a discharge point. Draw control – the process of controlling the amounts of ore drawn from individual drawpoints in order to achieve a number of mining objectives. Draw rate – the rate, often expressed as tonnes per day or shift, at which caved or broken ore is drawn from individual drawpoints or a group of adjacent drawpoints. Draw zone – the zone of caved or broken material that will eventually report to a particular drawpoint during progressive draw. Drawbell – the excavated structure, ideally having the shape of an inverted bell, which channels caved or broken ore to a drawpoint. Drawpoint – the excavated structure on the extraction or production level through which the caved or broken ore is loaded and removed from the cave. Drawpoint drift – an opening or drift through which a drawpoint is accessed from a production drift on the extraction or production level. Drawpoint spacing – the spacing in plan between like points in adjacent drawpoints, defined in a number of possible ways.


468

Ellipsoid of motion, draw or extraction – an approximately ellipsoidal zone within a mass of caved or broken ore or rock from which all the material discharged from a bottom outlet after a given period of time will have originated. Extraction or production level – the level in a caving mine through which the caved or broken ore is extracted and transported away from the cave. Extraction level layout – the arrangement of production and drawpoint drifts, drawpoints and other excavations on the extraction or production level. Flow ellipsoid – the approximately ellipsoidal shape of the volume of caved or broken material which moves or “flows” when a bottom outlet is opened or draw occurs. Fragmentation – the process, or the result, of blasting, caving and draw of initially in situ rock or ore. Front caving – a variant of block caving in which the cave is retreated on one or more levels from an initiating slot in the centre or on the boundary of the block. Geological Strength Index (GSI) – a rock mass classification scheme developed by Dr E Hoek as an index for use in estimating the strength and deformability of jointed rock masses. Gravity draw – the drawing of caved or broken ore by gravity, usually through a series of finger raises and ore passes connecting the cave to a lower transportation level. Gravity flow – the flow of broken material under the influence of gravity. Grizzly – a structure made of parallel steel beams or bars through which broken rock of up to a maximum size may pass under the influence of gravity. Hangingwall caving – the usually progressive failure or caving of the hangingwall rock overlying the inclined boundary of a mined orebody. Hangup – the trapping or wedging of one or more usually large blocks in or above the drawpoint or drawbell such that they will not move further by gravity. Haulage level – the level in an underground mine through which the ore and waste is transported away from the production area to crushers or for transport to the surface.


469

Hazard – a potential occurrence or condition that could lead to injury, delay, economic loss or damage to the environment. Height of the interaction zone (HIZ) – the vertical height at the base of the cave within which adjacent draw zones interact and produce lateral motion of caved ore. Hydraulic radius – in the context of underground mining, the ratio of the surface area to the perimeter of an excavated surface such as the roof of an undercut. Inclined drawpoint caving – a form of cave mining in which the drawpoints are developed not on a horizontal but on an inclined plane, often parallel with the footwall of an inclined oredody. Inclined undercut – a form of undercut in which the floor and roof are inclined at significant angles to the horizontal and which takes a chevron or zig-zag shape in vertical section. In situ fragmentation – the fragmentation of the in situ rock before it has been disturbed by undercutting or caving. In situ stress – the state of stress in the in situ rock before it has been disturbed by construction or mining activity. Interactive flow – a mechanism by which the broken ore or rock in adjacent draw columns may migrate from one draw column to the other. Isolated draw zone (IDZ) – a draw zone isolated from other draw zones as a result of drawing from an isolated drawpoint. Limit ellipsoid or loosening ellipsoid – the approximately ellipsoidal mass of broken ore or rock that will have moved, but not necessarily been discharged, when a bottom outlet is opened or draw occurs. Major apex – the shaped structure or pillar above the extraction level formed between two adjacent drawpoints but separated by the extraction drift. Mass flow – the mechanism by which a volume of broken ore or rock moves downwards uniformly and vertically during draw.


470

Mining Rock Mass Rating (MRMR) – a rock mass classification scheme developed by Dr D H Laubscher for use in mining, including block caving, applications. Minor apex – the shaped structure or pillar formed between two adjacent drawbells on the same side of the extraction drift. Mud rush – a sudden inflow of mud from a drawpoint or other underground opening. Narrow flat undercut – a flat undercut whose vertical height is limited to approximately that of the drill drifts, often in the order of 4 m. Narrow inclined undercut – an undercut that is both inclined and narrow. Panel caving – a form of cave mining in which the orebody is undercut and caved progressively in a series of usually parallel panels. Particle Flow Code (PFC) – a numerical analysis code based on Dr P A Cundall’s distinct element method in which the material is represented as an assembly of discrete particles and permitted to flow according to the laws of Newtonian mechanics. Plug subsidence – a form of discontinuous subsidence in which a plug of material overlying an underground opening subsides into the opening suddenly. Post- or conventional undercut – an undercutting strategy in which the undercut is mined after the development of the underlying extraction level, including the drawpoints, has been completed. Pre-undercut – an undercutting strategy in which the undercut is mined before the development of the underlying extraction level and the drawpoints. Primary fragmentation – the fragmentation defined by the blocks in the vicinity of the cave back as they separate from the cave back when the undercut is mined and caving is initiated. Production drift – one of the major set of parallel excavations or drifts on the extraction or production level through which the drawpoint drifts are accessed and ore is transported away from the drawpoints. REBOP (Rapid Emulator Based On PFC3D) – a numerical analysis code used to simulate the flow of broken ore or rock in draw columns and through drawpoints.


471

Reinforcement – a means of improving the overall properties of a rock mass from within the rock mass by techniques such as rock bolting and cable bolting. Risk – formally the product of the probability of occurrence of a hazard and the magnitude of the consequences of that occurrence. Risk analysis – a structured process which identifies both the likelihood and the consequences of the hazards arising from a given activity or facility. Risk assessment – the comparison of the results of a risk analysis with risk acceptance criteria or other decision parameters. Roadway – generally, a long excavation used for transportation in an underground mine. In caving mines the term is sometimes used to describe the surface on which vehicles travel on the extraction or production level. Rock burst – the uncontrolled disruption of rock associated with a violent release of energy. Rock mass characterisation – the systematic process of describing a rock mass both quantitatively and qualitatively for engineering purposes. Rock mass classification – a method of assigning numerical values to a range of characteristics considered likely to influence the engineering behaviour of a rock mass and of combining these values into one overall numerical rating. Rock Mass Rating (RMR) – a rock mass classification scheme on a scale of 0-100 developed by Dr Z T Bieniawski. Secondary breaking – the additional breaking of caved or broken ore or rock by mechanical means or by explosives to reduce large fragments to the sizes required for loading and transportation. Secondary breaking may be required to treat hangups. Secondary fragmentation – the fragmentation produced in caved ore or rock during residence in the ore column and draw. Shape factor – the ratio of the area to the perimeter of an excavated surface, ie the hydraulic radius. Slusher – a mechanically operated, chain pulled scraper which moves caved or broken ore from a row of drawpoints to a grizzly or ore pass.


472

Stacking – the collection or stacking of caved or broken ore or rock on the undercut level or on the major apices. Stress caving – a caving mechanism in which the in situ rock fractures and caves under the influence of the stresses induced in the cave back. Subsidence caving – a caving mechanism in which a large mass of rock subsides rapidly as a result of shear failure on the vertical or near-vertical boundaries of a block. Support – the application of a reactive force to the surface of an excavation to control the deformation of the rock mass. Sometimes support is taken to include reinforcement as defined above. Surface subsidence – the lowering of the ground surface following underground mining. Undercut – the approximately horizontal slot mined to initiate caving of a block or panel. Undercut level – the level of the bottom of the undercut from which the mining of the undercut takes place. Undercutting rate – the rate, expressed as an area per unit time (eg m2 per month), at which the undercut front is advanced. Void diffusion – a mechanism of the irregular flow and intermixing of ore or rock from adjacent draw zones which involves the upwards migration, filling and/or collapse of voids formed by large, angular particles. Wet muck flow – the sudden collapse and rapid run-out of wet granular material, usually from a drawpoint. Zone of influence – generally, the zone surrounding an excavation in which the pre-existing stresses are perturbed by a given amount (often taken as 5%) by the presence of the excavation. In block caving, the term is also used to describe the zone surrounding the cave in which the rock mass response is influenced by the presence of the cave.

473

APPENDIX B

RELATION BETWEEN CAVED COLUMN HEIGHT AND VERTICAL STRESS AT THE CAVE BASE

B.1 INTRODUCTION

he study reported here was carried out by Dr Loren Lorig of the Itasca Consulting Group, Inc, Minneapolis, Minnesota, USA, as part of the International Caving Study Stage I to support further development of the expert system BCF developed by

Esterhuizen (1994). The results of this study were incorporated into the revision of the BCF program outlined in Chapter 4 (Esterhuizen 1999). The content of this Appendix is taken almost verbatim from Lorig’s report of the study in the Final Report of the International Caving Study Stage I. As was noted in Chapter 4, secondary fragmentation in BCF is based in part on the estimated stresses acting at the cave base. One module in this program considers a block of caved material above the cave base and estimates what percentage of the block’s weight is transmitted to the base. BCF incorporates an estimation of the vertical stress (referred to as pressure) transmitted for plane conditions (when two of the sides of the assumed rectangular base have infinite length). In the 1994 version of BCF this estimation was based on the results of sand model tests conducted for Southern African chrysotile asbestos mines (Heslop and Laubscher 1981). The sand model had planar vertical sides made of steel. These boundary conditions result in a low friction angle that may over predict the vertical stress acting on the cave base. If the sides are frictionless, then 100% of the caved column height acts on the base - ie, the vertical stress is the product of the density, gravity and height. One objective of Lorig’s study was to make a quantitative estimation of transmitted vertical stress for the plane strain condition using limit equilibrium analyses and numerical models. A second objective was to extend the analysis to the more general case in which the base area is considered to have finite dimensions. A fundamental assumption inherent in the analyses is that the caved material is drawn uniformly, resulting in a relatively uniform vertical stress acting on the base. It is important to note that uniform draw is seldom achieved in practice. Vertical stresses can be much higher in areas without draw than in areas that are being actively

T

Appendix B – Relation between Caved Column Height and Vertical Stress at the Cave Base

474

drawn. This behaviour results from shear stresses at the margins of actively drawn areas that add vertical stresses to the areas not being drawn. The problem to be analysed is represented in Figure B.1. A ‘prism’ of caved material is assumed to develop above the cave base. The rock in its caved condition is assumed to be cohesionless (ie, c = 0), with a certain friction angle, φ, and a unit weight, γ. The base area has a width, B, and (out-of-plane) depth, D; the height of the ‘caved’ prism is H. In the simple model represented in Figure B.1, the total weight of the prism, W, is equilibrated by frictional forces that develop on the lateral walls (the resultant forces, F, represented in the figure) and the pressure, p, at the base. The frictional forces, F, are assumed to depend on the confinement (forces N) that develops in the direction perpendicular to the walls. The equilibrium of the ‘prism’ (and, consequently, the pressure transmitted to the base) is conditioned by the ‘constraining’ effect imposed by the lateral confinement. The specific goal of this study was to analyse quantitatively the relationship between the base pressure, p, and the geometry of the prism (eg the size of the base, the ratio between width and height, etc). For example, it can be expected that the greater the height of the prism, the larger will be the area over which the frictional stresses will act (on the sides of the prism) and therefore, the lower the percentage of pressure, p, transmitted to the base. Studying the dependence of the out-of-plane dimension (length D) on the pressure transmitted was of particular interest. The simplest approximation, which will be referred to as the two-dimensional analysis, considers a prism of infinite length in the out-of-plane direction (the mechanical equivalent of plane strain conditions). A better approximation, which will be referred to as the three-dimensional analysis, takes the third dimension, D, into consideration.

γ

p

F

BD

H

extractionlevel

c=0=35φ

o

W

N

Figure B.1: Definition of the problem


475

B.2 LIMIT EQUILIBRIUM MODELS

As a first approximation, the problem can be analysed using limit equilibrium methods. These methods have been applied in the past to the study of the roof stability in excavations in frictional materials (eg Terzaghi 1943). For example, Figure B.2a shows the limit equilibrium model used by E gger (1983) to study the stress distribution in the rock mass above the roof of a shallow tunnel. Note that there is some resemblance between the problem solved by Egger and the problem described in the previous section (compare Figures B.2a with Figures B.1 and B.2b). In a limit equilibrium model, the material constituting the prism of Figure B.2 is assumed to be in a failure (or plastic) state. Stating the force equilibrium for ‘layers’ of differential height dz (as shown in Figure B.2b), a relation between the mechanical and geometrical variables can be derived. In the sub-sections below, both two- and three-dimensional limit equilibrium models will be considered.

(a) (b)

Figure B.2: (a) Limit equilibrium model by Egger (1983), and (b) equilibrium analysis for a layer of infinitesimal height above the cave base

Two-dimensional case

Consider the prism of caved material that has an infinite length, D, as shown in Figure B.2b. If this caved mass is assumed to be cohesionless with a certain friction angle, φ, and a unit weight, γ, the vertical stresses, σz, and the horizontal stress, σx, can be written as


476

( )z H zσ γ= − (B.1)

x p zKσ σ= (B.2)

where 1 sin1 sinpK φ

φ+

=−

.

Equations B.1 and B.2 consider the ‘active’ type of plastic failure; in this case, the prism in Figure B.2b achieves the (critical) equilibrium limit due to a reduction of the support pressure at the bottom.

The differential expression of equilibrium for the layer of infinitesimal height shown in Figure B.2b results, then, in

2 tan 0zz p

d Kdz Bσ σ φ γ− + =

(B.3)

Solution of Equation B.3 with the boundary condition σz(0) = p leads to

2 tan( )

2 tan 2 tan

pK zB

zp p

B Bp eK K

φγ γσφ φ

⎛ ⎞= + −⎜ ⎟⎜ ⎟

⎝ ⎠

(B.4)

When the upper surface of the prism in Figure B.2b is considered stress-free, the relation between the normalised base pressure, p/(γB), and the ratio H/B can be obtained from Equation B.4, by making σz(H) = 0 ,


477

2 tan1 1

2 tan

pH KB

BP eHHB

φ

γ φ

−⎛ ⎞⎛ ⎞⎜ ⎟⎜ ⎟= −⎜ ⎟⎜ ⎟

⎜ ⎟⎜ ⎟⎝ ⎠⎝ ⎠

(B.5)

Three-dimensional case

A similar analysis to that presented in the previous sub-section can be carried out for a prism that has a length, D, in the out-of-plane direction (see Figure B.1). The differential expression of equilibrium, Equation B.3, now takes the form

2 tan 0zz p

H

d Kdz Rσ σ φ γ− + =

(B.6)

where RH is the hydraulic radius of the base region,

2( )HBDRB D

=+

(B.7)

Solving Equation B.6 with the boundary condition σz(0) = p leads to

tan

tan tan

p

H

K zRH H

zp p

R Rp eK K

φγ γσ

φ φ⎛ ⎞

= + −⎜ ⎟⎜ ⎟⎝ ⎠

(B.8)

Assuming the upper surface of the prism to be stress-free and combining the condition σz(H) = 0 in Equation B.8 with Equation B.7 gives


478

tan1 1tan

pH

H KR

pH

p eHH KR

φ

γ φ

−⎛ ⎞= −⎜ ⎟⎜ ⎟

⎝ ⎠

(B.9)

Note that, in the limit, when D tends to infinity, the hydraulic radius given by Equation B.7 becomes RH = B/2, and the same relation as in Equation B.5 for the two-dimensional case is recovered.

Interpretation of results

Equations B.5 and B.9 give the proportion of (prism) weight transmitted to the base area for both the two- and three-dimensional cases considered. These expressions are represented in Figures B.3a and B.3b, respectively. The percentage of weight transmitted to the base, expressed as the ratio p/(γH), depends on H/B and H/RH, respectively. Note that, according to Equations B.5 and B.9, the two- and three-dimensional solutions could have been represented by a single curve if the ratios 2H/B and H/RH had been considered instead. (This is because, in the limit, when D tends to infinity in the three-dimensional model, these two ratios must be equal).

As will be shown, values of base pressure computed from these limit equilibrium models result in smaller values than those obtained from plasticity continuum models (eg with FLAC (Itasca 1998a)). Despite the limitations that the limit equilibrium models may have (with regard to their ability to make accurate predictions of loads transmitted), it appears useful to employ them as starting points for more elaborate analyses. In this sense, it is interesting to note that the relation between normalised base pressure (p/γH) and the ratios H/B and H/RH, as represented in Figure B.3, obey a power law. As will be seen, numerical results obtained with FLAC and FLAC3D (Itasca 1997b) appear to follow the same law.

The curves also confirm the hypothesis mentioned first: as the H/B ratio increases (or the frictional lateral surface increases in size), the pressure transmitted to the base decreases. This is suspected to be because of the increasing importance of the ‘lateral-constraint’ effect.


479

0 1 2 3 4 51

10

100

0 5 10 15 20

10

100

2

a)

b)

p

γH

p

γH

HB

two-dimensional model

three-dimensional model

[%]

[%]

RH

H

Figure B.3: Results for the (a) two-, and (b) three-dimensional limit equilibrium models B.3 NUMERICAL EXPERIMENTS USING FLAC AND FLAC3D

Numerical models based on the continuum codes FLAC and FLAC3D were used to estimate the pressure transmitted to the base area. Figures B.4a and B.4b show the layouts for the FLAC and FLAC3D models, respectively. The models were set up simulating the conditions of the limit equilibrium models (ie cohesionless Mohr-Coulomb material, 35o friction angle, an initial plastic state of stresses with the vertical stress given by a lithostatic gradient, etc). The same model was then solved for decreasing values of base pressure, p (see Figure B.4a). For each run, the velocity history at the point A (the intersection of the symmetry axis and the base area) was recorded. The critical base pressure corresponding to a limit equilibrium state could then be determined.


480

The results obtained from these runs were analysed in statistical terms and compared with the values used in BCF and those obtained by the limit equilibrium models. These results are discussed in the following sub-sections.

sym

met

ry a

xisa)

b)

fixed

fixedundercut

H

H

B/2

B/2D/2

p=k Hγ

lithostaticgradient of

initial stresses

fixed block

fixed block

undercut

movingblock

symmetry axis

Case shownH/B=1.5

Case shownH/B=1.5B/D=2

A

A

Figure B.4: (a) FLAC, and (b) FLAC3D models used in evaluation of the pressure transmitted to the base area


481

Results for the two-dimensional case

As was noted in Section B.1, the first goal of this study was to compare the values of base pressure used by BCF and those obtained from mechanical models. Figure B.5 summarises the results for the two-dimensional case. It is clearly seen that the values of pressure obtained from a limit equilibrium model (Figure B.3a) are lower than those obtained with FLAC. The FLAC model appears to show some mesh dependency, particularly for relatively low values of B/L; however, as the B/L ratio increases, the mesh dependency disappears (note, for example, that the FLAC results are the same for B/L=20 and 40). It can also be seen that the values used by BCF are higher than those corresponding to limit equilibrium models and FLAC.

0 2 4 6

1 0

1 0 0

p p

γHγH

HB

HB

5

[%]

Limit equilibriummodel

FLAC

(B/L=10)

(B/L=20 and 40)

BCF(empirical)

C

B

A

Curves A, B, C are power functionsof the form

=

-m

q

for A, q=34 m=0.75 B, q=44 m=0.57 C, q=80 m=0.83

Figure B.5: Analysis of results for the two-dimensional models (L is the size zone used

in the FLAC models)

A regression analysis was performed on the scattered data available. It is quite noticeable that the results obtained with the different methods appear to obey a power law. In particular, the FLAC results, assumed to be rigorously computed mechanical solutions, are fitted by the equation,

0.57

44p HH Bγ

−⎡ ⎤= ⎢ ⎥⎣ ⎦

(B.10)


482

In order to provide a better estimation for the base pressure transmitted in the two-dimensional model, the BCF values are compared with those obtained from Equation B.10 in Table B.1.

Table B.1: Comparison of values used in BCF (Esterhuizen 1994) and results from the

FLAC model

Width-to-height ratio of the draw column

% of caved weight acting at the bottom of column

BCF FLAC

1:1 80 44

1:2 45 30

1:3 33 23

1:4 25 20

1:5 21 18

Results for the three-dimensional case

Figure B.6 summarises the results obtained with FLAC3D. It can be seen, again, that the limit equilibrium results in Figure B.3b correspond to lower values of pressure than do the numerical results.

0 5 10 15 20

10

100

p p

γH γHHR

H

2

[%]

FLAC 3 D

Limit equilibrium

A

Curve A is a power functionof the form

=-m

q

for A, q=96 m=0.99

RH

H

Figure B.6: Analysis of results for the three-dimensional models (ratio L/B = 20 was

used in the FLAC3D models)


483

The regression analysis performed on the FLAC3D values gave the relation:

0.99

96H

p HH Rγ

−⎡ ⎤

= ⎢ ⎥⎣ ⎦

(B.11)

Equations B.7 and B.11 allow Table B.1 to be extended for rectangular base geometries (Table B.2). Table B.2: Percentage of caved weight acting at the bottom of column for base aspect

ratios of 0, 1 and 2

Width-to-height ratio

of the draw column

% of caved weight acting

at the bottom of column

B/D = 0

(FLAC)

B/D = 1

(FLAC3D)

B/D = 2

(FLAC3D)

1:1 44 27 15

1:2 30 14 8

1:3 23 9 6

1:4 20 7 4

1:5 18 6 4

B.4 DISCUSSION OF RESULTS

Limit equilibrium models lead to the lowest values of pressure transmitted to the base. The 1994 version of BCF, on the other hand, used the highest values; in between are the numerical results obtained with FLAC and FLAC3D. As noted in Section B.1, the actual vertical stress acting at the base of any caved area is probably heterogeneous. It is reasonable to expect the the stresses at the base to have some influence on secondary fragmentation. However, it is doubtful that any estimate based on uniform draw can be used to obtain anything other than a general idea of stress levels. Secondary fragmentation occurring near the cave base would seem to depend on local conditions occurring above the drawpoint. For example, PFC3D (Itasca 1998a) models have demonstrated that very high contact forces can result when hangups occur above the extraction point (see Figure B.7). It has not been demonstrated that these forces are directly related to the vertical stresses acting at the base of the caved volume.


484

In a related matter, the vertical stress acting at the base of the caved volume is important in defining the loading of extraction level pillars during the operational phase. There is little information available regarding this important loading condition.

Figure B.7: Close-up of the draw cone region in a PFC model after hangup occurs (The black lines between particles indicate the directions of the contact forces. The line thickness indicates magnitude, in this case, the maximum magnitude is 607 kN)

B.5 CONCLUSIONS

The magnitude of the average vertical stress acting on the base of a caved region is used in BCF to estimate the contribution to the fragmentation process that occurs just above the drawpoint. The numerical and limit equilibrium results presented here suggest that the vertical stresses used in the 1994 version of BCF may have been overestimated. Estimating the contribution to secondary fragmentation that occurs near a drawpoint is made difficult because • it is impossible to measure the history of forces acting on rock particles as they travel

downwards to the drawpoint;


485

• it is impossible to measure the stresses acting within a caved rock mass; and • at present, secondary fragmentation is thought to result from several processes (see Section

4.2) that are difficult to separate and quantify. Although, in principle, PFC3D could be used to address these difficulties, it is likely that many simulations would be needed to obtain statistically meaningful results. The foregoing suggests that the prediction of secondary fragmentation will remain a difficult and largely empirical process for the foreseeable future. Improvement in predictive capabilities will come through the sound mechanical interpretation of empirical results.

486

APPENDIX C

NUMERICAL SIMULATION OF PARTICLE FLOW USING REBOP C.1 PURPOSE

he mechanics of particle flow and its influence on drawpoint design and draw control practice was discussed in Chapters 6 and 7. As was indicated, much of the current understanding of the flow of caved material during draw comes from laboratory model

studies and field experiments. It was suggested that modern numerical methods may provide a means of developing the improved understanding and knowledge that are clearly necessary if the state-of-the-art of the design and operation of block and panel caving mines is to be advanced. The purpose of this Appendix is to present the results of some simulations of particle flow towards drawpoints carried out by Lorig and Cundall (2000) as part of the International Caving Study Stage I using the codes PFC3D (Particle Flow Code in 3 Dimensions, Itasca 1998a) and REBOP (Rapid Emulator Based On PFC3D, Itasca 2000a). This study did not consider the effects of the migration of fines or secondary fragmentation during draw. This account of the study is taken directly from that of Lorig and Cundall (2000). It should be noted that this account concerns only the initial stage of a study which is continuing as part of the International Caving Study Stage II. PFC3D models three-dimensional assemblies of particles that move and interact under the laws of motion and laws relating to conditions at the contacts between particles. PFC3D has been used to study the flow of particles into a drawpoint during the caving process, but its application is limited to very few drawpoints, because of the large numbers of particles needed. Even a simulation of flow into one drawpoint can take several days to execute on a fast computer. Therefore, Lorig and Cundall (2000) sought to reproduce the mechanisms observed in PFC3D models in a more efficient (although approximate) manner, so that simulations can be performed in a matter of minutes, even for models that contain multiple, interacting drawpoints.

T

Appendix C – Numerical Simulation of Particle Flow using REBOP

487

This Appendix focuses on the code REBOP, which embodies rules based on mechanisms observed in PFC3D simulations and simulates the evolution of isolated draw zones (IDZs) and their interactions. REBOP makes no assumptions about the shape of the IDZs. It contains “micro” rules that govern how material flows from one layer to the next and how much material derives from mechanisms such as erosion and mass exchange between adjacent IDZ catchment volumes. The shape of each IDZ evolves continuously (in contrast to the fixed draw cone of PC-BC), and emerges “automatically” as the micro-rules are applied repeatedly. Potentially, this lower-level approach is closer to reality and allows local mechanisms and interactions to be reproduced. However, the approach is new and as yet untested, apart from the simple examples presented here. It is encouraging to note that apparently realistic IDZ shapes do indeed emerge from the application of quite simple micro-rules, but much more extensive testing is needed before it can be concluded that the approach is valid for the whole range of conditions encountered in caving mines. The results on which the relations used in REBOP are based derive from the physical model tests discussed previously and from simulations performed using PFC3D. Before REBOP is described, the results of some simulations carried out using PFC3D will be presented. The complete set of equations used in the REBOP code is given in the user’s guide (Itasca 2000a). C.2 PFC3D SIMULATIONS

C.2.1 The PFC3D Numerical Model

The PFC3D model consists of a three-dimensional collection of rigid particles or balls. All particles are spherical, but spheres may be rigidly clumped together to form particles of arbitrary shape. PFC3D models the movement and interaction of particles using the Distinct Element Method (DEM) first developed by Cundall (1971). Newton’s laws of motion provide the fundamental relations between particle motion and the forces causing that motion. The force system may be in static equilibrium — in which case, there is no motion — or it may be such that it causes the particles to flow. In addition to spheres, the PFC3D model also includes “walls” that provide boundaries to the models. The spheres and walls interact through the forces that arise at contacts, assuming linear springs in the normal and shear directions. Sliding is allowed in the contact shear direction by limiting the magnitude of shear force to the friction coefficient multiplied by the normal force. A detailed description of the method, including the assumptions made, is given by Cundall and Strack (1979).

C.2.2 PFC3D Approach to Modelling Gravity Flow for Isolated Drawpoints

Two problem geometries have been considered. Most problems involved a conical-shaped draw outlet, but some problems were run with a trench (or hopper) geometry. In all cases, the


488

complete problem geometry was established with planar walls. Comparisons with runs for “bumpy” walls showed little difference in results. A “matter replicator” was used to reduce problem size (and, hence, computational time) by introducing symmetry planes in the overlying material, while preserving full symmetry within the critical area of the draw cone. The matter replicator creates copies of particles as they fall through a particular horizontal plane (at the top of the draw cone). At the instant of replication, the copies are symmetrical, but symmetry is quickly broken as the material flows through the cone. Complete models that did not include the matter replicator gave results that were similar to those obtained using the matter replicator (apart from a significant difference in computer time). Material properties were established by performing simulated triaxial tests on particle assemblies under representative confining stresses. Realistic bulk properties for caved material were obtained when rock fragments were represented as clumps having two or three particles per clump. Although PFC3D can represent clumps containing arbitrary numbers of particles (and, hence, overall shape), there seemed to be no reason to increase complexity at this stage. Furthermore, the effect of clumps breaking into smaller clumps, as a function of the magnitude of the forces acting on the clumps, was not studied, although PFC3D is able to allow this mechanism to occur. In the results presented here, the clumps remain intact with there being no relative motion between constituent spheres.

Measurements during simulations

The items listed below were measured during the simulations. These measurements were used to develop a quantitative understanding of the mechanics of gravity flow. The postulated mechanisms operating in the gravity flow simulations are discussed in a subsequent sub-section. 1. Porosities were measured at various heights within the draw column. An initial porosity of

0.35 was used for most simulations. 2. Profile plots were made at selected stages. The profile plot shows the current locations of

idealised marker bands arranged in a rectangular pattern. Averaging of displacements was used in the circumferential direction for conical outlet geometries.

3. IDZ plots were superimposed on the profile plots as shown in Figure C.1. The IDZ is

defined as a contour of points within the caved rock mass that have experienced a given magnitude of vertical displacement relative to their initial positions. In PFC3D computer plots, this magnitude is taken to be 1.0 m. The conclusions are relatively insensitive to this choice.


489

Simulations of gravity flow for individual drawpoints

Full PFC3D simulations of a single drawpoint show generally similar results to earlier physical model studies. However, in contrast to the model studies, further detailed internal information can also be obtained from the simulations to guide the formulation of the relations used in REBOP. An important characteristic of the flow process is that the IDZ corresponds to a region of high porosity and low average contact force between particles. Figure C.2a shows an early stage in a PFC3D simulation, with the IDZ shown as a white line (denoting the contour of 1-m vertical movement). Figure C.2b shows a representation of the contact forces between particles. The gravity body forces within the IDZ are transferred to the region outside the IDZ, leaving the internal region lightly loaded.

Figure C.1: Typical PFC3D results for quantifying IDZ development (The plot shows the IDZ limit (white line with ellipsoidal shape) superimposed on a profile

plot of idealised marker bands. Development of the 45° surface slope corresponds well with the bulk friction angle of the material of 44°.)


490

Figure C.2: (a) PFC3D simulation of flow into a single drawpoint, and (b) contact forces, shown as black lines of thickness proportional to force magnitude

The rate at which the IDZ propagates upwards towards the surface depends on the difference between the final and initial porosities. The results obtained suggest that the mean value for final porosity is around 0.4 or slightly higher. This value agrees with data collected from Kimberlite mines (Guest 1999). Another important factor is the lateral expansion of the IDZ. Early physical model studies suggested that there is a connection between particle size and the eccentricity of the ellipsoid. PFC3D simulations have shown that particle shape has a strong influence on IDZ width. For example, the result shown in Figure C.3a is for spherical particles, while that in Figure C.3b is for particles of the same volumetric size distribution, but with aspect ratios of 1.5 to 1 (using the clump logic of PFC3D to bind pairs of spheres together). Lorig and Cundall (2000) postulated that the mechanism of erosion is responsible for the widening of the IDZ with increasing extraction. If a continuum simulation is performed using a frictional material, the IDZ is almost vertical. In contrast, for a material that consists of discrete particles, individual particles in proximity to the shearing interface at the IDZ boundary may be torn away from the intact matrix. This mechanism might resemble the erosion of particles in the bank of a flowing river. The contrast observed between the behaviour of spheres and clumps could be explained by the relative ease of detaching a sphere, compared to that of detaching a “peanut-shaped” particle, which can be interlocked with its neighbours.


491

Figure C.3: IDZ locations for PFC3D simulations for assemblies consisting of (a) spheres, and (b) 2-ball clumps of 1:1.5 aspect ratio.

C.2.3 PFC3D Simulations with Four Drawpoints

The effects of interaction between adjacent drawpoints were studied using models having four adjacent drawpoints. These models used the previously described matter replicator to reduce the problem size. The results presented here are for constant and equal draw in all drawpoints. In one set of runs, the tops of the cones have radii of 5.73 m and are separated from their closest neighbours by 1.54 m (1.54 + 2 * 5.73 = 13 m pitch). The distance across the diagonal between cone centre lines is 18.38 m. Therefore, the distance between cone rims across the diagonal is 6.92 m. In dimensionless form, the ratio of separation (between rims) to cone radius is either 0.27 or 1.21, depending on the path. Figure C.4 shows that there is complete interaction between the drawpoints, with plug flow occurring essentially over the whole height (ie there is no dead zone). In a second model (see Figure C.5), the drawpoints were spaced 18 m apart. The separation ratios here (using similar logic to the above) are 1.14 and 2.44. In this case, there is interaction only at some height above the draw level (depending on how well developed is the IDZ). At a late stage in the test, plug flow is seen at a height of about 1.5 times the cone radius above the draw level (below this, there is a dead zone).


492

Figure C.4: Four-drawpoint model with 13 m separation between drawpoints showing complete interaction (plug flow)

Figure C.5: Four drawpoint model with 18 m separation between drawpoints showing poor interaction


493

In summary, the flow pattern may be compared with that of a single drawpoint run with similar conditions (eg Figure C.1), and it is found that the IDZ only extends laterally to about 50% of the cone radius. Thus, there appears to be some interaction between IDZs in the model shown in Figure C.5, which, taken separately, would not intersect. In the model shown in Figure C.4, adjacent IDZs would be expected to touch (or nearly touch) soon after drawing begins, so the observed strong interaction is not surprising.

C.2.4 PFC3D Component Tests

“Component tests” were also performed to provide data on the erosion mechanism. Component tests attempt to reproduce selected parts of the full caving process, with lower computational costs. One such test was similar to the full PFC3D draw simulation, but the conical drawpoint object was eliminated and replaced by a prescribed-velocity boundary condition on particles on the base of the model. A linear velocity profile was imposed on these particles, with maximum velocity on the centre line and zero velocity at the radius of the opening. A 90° sector in plan was modelled, with frictionless walls acting as lateral constraints. Both opening radius and mean particle radius were varied, with particles represented by 2-sphere clumps of 1:1.5 aspect ratio. Notably, no systematic dependence of IDZ width on particle size was seen. The IDZ widths are almost identical, with that of the coarse material being, if anything, slightly smaller than that of the fine material. This is in contrast to a widely accepted finding of Janelid and Kvapil (1966), but it is in agreement with that of Peters (1984). The component tests also show that the rate at which the IDZ expands upwards appears to decrease strongly with friction angle. However, the maximum stable porosity also increases with friction angle. Thus, the porosity-jump (which occurs at the IDZ boundary) is greater for higher friction materials, and therefore, for a given extracted mass, the IDZ expands upwards at a lower rate than it does for lower friction materials.

C.3 THE PHYSICAL BASIS OF REBOP

C.3.1 Overview

Incremental relations that can mimic the micro-mechanisms acting in a PFC3D model of flowing granular material are essential to the development of a rapid simulator. It is not adequate simply to invent functions that fit the overall IDZ and deformation shapes, as was done by Janelid and Kvapil (1966), who based their calculations on the assumption that the IDZ is an ellipse. To model the interaction between neighbouring drawpoints and to be predictive in situations that have not been simulated previously, incremental equations are needed to describe how material migrates from point to point, rather than equations (eg for ellipses) that have been chosen to fit some observed overall shapes. The starting point is that mass is conserved. Mass balance equations can be written to govern movement of material from one layer to the next,


494

but equations are still needed that quantify the contributions from various sources. The mechanisms that are revealed by PFC3D simulations are reviewed first.

C.3.2 Mechanisms Observed in PFC3D Simulations

PFC3D results show the IDZ developing in a characteristic way. For example, Figure C.2a shows an early stage in a simulation for an assembly of spherical particles. The IDZ is equivalent to the “limit ellipsoid” or “ellipsoid of loosening” of Janelid and Kvapil (1966). Their “ellipsoid of motion” would consist of the locus of all points in the initial assembly that have moved into the cone. Behaviour similar to that in the full PFC3D simulations was observed in the “component tests” discussed previously. The following observations may be made. • The lines denoting the IDZs are the boundaries between high porosity material in the

flowing region and low porosity material in the undisturbed region. • The amount of material flowing through given horizontal cross-sections decreases with

height. The mean angle of the coloured marker lines in Figure C.2a decreases with height above the base.

• The IDZ consists of two main parts, a nearly vertical lower section, and a tapering upper

section that intersects the centre line at some finite angle. • The lower part of the IDZ bulges out, deviating from a vertical line. C.3.3 Postulated Mechanisms

Lorig and Cundall (2000) postulated that two main mechanisms combine to account for the shape of the IDZ. 1. When material is removed from one level, material at the next higher level “collapses,” so

that the high porosity area in the upper region expands. The area expansion depends on the difference in area (projected onto horizontal planes) between the two levels. When the areas are identical, material simply translates downwards: mass flow in the two slices is identical. When the upper level area is less than that of the lower level, material is passed to the lower level from the expanding volume of high porosity material in the upper layer.

2. The first mechanism would result in a vertical “wall” to the IDZ (since the mechanism is

only activated when neighbouring areas are different). However, at the interface between moving and stationary material, particles get dislodged and enter the flowing mass. Thus, the lateral IDZ boundary expands as a function of the local flow rate. The mechanism is


495

similar to the erosion of a river bank in which particles are progressively torn away from the bank by the flowing water.

The numerical experiments also provide evidence for the collapse mechanism that operates at the active IDZ front. The fact that the active front curves inwards towards the centre line means that a given layer is “undercut” at its outer edge by material withdrawn from the larger area of the layer below. It is proposed that the overhanging part of the upper layer partially fails (thus releasing mass, due to porosity increase) near the outer edge of the layer. This gives rise to the jump in porosity at the active front. Initially local, incremental relations that can reproduce the evolution of shape of a single IDZ were considered. Then, the logic was extended to interacting IDZs.

C.3.4 Single IDZ Evolution Relations Used in REBOP

Simple incremental relations were proposed that embody the collapse and erosion mechanisms, denoted by M1 and M2, respectively. By applying the relations to a multi-layered system, it can be determined if they lead to IDZ shapes similar to those observed in PFC3D simulations. In the initial version of REBOP (1.0), the cross-section of the IDZ is taken to be circular. It is planned, in later versions, to replace circular cross-sections with polygons, which can represent the non-uniform geometry at intersections of adjacent IDZs. Assume that the system can be represented by a series of horizontal layers and the IDZ by a series of circles, one for each layer. Consider a representative layer, i, in which the mass flow rate (to the layer below) is im& . This flow is supplied by: (a) the flow from the next layer above, 1+i ; (b) the extra mass contained in the expansion of layer 1i + due to collapse (M1); and (c) the extra mass supplied by erosion (M2). Hence,

( )( )( )21 M

1iM1i1ii m m m m +++ ++= &&&& (C.1)

A series of time increments (each of magnitude tΔ ) is executed and the layers are scanned from the lowest to the highest. The maximum incremental mass that can be passed from layer

1i + to layer i is maxi im m tΔ = Δ& , since the flow rate in layer i is already fixed at im& . The

incremental mass derived from the collapse and erosion mechanisms is expressed as a fraction of max

imΔ . The incremental relations describing the collapse and erosion mechanisms are described in detail in the REBOP user’s guide (Itasca 2000a). As soon as the IDZ intersects the free surface, the porosity within the IDZ remains relatively constant, and the mass balance is satisfied by (a) a mean reduction in elevation of the free surface, and (b) a lateral expansion of the IDZ, resulting in a greater volume of material with increased porosity.


496

C.3.5 Interaction Between Adjacent IDZs

The term IDZ strictly refers to the three-dimensional surface at which there is a porosity jump (from the in situ porosity, 0n , to the maximum stable porosity, 1n ) for a single drawpoint. However, the term will continue to be used to denote similar surfaces arising from multiple drawpoints. As the IDZs of neighbouring drawpoints approach each other, there is a point at which interaction occurs. Assume, initially, that two IDZs act independently until they actually touch. There can be no further expansion of the IDZs in the region of touching, because the material available for expansion is already in its high-porosity state, 1n . Therefore, the contribution,

within a layer, from IDZ expansion to mass flow is zero over that part of the IDZ in contact with its neighbour. In this region, the mass flow from the lower layer is simply passed to the upper layer, ie there is plug flow. C.3.6 Material Transport: Mixing

It appears from tests with PFC3D in which the full process of flow is considered (see Section 7.6.2), that the velocity profiles above a single drawpoint can be approximated as a series of inverted cones, with maximum velocities on the centre line decreasing linearly with radius to the IDZ boundary. In order to account for the correct mix of material types arriving at the extraction point, we need to represent the flow paths of distinct material fragments. To do this a cubic array of “markers” is superimposed on the initial volume of material. Each marker represents the volume of rock originally surrounding it. Markers are moved according to the local velocity field at each instant in time, carrying with them property information (eg grades and physical properties) of the associated material. Thus, material flowing from drawpoints also has associated markers, which allow the particular mix of material types to be calculated. The precision of this calculation depends on the initial density of markers chosen. Markers can also be used to modify the properties of each layer used in the mass flow calculations (such as density) if this is important. When a marker enters the drawpoint, it is considered to be “extracted,” and its associated mass is recorded, together with the grade values that it carries. The mass associated with a marker is given by: 0(1 )mark mark sm V n ρ= − (C.2)

where markV is the volume associated with the marker (typically 3S , where S is the original distance between markers in a cubic array). The initial porosity, 0n , and solid density, sρ , are the values that existed at the original location of the marker. At any stage in the simulation, graphs may be plotted of the total mass extracted and the proportions of the various grades, as calculated by the accumulated masses of the corresponding markers. A coarse grid of markers


497

can be chosen for a rapid, approximate solution (“noisy” graphs), while a fine grid can be chosen for a lengthy, but more accurate, solution (smooth graphs). An illustration of marker movement at an early stage in drawing is provided in Figure C.6. The density of markers is increased for this example, and markers are shown only if they lie within a thin slice parallel to the plane of the paper. The movement can be seen clearly near the top of the IDZ region, but lower down, the plot is confused by markers entering the plotting slice from “out-of-plane” locations.

Figure C.6: Marker positions shown at an early stage in drawing Note that the need for "mixing rules," as used in PC-BC, is avoided by the use of markers. The non-uniform motion of the markers means that a mixture of material, originating from different locations, arrives at a given drawpoint simultaneously. In REBOP, the specification of mixing rules is replaced by the specification of a flow field. It is believed that the latter is easier to determine and verify than the former, because flow fields are directly observable in experiments, whereas mixing rules must be inferred from indirect observations.


498

C.4 EXAMPLE REBOP SIMULATION: SINGLE DRAWPOINT

A PFC3D simulation of a single drawpoint was performed in which the overburden consisted of two layers, the lower containing ore and the upper containing waste (see Figure C.7 for the initial stage and Figure C.8 for the stage in the draw process at which about 10% of the total material has been extracted). Figure C.9 provides a history of the material produced. Dilution by waste is seen to start at about a 10% extraction ratio. The upper radius of the conical drawpoint was 5.6 m, the overall height of the overburden was 45.75 m, and the bottom of the waste was at 25.1 m, measured from the top of the drawpoint cone. The mean initial porosity was approximately 0.36, the particles were spheres, and the friction coefficient for contact between particles was 1.0 (equivalent to a bulk friction angle of about 33°.)

Figure C.7: Initial state of model (waste material shown in grey)

The simulation was reproduced with REBOP, using the same values for intact density, porosity, friction angle, drawpoint radius and heights of ore and waste. The draw rate was constant at 40 tonnes per day. Figure C.10 shows the plot of extraction histories from REBOP at approximately the same stage as Figure C.9, ie at double the time at which the first dilution is seen. Not only is the form of the curves similar, but the numerical values of dilution (expressed as the ratio of waste mass to total mass) at the final stages of both simulations are almost identical, as well. The IDZ profile calculated by REBOP at 145 days is shown in Figure C.11, and is similar to that of the PFC3D simulation shown in Figure C.8. It may be noted that the PFC3D simulation took about 2 days to execute on a 1-GHz Pentium machine, while the REBOP calculation executed in a matter of seconds.


499

Figure C.8: State of model at 10% extraction

0

5000

10000

15000

20000

25000

30000

35000

0 5 10 15 20 25

Number of balls extrac ted versus total % of extraction

% extrac ted

bal

ls e

xtra

cted

To ta l materia l

'Ore ' m ateria l

'W aste ' mate ria l

y p

Figure C.9: Histories of materials produced, PFC3D simulation


500

Figure C.10: Histories of total extraction (top line) and waste produced (lowest line), REBOP simulation

Figure C.11: Geometry of IDZ and surface profile at 145 days


501

C.5 CONCLUSIONS

PFC3D appears to offer some advantages over conventional (eg sand model) methods for understanding the flow of coarse granular material, produced in caving mines. These advantages relate to PFC3D’s ability to: • simulate coarse fragmentation, which is particularly important in caving mines in stronger

orebodies; • continuously monitor the evolution of key parameters (eg porosity) during the flow

process; and • visualise the evolution of the IDZ shape and size during the flow process. PFC3D also offers an advantage over sand models in the time necessary to set up and run a model. Nevertheless, there is significant computational time involved in obtaining PFC3D results. This limitation led to the development of REBOP. The main conclusion reached from the early development of REBOP, is that the application of simple incremental relations gives rise to evolving IDZ shapes that are similar to those seen in laboratory tests on the flow of broken rock and in PFC3D simulations. The relations govern the micro-mechanisms of collapse, erosion, free-surface slumping and interaction between IDZs, and the repeated application of these relations, together with the strict application of mass balance equations, leads to realistic flow patterns at the macro-level. The rules contain free parameters that can be adjusted to match REBOP “predictions” to calculated results from PFC3D simulations. A superimposed set of markers allows the paths of individual mass points to be tracked from their origin to the time at which they are extracted from a drawpoint. Thus, the composition of the extracted material is known at any time. REBOP allows for the tracking of any number of ore types, with any spatial distribution of grade for each ore type. A comparison between REBOP and PFC3D simulations shows good agreement, both qualitatively (IDZ shape and form of dilution history) and quantitatively (amount of dilution at a given stage). REBOP is in an early stage in its development. Further comparisons to known results are needed, using both laboratory and field data. Based on these comparisons, adjustments to the relations, or extra parameters, may be necessary. Assuming success in the comparisons, the code should be enhanced to the level at which it can function as a design tool, able to accept standard mine data and produce output in forms that are useful to planning and production engineers. When this is done, new comparisons should be made with full-scale field results, and code changes made if necessary. Lorig and Cundall (2000) have suggested the following specific changes to enhance REBOP.


502

• Modify the code such that it can be used routinely in design. For example, standard input

and output files should be used, standard sets of units should be included and standard nomenclature should be used (eg refer to “blocks” rather than “layers”).

• Add the ability to model the migration of fines. This extension will require PFC3D

simulations of material with wide size distributions in order to understand, and generalise, the mechanisms that will be added to REBOP.

• Add the ability to represent material fragmentation. Here again, PFC3D simulations will

be needed to guide REBOP development. As noted earlier, the studies reported in this Appendix represent only the initial stages of the continuing development of REBOP being undertaken at the time of writing as part of the International Caving Study Stage II. It is expected that REBOP will eventually provide the powerful numerical simulation capability required to significantly advance understanding of particle flow during draw in block and panel caving and to replace existing empirical mixing rules.

503

APPENDIX D

LIMITING EQUILIBRIUM ANALYSIS OF PROGRESSIVE HANGINGWALL CAVING

D.1 DERIVATION OF EQUATIONS

he assumptions made, and the variables involved in the limiting equilibrium analysis of the problem illustrated in Figure 9.16, are set out in Section 9.5.

Weight of wedge. The weight of the wedge of rock BCDNML in Figure 9.16 is

W2γ

= [ ( ) ( )( )

( ) ( )( )αΨΨ

Ψ+ΨΨ+α

αΨΨ

Ψ+ΨΨ+α

-

H -

-

H 21

p102

0p10

p202

02p022

sinsin

sinsin

sinsin

sinsin

( ) ( )αΨ

Ψα

αΨ

Ψα+

-

Z- -

Zp2

p222

1p

p121 sin

coscossin

coscos ] (D.1)

Base area of wedge. The area of unit thickness of the surface LM (Figure 9.16) on which failure takes place is

( )

( )αΨαα+Ψα

= -

Z- H A

p2

202

sincoscoscotsin

(D.2)

Thrust due to caved material. The thrust acting on the wedge BCDNML due to the caved material left in the crater is one of the most difficult parameters to estimate with confidence in this analysis. A simplified system of forces used to calculate T is shown in Figure 9.16.

T

Appendix D - Limiting Equilibrium Analysis of Progressive Hangingwall Caving

504

Resolving the forces Wc, Tc and T in the horizontal and vertical directions and applying the

equations of equilibrium of forces gives the solution

KH T 2cc2

1γ= (D.3)

where

( )( ) ( ) ( )w0wp1wp1

0p1

cotsincos

2cotcot

φΨφΨ+φΨ⎭⎬⎫

⎩⎨⎧

+Ψ+Ψ

= - - -

ΗS

K c (D.4)

The weight of caved material below the level of point B has been ignored in this calculation. Inclination of thrust to failure surface. It is assumed that the thrust T is transmitted through the wedge BCDNML to the failure surface without loss or deviation. Hence, the inclination of T to the normal to the failure surface LM is p1wp2 - Ψφ+Ψ=θ (D.5)

As shown in Figure 9.16, the angle θ may be either positive or negative. If θ is negative, the thrust T has a shear component that acts up the failure plane, and tends to stabilise rather than activate slip of the wedge. Water-pressure forces. The water-pressure force due to water in the tension crack is

2ww2

1 Z V γ= (D.6)

The water-pressure force U that acts normal to the failure surface is

AZ U ww21γ=

=( )

( ) ⎥⎥⎦

⎤

⎢⎢⎣

⎡

αΨαα+Ψα

γ -

Z- HZ w

2p

202w sin

coscoscotsin21

(D.7)


505

Conditions of limiting equilibrium. It is assumed that the shear strength of the rock mass in the direction of failure is given by the linear Coulomb criterion φ′σ+′=τ ' c tann (D.8)

The effective normal and shear stresses acting on the failure surface are

A

V - U - T W ' p22p sincoscos

nΨθ+Ψ

=σ (D.9)

and

A

V T W 2p2p cossinsin Ψ+θ+Ψ=τ (D.10)

The conditions for limiting equilibrium are found by substituting for 'nσ and τ into Equation

D.8, which, on rearrangement, gives

( ) ( ) ( )

0cos

sincossinsin 2p2p

A c-

U - V - T - W

=φ′′

φ′+φ′Ψ+φ′θ+φ′Ψ

(D.11)

Mining depth for new failure. Substitution for W, A, T, V and U from equations D.1, 2, 3, 6 and 7 into Equation D.9 and rearrangement gives a quadratic Equation for H2, the new mining depth at which failure will occur

( ) ( ) ( )

0

2p02p02

2

sin

sinsinsin

Ψ

φ′ΨΨ+ΨΨ+α⎟⎠

⎞⎜⎝

⎛′

γ

-

cH

2

( ) ( )

⎥⎦

⎤⎢⎣

⎡Ψ′

φΨ+αγΨ

φ′Ψ+α⎟⎠

⎞⎜⎝

⎛′

γ−

0

0ww

0

02

2

sin2sinsin

sincossin

2 c

Z -

cH


506

( ) ( )( ) ( )

( ) -

- -

cH

2 αΨΨ

αΨφ′ΨΨ+ΨΨ+α⎟⎠

⎞⎜⎝

⎛′

γ−

1p0

2p2p0p102

1

sinsin

sinsinsin

( ) ( )( )αΨ

αΨφ′ΨΨα⎟⎟⎠

⎞⎜⎜⎝

⎛′

γ+

- sin - sin - sin cos cos

cZ

p1

p2p2p12

1

( )φ′ΨΨα⎟⎠

⎞⎜⎝

⎛′

γ− -

cZ

p2 p2

22 sincoscos

( ) ( )αΨφ′θ⎟⎠

⎞⎜⎝

⎛′

γγγ

+ - - K cH

p2

2cc sinsin

⎟⎠

⎞⎜⎝

⎛ φ′′

γφ′

′αγ

+ sincZ

- cos c

Z

2cos2 ww2

( ) ( ) 0sincos 2p

2ww - -

cZ

p2 =αΨφ′Ψ⎟⎠

⎞⎜⎝

⎛′

γγγ

+ (D.12)

Equation D.12 gives a solution for the dimensionless group c/H ′γ 2 in the form

( ) 2/122 b a a cH

++=′

γ (D.13)

where

( ) ( )φ′ΨΨ+Ψ

⎟⎠

⎞⎜⎝

⎛′

φ′γφ′Ψ

= - sin sin

c2 sinZ

- cos sin a

2p02p

ww0

(D.14)

and


507

( ) ( )( ) ( )02p1p

2p01p2

1

sinsinsinsin

Ψ+ΨαΨ

αΨΨ+Ψ⎟⎠

⎞⎜⎝

⎛′

γ=

- -

cH

b

( )

( ) ( )02p0

0p2p12

1

sinsinsinsincoscosΨΨΨ+α

ΨαΨΨα⎟⎠

⎞⎜⎝

⎛′

γ−

- -

cZ

2

( ) ( )0p20

0p22

2

sinsinsincoscos

ΨΨΨ+α

ΨΨα⎟⎠

⎞⎜⎝

⎛′

γ+

-

cZ

2

( ) ( )

( ) ( ) ( )φ′ΨΨ+ΨΨ+α

ΨαΨφ′θ⎟⎠

⎞⎜⎝

⎛′

γγγ

− -

- - K

cH

c

p20p20

02

p22

c

sinsinsinsinsinsin

( ) ( ) ( )φ′ΨΨ+ΨΨ+α

Ψ⎟⎠

⎞⎜⎝

⎛′

φ′γφ′α

⎟⎠

⎞⎜⎝

⎛′

γ−

- sin sin sin

sin c2

sin Z - cos cos

cZ

22p02p0

02ww

2

( ) ( )

( ) ( ) ( )φ′ΨΨ+ΨΨ+α

ΨαΨφ′Ψ⎟⎠

⎞⎜⎝

⎛′

γγγ

− - sin sin sin

sin - sin - cos

cZ

p20p20

0 2

2p2p2

ww (D.15)

Critical tension crack depth. The left-hand side of Equation D.12 can be differentiated with respect to Z2 holding p2Ψ constant; the result is equated to zero to obtain the critical value of Z2 as

( )φ′ΨΨφ′

=′

γ -

cZ 2

p22p sincoscos

(D.16)

Critical failure plane inclination. By holding Z2 constant, differentiating Equation D.12 with respect to p2Ψ , putting 0/H 2p2 =Ψ∂∂ and rearranging, an expression for the critical failure plane angle may be obtained as


508

( ) ⎟

⎟

⎠

⎞

⎜⎜

⎝

⎛

++φ′=Ψ 2/122

-1 2p

Y X

X cos 21 (D.17)

where

( ) ( )

( )αΨΨ

αΨ+ΨΨ+α⎟⎠

⎞⎜⎝

⎛′

γ=

-

cΗ

Xp10

20p10

sinsin

cossinsin21

( )

02

002

2

sincossin

Ψ

ΨΨ+α⎟⎠

⎞⎜⎝

⎛′

γ−

cH

( ) ( )w1p

2cc

p1

2p1

21 - cosK

cH

- - sin

cos cos

cZ

φα+Ψ⎟⎠

⎞⎜⎝

⎛′

γγγ

αΨ

αΨ⎟⎠

⎞⎜⎝

⎛′

γ−

α⎟⎠

⎞⎜⎝

⎛′

γγγ

+ cZ

cos2

ww (D.18)

and

( ) ( ) ( )

( )αΨΨ

αΨ+ΨΨ+α⎟⎠

⎞⎜⎝

⎛′

γ+

ΨΨ+α

⎟⎠

⎞⎜⎝

⎛′

γ=

-

cH

cH

Yp10

201p0

21

0

02

2

sinsin

sinsinsinsin

sin

( ) α⎟⎠

⎞⎜⎝

⎛′

γαΨ

Ψαα⎟⎠

⎞⎜⎝

⎛′

γ−

cZ

- -

cZ

cossin

cossincos 22

p1

p12

1

( ) α⎟⎠

⎞⎜⎝

⎛′

γγγ

+φα+Ψ⎟⎠

⎞⎜⎝

⎛′

γγγ

− cZ

- K cH

sinsin2

www1p

2cc (D.19)

Angle of break. By simple trigonometric manipulation it is found that


509

( )( )

⎥⎦

⎤⎢⎣

⎡α

ΨΨ+α

Ψ

αΨ+Ψ=Ψ

Z-

H

- Z

2 cossin

sincos

sintantan

20

02p

2p22pb (D.20)

D.2 CALCULATION SEQUENCE

The following sequence of calculations allows angles of failure and angles of break to be estimated using the equations derived above. For dry conditions, U = V = 0. If the upper surface is horizontal, α = 0. Positive values of α are as shown in Figure 9.16. The equations listed also apply for negative values of α. The sequence of calculations is: (1) Calculate K from Equation D.4. (2) Let ( )φ′+Ψ=Ψ 1p2

1e2p

(3) Calculate cZ′

γ 2 from Equation D.16.

(4) Calculate a from Equation D.14. (5) Calculate b from Equation D.15.

(6) Calculate cH′

γ 2 from Equation D.13.

(7) Calculate X from Equation D.18. (8) Calculate Y from Equation D.19. (9) Calculate the estimated value of 2pΨ from Equation D.17.

(10) Compare 2pΨ from step 9 with e2pΨ from step 2. If e2p2p Ψ≠Ψ , substitute 2pΨ for

e2pΨ in steps 3, 4 and 5 and repeat the calculation cycle until the difference between successive values of 2pΨ is <0.1%.

(11) Calculate bΨ from Equation D.20.

INDEX

Advance or advanced undercut ........ 4, 18-

................................. 19.22-23,247,463

Air blast... ............ 12, 28, 31. 248. 346, 376,

................... 378. 380, 393. 396-399, 463

Airgap ....................... 9.11-12, 22.332-333.

............................. 337, 398, 399, 463

Andina Mine, Chile

Cavability data .......................... 129. 136

Production costs ................................... 8

Type of operation ............................ 15

Undercut support and

reinforcement ...................... 224-225

Angle of break or angle of

subsidence ....... 347. 352. 367-373. 463

Angle of draw ................................. 347. 463

Arching .............................. 9.177.396.463

Areal (window) mapping ........................ 52

Athens Mine, Michigan, USA

Plug subsidence ...................... 351, 362

Austro-American Magnetite

Company Mine, Austria .......... 248-249

Back analysis ................................. 328.330

Bedding .................................................... .41

Bell Mine, Canada

Drawpoint support and

reinforcement ............... 279. 281-283

Fan undercut ............................. 205-206

Undercut height... ............... 204.235-236

Undercutting rate ............................... 213

Bingham Canyon, Utah, USA

Introduction of block caving ................. 16

Ore crushing and transportation ........ 254

Block Cave Fragmentation (BCF) ............. .

................................... 157, 172-181. 190

Block caving

General description ............................. 3-8

History .......................................... 12-15

Production costs .................................. 8

Brier Hill Mine, Michigan, USA

Subsidence to surface .................... 362

Brow ............... 5. 15. 23, 248, 266-268, 464

Bulking or swelling ................. 12, 302.464

Bulking factor ................... 12.302-303,464

Bull nose ........................................ 275, 464

Bultfontein Mine, South Africa


Camel back ..................................... 275. 464

Cavability ........ 11, 16.27-30.126-128,130.

......... 136-138.141.154,409-410,464

Cavability assessment

Continuum modelling ................. 139-147

Definition of cavability .................. 28. 464

Laubscher's caving chart ............ 126-130

Mathews' stability graph ............. 130-138

PFC3D modelling ....................... 147-155

Probabilistic approaches ............ 134-137

Cave back or crown ............ 7-12. 331-333.

............................................... 336.464

Cave back monitoring .................... 331-336

Cave inducement ............. 11. 168. 396. 464

Cave initiation ............... 3. 27, 31,126.191,

........................ 198.226.332-335.464

Cave propagation ..................... 12. 28. 192.

................................ 198.204.344,464

CaveRisk .................. 31, 375, 400, 406-422

Caving mechanics ................................ 8-12

Caving methods of mining

Classification .......................................... 7

General description ............................ 3-8

Caving rate ............................... 12, 201,464

Cavity monitoring .................................. 336

Century Mine, Australia .......................... 35

Chimney caving or chimneying ............ 31,

......................... 348-357, 361-362. 464

Cleavage or schistosity .......................... 40

Climax Mine, Colorado, USA .................. 13

509

Continuous subsidence ................ 346, 465

Core logging ................................ , ..... .43-48

Core orientation .................................. 43-46

Crown pillar ............... 7, 350, 359, 361-363,

................................................ 396, 465

Dilution ............. 27,293,297-298,304-305,

................................................ 308,465

Digital image processing systems ......... ..

................................................. 162-169

Discontinuities

Analysis and presentation of data ......... ..

............................ , ........................... 55-77

Areal frequency .................................... 61

Areai (window) mapping ...................... 52

Data collection ............................... .42-55

Definition ................................. 36-37, 465

Description ...................................... 41-42

Error and uncertainty in analysis .... 56-58

Frequency I spacing analysis ......... 60-61

JK Jointstats system ...................... 67-77

Lineal (scanline) mapping .............. 50-52

Linear fracture frequency ..................... 61

Orientation analysis ........................ 58-59

Persistence (size) analysis ............. 64-66

Set definition .................................. 71-73

Spacing ........................... 62-64, 104-105

Spot mapping ................................. 49-50

Termination ....................... 92-95,99-100

Volumetric density ............................. 60

Discontinuous subsidence .......... 346-348.

........................................ 353, 364. 465

Displacement monitoring ...... 325-326, 329

Draw ......................... 2, 9,12-13, 28,31,

......................................... 293-321; 465

Draw column ................. 294, 297, 306, 309,

....................... 313-314,316,318,465

Draw cone ....................... 312, 315-316, 465

Draw contro!.. ......... 9, 12.28, 31. 293-321;

............. , .................................. 397, 465

Draw rate ......... 300, 304-305, 309-310. 465

Draw zone ................................ 299-300. 465

Index

510

Drawbell ............ .4, 19, 22, 25, 28, 31,192-

....................... 196,200,204,209,215,

............................... 226, 241-244, 465

Drawpoint... ...... 4-6, 18-23,26,28,31,245,

......................... 247-273, 275-292, 465

Drawpoint drift ............ 19, 31,291-292.465

Drawpoint spacing ..... 23,26,28,247,255,

................. 258-262,265-266,283,466

Ellipsoid of motion draw or extraction ....

................ .255, 257-250, 265, 296,466

El Teniente Mine, Chile

Advance undercut ............................. 194

Angle of break ........................... 370, 372

Cavability data ........................... 129,136

Caving rates ...................................... 12

Drawbell shapes ................................ 270

Drawpoint spacing ............................. 266

El Teniente layout... .................... 252-253

Extraction Isvellayouts ............ 215, 262,

................................................ 264-265

Extraction level support

and reinforcement... ............. 286~288

Faults ................................................... 40

General description ........................ 16-19

Macro-trench ......................................... 5

Major collapses .................. 377, 379-380

Pillar damage ............................ 342-343

Production costs .......... " ........................ 8

Rock bursts ................................. 385-388

Stresses induced on extraction

leveL ............................................. 216

Subsidence crater ...................... 370-374

Undercut height... ............... 203-204, 234

Veins .................................................... 39

Zone of influence of cave ................ 347 .

................................................ 370,374

Energy Release Rate (ERR) .... ......... .. 384

Esmeralda Section, El Teniente Mine, Chile

Cavability ......................................... 137

Drawbell design ................................. 243

Flat undercut .............................. 206-208

Fragmentation ............................ 168-169

Pre-undercut 45° rule ......................... 193

Undercut drift support and

reinforcement ............................... 225


Excess Shear Stress (ESS) .................. 384

Extraction or production level ....... .4, 6, 9,

......................... 15, 18-19,22-23,26,28,31,

................................. 245-292, 294, 299, 466

Finsch Mine, South Africa

Blasting ............................................. 233

Fisher constant ........................... 59, 89-91

Fisher distribution ................. 59, 71, 88-91

Flow ellipsoid ......................... 225, 257, 466

Fragmentation ......................... 4, 23, 27-31,

................................................. 156-190,466

Fragmentation assessment .......... 156-190

BCF ............................................ 172-181

Digital Image processing ........... 162-169

Factors influencing

fragmentation ....................... 157-158

Fragmentation

measurement ....................... 159-169

In situ fragmentation ................. 169-172,

.............................................. 186-187

JKFrag ....................................... 181-190

Primary fragmentation .............. 174-175,

...................................... 178, 187-190

Secondary fragmentation .......... 175-178,

...................................................... 190

Freeport Mine, Indonesia

Air blast from ore passes ................... 397

Production costs .................................... 8

Type of operation ................................. 15

Wet muck .......................................... 389

Front caving .................................. 4, 5, 466

Gath's Mine, Zimbabwe

Progressive hangingwall caving ....... 364,

................................................... 368-369

Index

511

Geological or structural domains

.............................................................. 66-67

Geological Strength Index (GSI) .. 114-116,

............................................ 119-120, 123

Geophysical techniques .................... 53-55

Geotechnical core logging ................ 43-49

Geotechnical monitoring

Cavity monitoring ............................... 336

Convergence measurements

............................................ 326, 338-339

Displacement measurements

............................................ 325-326, 329

Extraction level monitoring ......... 337-342

Geotechnical monitoring

systems ................................. 324-330

Initiation and development of caving

.................................................... 331-336

Microseismic methods ................ 333-336

Pillar damage .............................. 342-343

Purposes ..................................... 322-323

Reinforcement loads ................... 340-342

Stress changes ........................... 339-340

Subsidence and ground movement

.................................... 342-345, 352-356

Time Domain Refiectometry ....... 332-333

Grangesburg Mine, Sweden

Progressive hangingwall caving ....... 364

Gravity draw .......................... .4, 13-15,466

Gravity flow ............................. 17, 255, 257,

.................................. 264-265,296-297,466

Grizzly ................................... .4, 15, 20, 245,

......................................... 253, 281,283,466

Ground-support interaction ............. 72-73,

.................................................. 283-285, 292

Hangingwall caving ................. 31, 348-349,

.................................. 364-369,466,501-508

Hangup .............................. 23, 28, 178,247,

......................................... 281, 381, 393, 466

Haulage level ........................ .4, 13,26,245,

................................................. 253, 291, 467

Minor apex .................... 249, 253, 259, 266,

......................................... 277-278, 283, 468

Mont Porphyre Project, Canada

Core orientation .................................. .45

Mud rush ............................... 248, 250, 376,

................................................. 387-399, 468

Narrow flat undercut ..... 196-197, 200-203,

......................... 205, 207, 234-236, 241, 468

Narrow inclined undercut ............ 195-196,

.......................... 200-205, 210-212, 234, 468

Newcrest Ridgeway Project

Core photographs ........................... 73-74

Noranda Technology Centre, Canada

Validation of digital imaging

systems ........................................ 165

Northparkes E26 Mine, Australia

Air blast .............................................. 396

Cavability ............................ 129, 136-137

Cave monitoring with TDR ................. 333

Caving mechanics ............................. 9-11

Caving rate ........................................... 12

Drawbell blasting ................................ 242

Drawpoint fragmentation .................... 168

Drawpoint spacing, Lift 2 ................... 266

Explosive energy distribution ..... 239-240

High undercut... .......................... 202-204

Hydraulic fracturing ............................ 168

Microseismic monitoring ............ 334-346

Narrow inclined undercut, Lift 2 ......... 21 0




Undercut advance direction ............... 197

Undercut drift support and

reinforcement ............................... 225

Undercut powder factors .................... 232


Optical imaging ... .................................... .48

Ore crushing and transportation. 253-254

Index

513

Palabora Mine, South Africa

Development of BCF ......................... 172

Drawpoint spacing ............................. 266

Extraction levellayout... .............. 250-251


Narrow inclined undercut ................... 210



Undercut level support and

reinforcement ........................ 225-226

Undercut sequence ........................... 197

Panel caving ................ 4-5, 7, 16-20,22-27,

........................ 245, 247-248, 250, 254, 259,

.......... 260-262, 266, 272-275, 286, 289, 468

Particle Flow Code (PFC) ............. 147-155,

.................................. 295-296, 468, 484-500

Pewabic Mine, Michigan, USA

First form of block caving ..................... 13

Phi lex Mine, The Phillipines



Plug subsidence .................... 348, 350-351,

............................................ 356-361. 468

Post- or conventional undercut ............ 18,

....................... 22-23, 192-193, 195,200,

.................................... 217-218, 224, 468

Premier Mine, South Africa

Caving rates ......................................... 12

Drawbell blasting ............................... 242

General description ........................ 20-23

Fragmentation .................................... 160

Ground reaction curves .............. 283-285

Introduction of mechanised

panel caving ................................... 15

PC-BC for draw control.. .................... 318

Post-undercut .................................... 193


Undercutting strategy ........................ 213

Pre-undercut .............. 19-20, 192-193, 195,

......................................... 208, 215, 225, 468

Primary fragmentation .......... 156-158, 160,

................... 172-176, 178-182, 187-190,468

Principles of good support and

reinforcement practice ............. 272-274

Production or extraction drift ...... 310,468

Questa Mine, New Mexico, USA

Haulage drift monitoring ............ 338-339

Rapid muck pile compaction

mechanism ................................ 389-390

REBOP (Rapid Emulator Based

on PFC3D) ................. 296, 468, 484-500

Reduced drainage mechanism ..... 389-390

Reinforcement ...................... 191, 193, 195,

................................. 224-226,245, 247-248,

........................... 270-276, 279-282, 292 469

Rio Blanco Mine, Chile

Effect of clayey fines .......................... 393

Structurally controlled collapse ............... .

.................................................... 379-380

Zone of influence of cave .................. 370

Risk .................................... 27, 29, 293, 275,

........................................... 376, 393, 469

Risk analysis ............... 29-30, 400-422, 469

Risk assessment ............ 29, 393, 402, 469

Roadway .................................. 254-256, 469

Rock burst.. ........... 200, 248, 271, 286, 338,

........................... 376, 382-388, 392, 469

Rock mass characterisation .......... 32-125,

................................................... 292, 469

Rock mass classification ................ 36. 46.

.................................................. 100-116.469

Rock mass deformation modulus

.......................................................... 122-123

Rock Mass Rating (RMR) .............. 100-105.

.......................................... 108-109, 118,469

Rock Quality DeSignation (RQD) ..... 46-47.

.......................................................... 101-104

Salvador Mine, Chile

Air blast ....................................... 396-398

Index

514

Cavability data ................................... 136

Extraction level support and

reinforcemenL ...................... 291-292


Type of operation ................................ 15

San Manuel Mine, Arizona, USA

Continuous trough or trench

layout .................................... 248-249



Surface subsidence monitoring

............................................ 345, 352-354

Secondary breaking ...... 117, 159,247,469

Secondary fragmentation ............ 156, 158,

.................................... 175-178.190.469

Shabanie Mine, Zimbabwe

Air blast .............................................. 396

Continuous trough or trench

layout ........................................... 248

Direction of undercut advance .......... 199

Front caving ........................................... 4

Herringbone layout.. .......................... 250


Shape factor (or hydraulic radius)

.................................. 127.130,133-138.469

Simulation of rock mass geometry ......... ..

............................................................ 77-100

Forward modelling techniques ....... 84-85

Fractal based models .......................... 82

Geostatistical models ..................... 83-84

Hierarchical models ........... 82-83, 92-100

Random coplanar polygon

models ...................................... 80-82

Random disc (Poisson location)

models ........................... 79-80. 85-92

Slusher ................. .4, 15. 205. 245-246, 469

Stacking ......... .204, 208-210, 247, 268, 470

Stereographic projection .......... .46, 58-60,

. ............................................................. 71-72

Stress caving ............................ 9, 158,470

Subsidence caving ............. 11-12, 158,470

Support ................. ...... 23. 33. 43. 100. 108 •

....... 191-195.213.224-227.279-292 •

...................... 322. 325. 338-339. 470

Surface subsidence .............. 324.342-345 •

...................................... 346-374.470

Time Domain Reflectometry

(TOR) ......................... 332-333. 345. 355

Tongkuangyu Mine, China

Cave monitoring ................................. 332


Undercut ................... 23. 126-127. 136-138 •

............................ 140-145.148-151.470

Undercut level.. ...... 192-195. 204-205. 211 •

............... 213.215-216.221-226.470

Undercutting rate .................. 200-201.470

Underground mining methods .................... 1-3

Unstable slip .................................. 382-384

Urad Mine, Colorado, USA

Air blast .............................................. 396

Subsidence crater .............................. 346

Veins ........................................... 38-39. 113

Void diffusion ........................ 296-300. 304 •

...................................... 307-308.470

Wet muck flow ....................... 388-389. 392-

...................................... 393. 397.470

Zone of influence .................. 347. 369-370 •

...................................... 372.374.470

Index

515

block caving geomechanics s

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