DEVELOPMENT, CHARACTERIZATION AND APPLICATION OF A REACTIVE BULKING AGENT FOR WALL CONTROL

Transcription

1 DEVELOPMENT, CHARACTERIZATION AND APPLICATION OF A REACTIVE BULKING AGENT FOR WALL CONTROL By: Guillermo C. O. Silva A thesis submitted to the Department of Mining Engineering in conformity with the requirements for the degree of Doctor of Philosophy Queen s University Kingston, Ontario, Canada November, 2007 Copyright Guillermo C.O. Silva, 2007

2 ABSTRACT This research thesis is focussed on the development of a novel low density explosive composition whose main application is wall control in open pit mining. The product has, however, the potential to be used in a variety of applications and rock conditions where customization of the explosive s energy output is required. Experimental observations on the novel low density explosive showed that the product is capable of initiating and sustaining stable detonations at densities as low as 0.10 g/cm 3. Given the extreme low densities at which the novel product maintains its detonating characteristics, it will be appropriate to treat it as a reactive bulking agent, hence its name: Low Density Reactive Agent or LDRA for short. When mixed with standard ANFO prills, the reactive nature of the LDRA ensures a detonable mixture regardless of the final dilution sought or the degree of segregation eventually produced during mixing and loading. If operational constraints are such that a lower energy is required, the LDRA can then be used on a stand alone basis, without mixing it with other explosive compositions, such as ANFO or emulsions. The detonation characteristics of the LDRA at a target density of 0.15 g/cm 3 were evaluated, with particular effort placed on measuring the detonation and explosion pressures, parameters having the greatest influence on damage. The effects of diameter, confinement and primer on LDRA performance were evaluated through velocity of detonation (VOD) measurements. In addition, VOD experiments were conducted in the ii

3 LDRA to evaluate the stability of propagation in longer columns, the behaviour in a decking configuration and the ability to initiate and be initiated by a column of ANFO. The low pressure regimes characterizing the LDRA provided the opportunity to investigate the full pressure history of the detonation gases by designing experiments of a non destructive nature that allowed the repetition of tests under different loading scenarios. Following the characterization stage, the opportunity to test the LDRA as a damage control tool under a true operational scenario arose at the Chuquicamata Mine, in northern Chile. The project provided important input as to the feasibility of manufacturing the LDRA at a semi industrial scale and to evaluate the performance of the product in the large diameter blastholes used at the mine. iii

4 ACKNOWLEDGEMENTS Long journey this thesis; journey that would not have been possible if it wasn t for the support that represented my wife Mariana. Her sole presence and example was the driving force behind my undertaking. To her goes my love, my recognition and the sharing of this effort; it belongs to her as much as it does to me. Following in the list is Takis Katsabanis, an old time fellow co worker, then a business partner, then my supervisor, but overall a good friend, which at the end of the game is what prevails. He was instrumental in directing me in the right path and patient with the many times I lost the compass during this expedition, providing guidance as to the what, why, when and where of this thesis. To him goes my gratitude and friendship. Many people in the Department, in one way or another, accompany me during this endeavour. Special thanks goes to Dr. Yen for his kindness and willingness to help. Charlie Pelley, with whom I shared issues that go far and beyond this research and whom I am glad to have known; Jamie Archibald, for his willingness to help and participate and last but not least, the support provided by the Departmental Staff, starting from Dear old Bev to Peter, Maritza, Ray, Wanda, Michelle, Jenny and the countless list of people that passed by during my long stay at Queen s. They really made my miserable days enjoyable. iv

5 TABLE OF CONTENTS Abstract... ii Acknowledgments... iv Table of Contents... v List of Tables... x List of Figures... xii Chapter 1. Thesis Objective and Scope of Work Introduction Research Objectives Thesis Outline 5 Chapter 2. Review of Low Density Explosive Issues Introduction Review of Pressure Concepts Detonation Pressure Explosion Pressure Borehole Pressure Summary of Pressure Concepts Role of Density on Damage Control General Considerations Related to Low Density Explosives Russian Experiences using EPS based Explosives The Shock Behavior of Expanded Polystyrene (EPS) Detonation Behavior of EPS and AN Mixtures Chapter Conclusion 34 Chapter 3. Rock Breakage and Wall Control Topics Introduction 35 v

6 3.2 Mechanisms of Rock Breakage Wall Control Techniques Empirical Approaches Presplitting Suggested Presplitting Mechanism Buffer Blasting Cushion or Trim Blasting Smooth Blasting Air Decking Particle Velocity Damage Criteria Design Methods Swedish Approach CSM Approach Innovative Contributions to Wall Control Chapter Conclusions 56 Chapter 4. Development of the Low Density Reactive Agent Introduction Selection of Bulking Agents Expanded Polystyrene (EPS) Beads Developmental Stages of the LDRA Laboratory Tests with Hot Solutions of Ammonium Nitrate Experiments Using Mineral Oil as Surfactant Selection of Ammonium Nitrate Fines LDRA: Final Mixture Oxygen Balance Water Resistance Potential Improvements for Field Applications Chapter Summary 84 Chapter 5. Characterization of the LDRA Introduction Velocity of Detonation Experiments VOD Measuring Technique VOD Diameter Relationship 90 vi

7 5.2.3 Effect of Confinement Effect of Density Initiation Behavior of the LDRA Priming Recommendations for the LDRA Propagation Stability of the LDRA in Longer Charges Behavior of LDRA in Decking Configurations Mixtures of LDRA and ANFO Mixtures of LDRA with Ammonium Nitrate (AN) Prills Segregation Tests on LDRA and LDRA/ANFO Mixes Segregation Tests on the LDRA Segregation Tests on LDRA/ANFO Mixes Pressure Measurements on the LDRA Modified Aquarium Technique Pressure Sensors PVDF Film Sensors Assembly and Recording Instrumentation Experimental Analysis and Results Carbon Composition Resistors Assembly and Calibration Equation Selection Detonation Pressure Results Novel Explosion Pressure Measurement Development of a Calibration Equation Validation of Experimental Approach Explosion Pressure Experiments with LDRA Chapter Summary 159 Chapter 6. Field Application of the LDRA Introduction Typical Blast Design Applied in Chuquicamata Proposed Objectives of the Field Project Large Scale Manufacturing of the LDRA Manufacturing Process of LDRA at Enaex Facilities Quality Control at Manufacturing Plant Manufacturing of LDRA/ANFO Mixtures Field Experiments at Chuquicamata Mine 177 vii

8 6.5.1 Description of Field Experiments Experiments involving the LDRA mixture Baseline Experiments Velocity of Detonation Experiments Explosive Loading and Measurement of VOD Analysis of VOD Experimental Results Gas Pressure Experiments Gas Sensor Assembly and Field Implementation Gas Pressure Experimental Results Conclusions of Gas Pressure Results Near Field Vibration Experiments Field Implementation of Vibration Experiments Vibration Analysis of LDRA Trials Vibration Analysis of Baseline Experiments Comparison of Vibration Experiments Post Blast Inspection of LDRA Field Tests Suggested Design Alternatives Conceptual Design of a Manufacturing Plant for the LDRA Pending Research The Problem of Flammable Gases Chapter Summary 250 Chapter 7. Research Contributions Summary of Main Thesis Contributions Patent Information on the LDRA 256 References Appendix 2.1: Survey of Explosive Technologies Applied to Wall Control Appendix 4.1: Information on Bulking Agents Appendix 4.2: Recommendations on LDRA Final Mix Appendix 4.3: Potential Field Improvements Appendix 5.1: Velocity of Detonation Records viii

9 Appendix 5.2: Detonation Pressure Techniques Appendix 5.3: PVDF Pressure Records Appendix 5.4: CCR Pressure Records (gel and water) Appendix 5.5: CCR Pressure Records (cardboard tubes) Appendix 5.6: Reflected Hugoniot Technique Appendix 6.1: Quality Control of LDRA Appendix 6.2: Geotechnical Data of Testing Areas Appendix 6.3: Blast and Vibration Data of LDRA Experiments Appendix 6.4: Blast and Vibration Data of Baseline Experiments Appendix 6.5: LDRA Processing Plant ix

10 LIST OF TABLES Table 3.1: Peak particle velocity damage thresholds Table 4.1: Laboratory results showing the effect of oil content on adhesion of ammonium nitrate fines and mixture density for different EPS bead diameters Table 4.2: Experiments conducted using a preliminary composition of a LDE mixture consisting on oil wetted EPS beads and ammonium nitrate powder Table 4.3: Adhesion of ANMO fines for different tackifying and wetting oil contents Table 4.4: Sieve analysis of ammonium nitrate fines as supplied by Nitrochem Table 5.1: Initiation of LDRA at 0.15 g/cm 3 with detonating cord Table 5.2: VOD for LDRA at 0.15 g/cm 3 in 3.20 meter long steel pipes Table 5.3: Evaluation of LDRA in a decking configuration Table 5.4: VOD of LDRA/ANFO and EPS/ANFO at different diameters and mix ratios (Sudweeks, 2000) Table 5.5: VOD of LDRA/AN mixes at various densities Table 5.6: Characteristics of PVDF sensors used during experimentation Table 5.7: Detonation pressure results obtained from PVDF gages on LDRA at 0.15 g/cm Table 5.8: Average detonation pressures of LDRA under different loading conditions Table 6.1: Blasthole loading features for LDRA Test Table 6.2: Blast design data for LDRA Test Table 6.3: Blast design data for LDRA Test 5 using LDRA at 0.20 g/cm Table 6.4: Blast design data for LDRA Test x

11 Table 6.5: Summary of VOD measurements conducted during field trials at Chuquicamata Table 6.6: PPVcritical for the mine expansions where LDRA field experiments were conducted Table 6.7: Maximum PPV recorded during field trials using the LDRA and mixes with ANFO Table 6.8: Blast design data for LDRA Test Table 6.9: Data used to generate Holmberg and Persson predictive model for LDRA Test Table 6.10: Relevant data from baseline trials for typical blasting practices used at the mine Table 6.11: Berm widths as an indication of damage produced by the LDRA and Enaline sections corresponding to presplit LDRA Test Table 6.12: Full Control program results for LDRA Test Table 6.13: Full Control program results for LDRA Test xi

12 LIST OF FIGURES Figure 2.1: EPS at g/cm 3 subjected to shock from 20 g of ammonite through a steel plate (Nifad ev et al, 1992/a) Figure 2.2: Before and after effect of shock on EPS at a density of g/cm 3. Shock was provided by 20 g of ammonite through a steel plate barrier (Nifad ev et al, 1992/a) Figure 2.3: Partial gasification of EPS granules at a density of 0.05 to 0.10 g/cm 3, subjected to explosive shock induced by 100 g of Ammonite without a steel plate barrier (Nifad ev et al, 1992/a) Figure 2.4: Bulging and banding effect produced by low density mixtures of ammonium nitrate and expanded polystyrene (Nifad ev et al, 1992/b) Figure 3.1: Notation used for the calculation of particle velocity using Holmberg s approach Figure 4.1: Cumulative weight retained for ammonium nitrate fines provided by Nitrochem Figure 5.1: VOD record using a brass sensor on LDRA at 0.12 g/cm 3 in 75 mm steel pipe Figure 5.2: Aluminum and brass tubes used for VOD sensors Figure 5.3: VOD trace with an aluminum sensor. LDRA at 0.12 g/cm 3 in 75 mm steel pipe Figure 5.4: VOD Diameter relationship. LDRA at 0.15 g/cm 3 under steel confinement Figure 5.5: Average VOD as a function of diameter for LDRA at 0.15 g/cm Figure 5.6: Ideal VOD inferred from experimental data for LDRA at 0.15/gcc under steel confinement Figure 5.7: Effect of confinement on stable VOD regime and critical diameter Figure 5.8: VOD of LDRA at 0.27 g/cm 3 confined in heavy duty cardboard tubes Figure 5.9: Density vs. VOD for 50 mm and 75 mm confined LDRA charges xii

13 Figure 5.10: Before and after events of VOD experiments conducted on 3.20 m long, 50 mm and 75 mm diameter steel tubes loaded with LDRA at 0.15 g/cm Figure 5.11: Composite traces for the two aluminum probes used to measure VOD in the 3.20 meter long, 75 mm steel tube loaded with LDRA at 0.15 g/cm Figure 5.12: VOD record of a 75 mm diameter steel tube loaded with similar column lengths of LDRA at 0.15 g/cm 3 and ANFO at 0.85 g/cm Figure 5.13: Test configuration used for the evaluation of the LDRA in decking applications Figure 5.14: Comparison of VOD traces generated by LDRA/ANFO and EPS/ANFO mixes Figure 5.15: VOD tests on mixes of ANFO/LDRA and ANFO/EPS at different densities Figure 5.16: Quality of VOD traces for two experiments conducted on AN/LDRA mixtures Figure 5.17: Bottom and side view of a 150 mm plexiglas tube loaded with LDRA after being subjected to a 24 hour vibration test Figure 5.18: Effect of tapping on LDRA density Figure 5.19: Test setup to measure density gradients Figure 5.20: Density readings of LDRA taken from four sections of a 75 mm tube Figure 5.21: Density readings of LDRA taken from four sections of a 125 mm tube Figure 5.22: Density gradient in mixtures of LDRA/ANFO Figure 5.23: Density gradient in mixtures of EPS/ANFO Figure 5.24: Photograph and schematic arrangement of the modified aquarium test Figure 5.25: Record obtained from the modified aquarium method showing a constant VOD on the explosive section and a decaying shock velocity in water Figure 5.26: Modified aquarium test conducted on LDRA at 0.15 g/cm xiii

14 Figure 5.27: Schematic of the experimental arrangement used to infer detonation pressure by measuring explosive VOD and shock velocity in water via two aluminum probes Figure 5.28: Records produced by aluminum probes placed in the LDRA and in water in order to measure VOD and shock velocity respectively Figure 5.29: PVDF film gage manufactured by Ktech Corp Figure 5.30: Dynasen s RC hardware integrator for charge mode recording with PVDF gages Figure 5.31: Charge mode recording with the RC circuit connected at the PVDF leads Figure 5.32: PVDF gage assembly set for charge mode recording Figure 5.33: PlotData screen plot showing results on LDRA test using current mode configuration. Both the current trace and the corresponding pressure profile are shown Figure 5.34: Comparison of charge and current mode records for LDRA in 50 mm steel pipes Figure 5.35: Setup used to protect the carbon composition resistors (CCR) Figure 5.36: Pressure generated by a CCR placed in water and shocked with LDRA at 0.15 g/cm 3. Detonation pressure of the LDRA is then determined from this record Figure 5.37: Summary of detonation pressure experiments conducted in 51 mm diameter steel tubes using carbon composition resistors under different configurations Figure 5.38: Threaded thick wall steel tube and steel plug used during the explosion pressure experiments. Cross section of assembly showing pressure chamber and resistor placement Figure 5.39: Components of the steel plug to house the resistor. A stainless steel thermocouple cable and a duplex wire cable connected to the CCR embedded in epoxy are also shown Figure 5.40: Aquarium arrangement used by Katsabanis (1997) Figure 5.41: Aquarium test used for developing calibration equation for low pressure regimes xiv

15 Figure 5.42: Experimental setup used to measure explosion pressure of detonating cord Figure 5.43: Borehole pressures generated by detonating cord of different strengths, evaluated with the experimental arrangement and calibration equation developed at Queen s Figure 5.44: Validation of experimental setup used to measure borehole pressures with carbon composition resistors: experimental results versus output estimated by Canmet s equation Figure 5.45: Composite of explosion pressure records generated by a fully coupled LDRA Figure 5.46: Explosion pressures generated by the LDRA as a function of charge length Figure 6.1: Overview of Chuquicamata Open Pit Mine looking north Figure 6.2: Plan view of a typical blast pattern used in Chuquicamata Mine Figure 6.3: NCN mixing facilities within Enaex s plant Figure 6.4: Bin container for the AN prills and the grinder used to produce AN fines Figure 6.5: Vibrating screen used to classify the AN fines Figure 6.6: Stainless steel rotating mixer used during manufacturing of the LDRA Figure 6.7: LDRA 10 kg bags ready for storage Figure 6.8: Quality control test conducted on the LDRA at Enaex s testing grounds Figure 6.9: Blast and monitoring hole coordinates corresponding to LDRA Test Figure 6.10: Blast and monitoring holes coordinates corresponding to LDRA Test Figure 6.11: Diagram of blast and monitoring holes location corresponding to LDRA Test Figure 6.12: Presplit and monitoring holes coordinates corresponding to LDRA Test xv

16 Figure 6.13: Presplit and recording holes corresponding coordinates to LDRA Test Figure 6.14: Mine coordinates corresponding to LDRA Test Figure 6.15: Manual loading of LDRA into test holes Figure 6.16: Field control of LDRA density Figure 6.17: Bucket loading of LDRA (blue) and ANFO (grey) for a 50/50 volume mix Figure 6.18: VOD records of two presplit holes loaded with LDRA at 0.20 g/cm Figure 6.19: VOD trace of a LDRA/ANFO at 0.37 g/cm 3 in 350 mm diameter hole Figure 6.20: VOD Density data for the various diameters tested at Chuquicamata Figure 6.21: VOD Diameter relationship for the different LDRA mixtures Figure 6.22: VOD trace for a LDRA/ANFO mixture at 0.49 g/cm Figure 6.23: Negative air pressure trace, indicative of damage, at 5 m from contour blast (Ouchterlony, Nie et al, 1996) Figure 6.24: Positive pressure trace, indicative of gas pressurization, at 2.8 m behind a presplit (Ouchterlony, Nie et al, 1996) Figure 6.25: Sensor assembly used to monitor gas pressure Figure 6.26: Calibration curves of the gas pressure sensors used during field experiments Figure 6.27: Experimental setup used for measuring gas pressure behind the last row Figure 6.28: Gas pressure sensor assembly in a 311 mm hole Figure 6.29: Gas pressure trace from 165 mm blastholes. LDRA Test 1 (MicroTrap record) Figure 6.30: Timing of VOD traces corresponding to LDRA Test 1 (MicroTrap record) Figure 6.31: Gas pressure trace 6.5 m behind a 350 mm hole loaded with LDRA at 0.20 g/cm 3 corresponding to LDRA Test xvi

17 Figure 6.32: Pressure at 3.1 m behind presplit section loaded with LDRA at 0.20 g/cm Figure 6.33: Pressure at 3.1 m behind presplit section loaded with Enaline emulsion cartridges Figure 6.34: Geometric parameters used in Holmberg and Persson s equation Figure 6.35: Vibration sensor embedded in epoxy and setup of PVC tube in the blasthole Figure 6.36: Diagram showing field implementation of vibration measurements Figure 6.37: Mine coordinates showing the relative location of the six blastholes and two geophones used during the first experiment conducted with LDRA Figure 6.38: Vibration records showing the three particle velocity components of the PPV corresponding to LDRA Test 1. Sensor located behind 165 mm diameter blastholes Figure 6.39: H&P predictive model for LDRA Test Figure 6.40: H&P predictive model for LDRA Test 6. Buffer row, 350 mm. LDRA at 0.20 g/cm Figure 6.41: H & P predictive model for Baseline Test B3. Simulated Buffer row, 350 mm loaded with heavy ANFO Figure 6.42: Enaline emulsion cartridge used for pre splitting Figure 6.43: Presplit of Test 4 indicating Enaline (back) and LDRA (forefront) sections Figure 6.44: Surface damage behind LDRA blastholes Figure 6.45: Transition area from LDRA to Blendex Figure 6.46: Trench behind Blendex 945 blastholes Figure 6.47: PPV isolines generated by an emulsion at 1.20 g/cm 3 in a 311 mm blasthole Figure 6.48: PPV isolines generated by LDRA at 0.20 g/cm 3 in a 311 mm blasthole Figure 6.49: PPV isolines generated by LDRA at 0.20 g/cm 3 in a 165 mm blasthole xvii

18 Figure 6.50: PPV isolines for the present and alternative design for Expansions 40 E and 41 E Figure 6.51: PPV isolines for the present (top) and alternate (bottom) design for Expansion 47 West Figure 6.52: Timing sequence of LDRA Test 6 showing potential areas of vibratory synergies Figure 6.53: Frame sequence at 2, 10 and 28 ms after initiation. LDRA at 0.15 g/cm Figure 6.54: Frame sequence at 2, 16 and 26 ms after initiation. LDRA and perlite mixed at 50/50 volume. Mixture density of 0.20 g/cm Figure 6.55: Frame sequence at 2, 4 and 6 ms after initiation. Perlite ammonium nitrate based mixture (PANFO) at a density of 0.28 g/cm Figure 6.56: Frame sequence at 2, 4 and 6 ms after initiation. LDRA at 0.15 g/cm % H 10 flame suppressor powder Figure 6.57: Frame sequence at 2, 4 and 6 ms after initiation. LDRA at 0.15 g/cm % H 10 flame suppressor powder xviii

19 Chapter 1. Thesis Objective and Scope of Work 1.1 Introduction During the last decades, there has been a committed effort by explosives manufacturers, researchers and users to develop low density explosive mixtures and to experiment with loading methods for wall control applications. The objectives of these explosive technologies and loading techniques are clear cut: to reduce over break beyond the limits planned for the blast, to increase safety and to reduce overall costs. Controlled blasting techniques such as pre splitting and smooth blasting, loading practices such as decoupling and decking, special explosives technologies for wall control such as string charges and low density mixtures are examples of such an effort. These techniques and technologies are not restricted to mining operations solely, they are also applicable in tunnels, highway cuts, ornamental and construction stone quarries as well as underground mine developments such as shaft sinking, drifting and wherever the need to preserve the natural strength of the rock is sought. When applied to open pit mining, the use of wall control methods and related explosives technologies will increase the stability of the final pit walls. The importance of stable slopes cannot be overstated, as they will allow steeper pit slope angles, which in turn will increase ore to waste ratios and, therefore, increase the recoverable ore reserves of the mine. Moreover, increasing the ore to waste ratio implies a reduction in the volume of waste rock that needs to be excavated and thus the high expenses 1

20 associated with removing said waste material. The costs of implementing a wall control program are, more often than not, compensated by the savings associated with a small increase of the overall pit slope angle. An illustrative example of the impact of implementing a wall control program was presented by Jackson (1981), where he considered a 152 m deep conical shaped open pit mine having a 183 m radius at the bottom and an overall pit slope angle of 45 degrees. Ore was assumed to be chalcopyrite disseminated in granodiorite with a density of 2.9 g/cm 3. According to Jackson s example, an over break of 1.83 m resulting from a poor wall control program generated an excess of 253,000 metric tons (approximately 0.7% of the final pit volume) of extra waste material that needed to be removed. In addition to the mentioned issues, a stable slope will benefit the operational performance of the mine by increasing the factor of safety and reducing the need for costly remedial measures such as artificial support and scaling. Similar principles apply to underground mines, whereby reducing dilution in stopes and pillars due to excessive damage is critical for the success of the operation. Many variables play an important role in the success of a blast. Those that cannot be controlled by the operator, such as geological features, limit the design to minimizing their potential effects. However, there are a number of variables (Canmet, 1977) that can be controlled by the engineer for designing a blast, these include burden and spacing, borehole diameter, collar height, sub grade depth, drilling accuracy, explosive type and 2

21 loading density. In turn, the latter can be attained by means of decoupled charges, decked charges or the use of low density explosive mixtures. From the above list, explosive type and loading density are of most interest for the present research thesis, which focuses on the development of a bulk loaded, low density explosive composition aimed for damage control applications, the characterization of relevant performance parameters and the evaluation of its capability to reduce damage under real operational scenarios. 1.2 Research Objectives At the core of this study is the search for a viable low density explosive mixture that better addresses the shortcomings of the existing explosive technologies and loading methods that are presently used in wall control. As with every research thesis, questions on pertinent issues are raised and proposals to address them postulated. This particular thesis seeks to develop a low density explosive that tackles operational issues such as: back break and the poor top fragmentation normally found in open pit and quarry benches due to excessive or badly distributed explosive energy poor loading rates associated with decoupled explosives or decking methods ingredient segregation and its effect on initiation and/or propagation failure 3

22 desensitization due to pre compression (channel effect), sympathetic detonation and similar forms of explosive malfunction observed in decoupled and decked explosive charges need for ultra low pressure ranges in order to represent a feasible alternative to decoupled products while keeping its bulk loading characteristics additional relevant issues including viability of the mixture for mechanized bulk loading equipment, availability of ingredients on an industrial scale, ease of handling and transportation requirements, etc. Based on the above points, the following basic goals are sought for this research thesis: 1. To develop a low density explosive mixture such that it could be used in wall control applications in open pit mines and quarries. To make it a feasible alternative, as both a bulking agent and/or a stand alone product, the mixture should initiate and propagate in a stable manner at densities an order of magnitude lower than those of typical commercial explosives, such as standard ANFO. This characteristic should satisfy any energy requirement between its own low density and the density of the commercial explosive, providing adequate flexibility for damage and fragmentation control. In addition, the low density mixture should be a bulk loaded product, making it a valid substitute to the decoupling and or decking techniques commonly used in mining operations as a tool to control borehole pressure and damage. 4

23 2. To characterize performance parameters of the novel low density composition, focusing on those properties that are relevant to wall control. Given the influence that detonation and explosion pressure have on damage control, a sustained experimental effort was focused on measuring these important parameters, which otherwise are estimated from theoretical considerations. 3. To assess the product s performance in an operational environment as an alternative to existing technology for wall control and to examine operational issues related to its manufacturing process at a semi industrial scale. In summary, the proposed research will consist of three basic stages: first, the developmental stage of the explosive composition, second, the characterization of detonation parameters and other properties that are relevant to the application it was developed for, and third, the large scale manufacturing and evaluation of the product s performance as a damage control tool under real operational scenarios. 1.3 Thesis Outline The remainder of this thesis is organized in the following six chapters: Chapter 2: this chapter starts with a brief description of the alternatives available to reduce explosive energy concentration in the borehole and continues with a review of pressure concepts and the practical role that density has a damage control tool. The benefits of using low density mixtures as an alternative to other explosives and loading techniques are also discussed. The shortcomings of present low density mixtures were 5

24 exposed in an attempt to pave the way for the introduction of the novel low density concept and widen the alternatives in search for solutions to blast induced damage. Experimental research conducted in Russia on polystyrene based low density explosives was relevant literature for this research and has been included as a section in the chapter. A survey of explosive products used in wall control applications however, has been included as reference in the appendix section of the thesis. Chapter 3: this chapter addresses rock breakage mechanisms as they relate to explosives in general but more important, as a way to expose the need for new technology in low density products. In more detail, wall control techniques, including established empirical methods such as buffer, presplitting and cushion blasting are briefly discussed. In addition, design approaches based on particle velocity criteria and other innovative methods contributing to assess and control damage are presented. Chapter 4: this is one of the core chapters of the research and deals with the developmental stages of the novel low density mixture, referred to as low density reactive agent or LDRA for short. Potential bulking agents are outlined and the characteristics of expanded polystyrene, the bulking agent of choice, discussed in more detail. Oxygen balance of the final LDRA mix is discussed in more detail, however, recommendations regarding ingredients, mixing process and density of the final LDRA composition are presented as reference material in the appendix section. Chapter 5: this chapter relates to the explosive characterization stage, where relevant detonation properties of the LDRA are measured, with particular emphasis on velocity 6

25 of detonation (VOD), detonation pressure and explosion pressures. In addition, laboratory observations of the LDRA and its mixtures with regular ANFO prills are also described. Chapter 6: this chapter summarizes the effort involved during the field tests conducted at Chuquicamata Mine, in northern Chile, where the performance of the LDRA and its ability to reduce vibration levels were measured. Issues dealing from the semi industrial scale manufacturing of LDRA to the field experiments in large diameter holes and their results are described. In addition, the conceptual design of a manufacturing plant and studies addressing technical problems exposed by the LDRA during field testing are presented. Chapter 7: this final chapter summarizes the thesis contributions to research and provides information on the patent rights granted for the invention of the LDRA. 7

26 Chapter 2. Review of Low Density Explosive Issues 2.1 Introduction The implementation of any wall control method aims at the reduction of the explosive energy concentration at the final rows of the blast. This in turn can be implemented in the field by any combination of the following basic forms: 1. Decking the charges 2. Decoupling the charges 3. Reducing VOD 4. Reducing the density Although all four alternatives will generate lower borehole pressures and reduce damage, they do it through different avenues: the first two as a consequence of controlling the loading procedure and the last two by controlling explosive properties. The fourth option, controlling pressure by reducing the density of the explosive, is at the core of this research and will be treated separately; however, a brief description on the advantages and limitations of the first three alternatives will be described next for thesis completion. 1. Decking: refers to the axial or longitudinal decoupling of an explosive column by dividing it into decks spaced by gaps. These gaps are usually left with air although other inert materials such as sand or drilling detritus have been used. For a given charge 8

27 length, decking improves energy distribution along the length of the hole; however, it will tend to generate an excess of fines from those sections of borehole in contact with the explosive and a poor fragmentation from those sections which are not. In addition, the need for priming each individual explosive deck introduces an element of risk as a consequence of explosive and/or initiator malfunction due to shock that could lead to misfires or propagation failures. Moreover, the need to prime each individual deck translates into higher requirements for consumables (i.e. initiators and primers) and more important, into a labor intensive and time consuming loading operation seldom justifiable but for very long holes. 2. Decoupling: refers to the radial decoupling of a charge along the borehole length. The borehole pressure is reduced by allowing the gases to expand into the annular space left between the explosive charge and the borehole wall. Loading explosives in a decoupled configuration solves the issue of energy distribution; however, it introduces the risk of desensitization due to channel effect. This effect is characterized by the generation of a precursor air shock traveling in the annulus between the explosive charge and the borehole wall. Under a particular set of conditions, this shock could dead press the explosive ahead of the traveling detonation front to levels above its critical density and result in a propagation failure. To reduce the occurrence of channel effect, special more expensive explosive formulations were developed. Emulsion or slurry explosives, packed in continuous poly sleeves and traced with detonating cord are examples of such products. These are 9

28 relatively expensive products requiring the manual handling of heavy boxes. In addition, loading cartridge sleeves is both a labor intensive and time consuming operation. Bulk products are clearly the best alternative to blasthole loading, however, much of their advantage is lost when poly sleeves, cardboard or plastic tubes are used to attain decoupling. 3. VOD reduction: Mixtures of crystalline ammonium nitrate fines, flaked aluminum particles and lead dioxide (PbO2) powders have proven successful in reducing VOD to values below 1000 m/s while increasing the density beyond 2.0 g/cm 3 (Maranda, 2001). The use of stonegrit (Hardwick, 2002), a powdered by product of furnace operations, has also been mixed with explosives to reduce its VOD, although such mixtures have not developed into wall control field applications. Similarly, dilution of explosives with salt grains reduces the VOD without necessarily increasing mixture density. Being chemically inert, salt does not react during detonation, acting as a coolant and lowering the reaction temperature. However, care should be exercised if this approach is taken, since the resulting mixture will prove less sensitive to initiation. Furthermore, segregation of the salt grains would result in a poor quality mix, which in turn could lead to propagation failures. A 20% salt (weight %) composition has been recommended (Crossland, 1982) as the maximum concentration to be used with ANFO mixtures. 10

29 Slow burning combustible products such as rubber beads will also reduce the VOD of explosives, not necessarily as a consequence of a reduction in density but as a result of a poor contact between fuel and oxidizer. It is worth noting that the addition of ingredients that promote non ideal reactions will induce a reduction in VOD and in turn shock energy, although gas energy and total explosive energy, which influence borehole pressure histories the most, will not necessarily be reduced. VOD decrease by non ideal detonation is not considered a proper way to achieve wall control given that it does not take into account the effect of gases. Explosives capable of delivering low energy, pressure and charge concentration are desirable for wall control, all characteristics that can be achieved by low density explosive mixtures. 2.2 Review of Pressure Concepts From a damage control point of view, the most important parameters related to the performance of an explosive are the detonation pressure, the explosion pressure and the borehole pressure. Since this research effort deals with the development of a lowdensity, therefore low pressure explosive, it is important to review some of the basic concepts defining these particular parameters and the formulae that have been proposed to estimate them Detonation Pressure The detonation pressure refers to the pressure developed right behind the detonation front before any gas expansion takes place. This pressure is usually referred to as the 11

30 Chapman Jouget state, or CJ state for short. For ideal explosives, full chemical reaction (complete oxidation) has been achieved at this state; however, in the case of non ideal explosive products, the chemical reaction continues beyond the CJ state and into the expansion zone of the reaction products. From the conservation of momentum equation describing shock phenomena, the detonation pressure is given by: P d = ρ D u Equation 2.1 where ρ is the original density of the explosive, D the velocity of the detonation front and u the particle velocity behind the detonation front. Equation 2.1 can also be expressed as: P d 2 ρ D = Equation 2.2 (1 + γ ) CJ where γcj is the coefficient of adiabatic expansion (i.e. specific heat ratio) at the CJ state. For condensed high density explosives (density between 1.0 g/cm 3 and 1.80 g/cm 3 ), γcj is approximately equal to 3. By replacing this value in Equation 2.2, the well known expression used to estimate detonation pressure is obtained: P d 2 D = ρ Equation

31 In addition, semi empirical formulae to estimate detonation pressure have been proposed by different authors. Cooper (1997) suggests the following equation: P d = ρ D ( ρ ) Equation 2.4 where Pd is given in (GPa), ρ in (g/cm 3 ) and D in (km/s). Zhou and Yu (1992) proposed an equation based on the atomic composition of typical Ca Hb Nc Od explosive: 2 2 P d = ( 1.60 G ) ρ Equation 2.5 where the constant G = ( 0.5 b + c + d) /(2 a + b + c) and ρ the density in g/cm 3. It is worth noting at this time that the value of the coefficient of adiabatic expansion γ is a function of pressure and does not remain constant during the expansion process. Moreover, it has been shown experimentally (Hustrulid, 1997) that the magnitude of this coefficient is affected by external factors such as charge diameter, density and packing material. Several authors have proposed semi empirical formulae to estimate γ at the CJ state as well as at lower pressure ranges. These include Kamlet and Jacobs (1968): γ CJ = 0.655/ ρ ρ Equation

32 Defourneaux (1973): γ CJ = ρ Equation 2.7 and Cooper (1997): 0.04 γ CJ = 1/(1.386 ρ 1) Equation 2.8 where ρ (g/cm 3 ) is the density of the unreacted explosive Explosion Pressure The explosion pressure is defined as the pressure exerted by the expanding detonation products when they occupy the original volume of the explosive charge. If we assume a γ law equation of state with a value of γ remaining constant throughout the expansion process from the CJ state to the original volume of the charge, then the explosion pressure equals one half the detonation pressure (Zerill, 1981). Therefore, using Equations 2.2 and 2.3, the corresponding explosion pressure expressions are obtained: P e 2 1/ 2 ρ D = Equation 2.9 (1 + γ ) CJ P e 2 D = ρ Equation Several empirical formulae were also developed to estimate explosion pressure. These include the one proposed by Gehrig (1982) from Gulf Oil Chemicals, with Pe given in (Pa), D in (m/s) and ρ in (g/cm 3 ): 14

33 P e 2 = ρ D Equation 2.11 The one given by Sanchidrian (1996), with Pe in (Pa), D in (m/s) and ρ in (g/cm 3 ): 2 P e = 228 ρ D /( ρ) Equation 2.12 The one proposed in the Canmet Report (1977): P e = N ( ρ D 2 ρ ) Equation 2.13 where ρ is expressed in (g/cm 3 ) and N(ρ) is a coefficient that depends on explosive density and basically substitutes the 1/8 factor embedded in Equation 2.10 In addition to the above, Ouchterlony (1997) proposed the following equation: 2 γ ( γ + 1) P e = ρ D γ ( γ + 1) Equation 2.14 where γ is actually γcj, the coefficient of adiabatic expansion at the CJ state calculated 2 from Fickett and Davis (1979) according to γ = ( 1+ D / 2 Q) explosion expressed in (kj/kg). 1/ 2 where Q is the heat of Borehole Pressure The borehole pressure refers to the pressure exerted by the expanding detonation gases against the wall of the borehole. When an explosive charge is bulk loaded into the blasthole (i.e. fully coupled charge), the borehole pressure and the explosion pressure 15

34 equate, however, when decoupled explosive charges are used, the borehole pressure will have a reduced value which will depend on the extent to which the products of detonation are allowed to expand. Assuming the expansion process from the explosion state to the borehole state is adiabatic (i.e. PV γ = constant) with a constant γ coefficient, the borehole pressure (Nie, 1999) will be given by: 2 γ P b = Pe ( re / rb Le / Lb ) Equation 2.15 In other words, the borehole pressure is a function of the explosion pressure (Pe), the ratio of charge to blasthole radius (re/rb), the ratio of charge to blasthole length (Le/Lb) and the value assigned to γ. The formula given in Canmet (1977) presented below has basically the same form as Equation 2.15 above: P b = N ( ρ ) ρ VOD ( re / rb C ) Equation 2.16 where C represents the axial coupling ratio (Le/Lb) and γ = 1.2 (thus 2γ = 2.4). The specific heat ratio Gamma (γ) in Equation 2.15 is the coefficient of adiabatic expansion for the detonation products at the lower pressure range, in other words, at venting conditions, where its value greatly differs from the γ = 3 typically assigned to 16

35 the higher CJ pressure state. According to Ouchterlony (1997), γ values ranging between 1.1 and 1.5 have been used by different authors. Atlas Powder (1987) suggested γ = Summary of Pressure Concepts Summarizing the previous paragraphs, three interrelated and particularly important pressure concepts are to be considered when selecting an explosive product for wallcontrol applications; these are the detonation, the explosion and the borehole pressures. The detonation pressure is a function of the explosive density, the velocity of detonation and the coefficient of adiabatic expansion (usually assigned a value of 3); its amplitude is determined from the conservation of momentum equation. This conservation equation is appropriate for ideal conditions involving full chemical reaction at the CJ state, behavior that is best represented by condensed high density explosives such as most military formulations. However, commercial explosives exhibit a higher dependence on external and internal factors such as charge diameter, heterogeneity of the mixture and degree of confinement. These factors will affect the explosive s ability to reach full chemical reaction at the CJ state and, as a result, they will exhibit the non ideal behavior that characterizes most commercial products. The explosion and borehole pressures are both functions of the detonation pressure; thus, all the comments mentioned for the detonation pressure also apply. The amplitude of the explosion pressure is usually taken as being half the value of the detonation pressure; otherwise, the empirical relationships presented can be used to estimate it. The borehole pressure, in turn, is expressed as a function of explosion pressure and 17

36 estimated assuming a constant value for the specific heat ratio (about 1.3) during the gas expansion in the annular space created by the decoupled charge. 2.3 Role of Density on Damage Control Regardless of the mechanism responsible for fracturing the rock and the technique used to control it, the importance of detonation pressure (hence explosion and borehole pressures) cannot be overstated. A borehole pressure above the dynamic compressive strength of the rock will generate excessive fines from the pulverized zone around the borehole and produce unwanted over break. On the other hand, a pressure below the dynamic tensile strength will not fragment the rock. The pressures generated by practically all military and commercial explosives are well above the compressive strength of rocks. As a consequence, in fully coupled holes, a certain degree of pulverization and fracturing is to be expected unless decking or decoupling is implemented to lower the loading density. Given the limitations of these loading techniques, a window of opportunity arises for the introduction of explosive technologies that help control these pressure amplitudes. For convenience, explosion pressure Equation 2.10 is recalled below, P e e D = ρ 8 2 The above equation implies that explosion pressure depends on both, density and velocity of detonation; however, the stronger dependence on VOD that the equation 18

37 seems to indicate is somehow misleading. It has been empirically determined that a linear relationship exists between velocity of detonation and density over reasonable density ranges (Cooper, 1997). This can be expressed as: D = a + b ρ Equation 2.17 e where a and b are empirical constants specific to a particular explosive product. Rearranging Equation 2.10 and Equation 2.17, the following expression is obtained: P e ( b ρe + 2 b ρe + a ρe ) = Equation The cube function dependence that pressure has on density can now be clearly inferred from Equation As a consequence, from all the parameters defining explosion pressure, density is the one that plays the greatest role. It is for this reason that adjusting the density to match specific requirements is a very effective and practical way of reducing the pressure amplitude and controlling damage. Density reduction finds its main application in low density explosive mixtures, commonly applied in wall control blasting. In addition, low density mixtures amenable to bulk loading will render unnecessary the implementation of decoupling and or decking methods, both of which represent a labor intensive and costly approach towards damage control. 19

38 Reducing density by the addition of inert or combustible bulking agents (rather than chemical gassing) is considered one of the simplest ways to reduce the explosion pressure. However, serious consideration should be exercised upon mixing and loading to prevent excessive segregation of the bulking agent from the rest of the explosive ingredients. Selection of ingredients, manufacturing process and loading techniques will largely dictate the chances of success or failure of a low density mixture in overcoming the undesired effects of segregation, in particular the occurrence of propagation failures. Developing an explosive with those considerations in mind is at the core of this research thesis. 2.4 General Considerations Related to Low Density Explosives Much effort has been placed in developing low density explosives that are both cost effective and can keep up with high loading rates. A literature review of selective low density explosive products and techniques used in wall control throughout the last years has been included in Appendix 2.1 for thesis completion. Many of these low density explosives consist of ANFO prills mixed with bulking agents, the latter products limiting to a great extent the behaviour of the final dry mixture. One such limitation refers to the degree of dilution that can be attained by addition of bulking agents before the mixture fails to reliably initiate and or propagate. 20

39 In addition, the sensitivity of the mixture will depend not only on the degree of dilution attained but on the type of bulking agent used. For a given bulking agent density, greater dilutions can be obtained when combustible agents such as sawdust and polystyrene rather than inert agents such as perlite and vermiculite, are selected. Contrary to inert fillers, combustible bulking agents will participate in the chemical reaction and contribute to sustain detonation. The reduced pressures generated by low density explosives makes them ideal for a variety of situations where preserving the stability of the wall is required. For a given explosive weight, the use of low density products will result in longer column charges; therefore, an improved energy distribution within the blasthole without increase of the powder factor. Low density, well distributed charges will produce a more evenly distributed fragmentation, eliminate oversize from large collars, generate less environmental impact due to lower vibration levels and reduce damage; the latter more often than not, being the sole objective for its use. Another important consideration related to dry low density mixtures refers to segregation of the bulking agent from the ANFO prills during mixing and loading. The higher the dilution of the ANFO mixes, the higher the risk of initiation and or propagation failure, consequently, the more important the control of segregation becomes. Although segregation is to some extent controllable, it is nevertheless an unavoidable fact that will contribute to increase the risk of initiation and or propagation failure of the mixture. In many instances, a bulk emulsion matrix has been used as a 21

40 binding material to make the mixture sticky and prevent segregation. This approach not only carries a cost but, depending on the bulking agent used, could result in the breakdown of the emulsion and the loss of its binding characteristics which in turn could lead to failure. Perlite mineral has proven to induce breakdown of emulsions. The reduced shock energy that characterizes low density, low VOD products makes them good candidates for fractured or soft rock applications such as coal. In open pit coal mines having fairly undulating coal seams, the use of low shock products will allow all blastholes to be drilled to coal (i.e. the top of the coal seam) rather than every 4 to 6 holes as it is usually done. Loading to coal without a need to provide for stand off stemming to protect the seam from blast damage will not only be faster and less labour intensive but more importantly, it will generate more even floors and reduce the chances of leaving hard layers of unbroken rock above the coal seam, the latter requiring the use of heavy earth moving equipment (dozers and rippers). If the bulking agents used for diluting the explosive are such that they promote gaseous products of detonation rather than shock energy, the resulting higher gas component will induce gas penetration into the coal rock interface and promote separation between them. When drilling and loading to coal, this will ease the cleanup phase of the coal roof and reduce costs. Another benefit arising from the reduced pressures and shock velocities inherent in low density mixtures relates to the reduction of fines produced around the blasthole as a consequence of the compressive stresses generated by the detonation. Excessive fines, particularly in some commodities such as coal and iron ore, have detrimental effects 22

41 upon its product and the revenue of the operation. Field experimentation conducted by Sheahan et al (1998) using 65 mm 102 mm diameter holes in a 5 m high granite quarry bench showed a linear correlation between fines generated and the relative bulk strength of the explosives. During the experimental phase, Sheahan et al defined fines as those fragment sizes less than 16 mm and their results showed that a polystyrene/anfo mixture at a density of 0.28 g/cm 3 generated 1.1% of fines compared to 3.2% if regular ANFO prills were used instead. If fines were defined as particle sizes less than 38 mm, then these percentages become 3.2% and 5.8% respectively. Although the basis by which fines were defined is somewhat arbitrary, the results clearly indicate a definite decrease in the generation of fines as the density of the explosive is reduced. Other issues worth considering refer to water resistance and oxygen balance. Emulsion based low density products are capable of providing water resistance; however, diluted ANFO mixtures can be applied only where dry ground conditions prevail unless protection from the detrimental action of water is provided. Oxygen balance is an issue of concern when developing low density compositions, something that becomes particularly difficult to achieve when diluting ANFO with organic fillers such as sawdust, polystyrene, bagasse and similar agents. The use of organic fillers will result in fuel rich mixtures that will generate toxic fumes; however, these fumes are for the most part carbon monoxide (CO) and carbon dioxide (CO2) and as such they have a relatively low toxicity. Oxygen rich mixtures on the other hand, will generate nitrogen oxide fumes, highly toxic, which will form clouds that tend to hang 23

42 up for considerable long times before diluting in the air. This is a serious concern in operations such as open cut coal mines where the emulsion based Heavy ANFO typically used tends to generate this kind of problem in the presence of water. Underground operations are a much more restricted ground for applying low density products if their compositions are not oxygen balanced. The generation of toxic fumes precludes the use of many products unless the ventilation conditions provided by the mine are such that the mine inspector allows use of the explosive. Nevertheless, low density mixtures can be balanced and have been used in underground operations as recorded by Forsyth et al (1997), who recommended the use of a perlite/anfo mixture as a replacement of standard ANFO alone. 2.5 Russian Experiences using EPS based Explosives As discussed in coming chapters, expanded polystyrene (EPS) beads were selected as the bulking agent of choice for the low density reactive agent, therefore, a literature review of the use of EPS in explosive mixtures is deemed appropriate. As it will become apparent in the next sections, much of the investigations on the detonation physics of binary explosives containing expanded polystyrene, as well as those on the shock behaviour of polystyrene alone, were found to come from work conducted by Russian scientists during the last three decades. 24

43 Rozovoski (1976) conducted experiments to determine detonation rates of asphalt, polystyrene foam, polyurethane foam, porous rubber and other combustible materials mixed with liquid oxygen. Kuznetov (1977) conducted studies on pulsating detonations observed during shocking of low density explosives containing mixtures of powdered trotyl/eps and Ammonite/EPS. Ammonite consists of a mix of 21% trotyl and 79% ammonium nitrate prills. The phenomenon of pulsating detonations, whose existence was first detected in gases and later in a number of liquid explosives, is characterized by an unstable detonation front indicative of the temporary collapse or breakdown of the chemical reaction. Unstable detonation fronts are not characterized by the smoothness typical of high density explosive detonation fronts, but by a rougher uneven front as a consequence of the chemical reaction not being accomplished simultaneously over the cross section of the charge. According to Dremin (1983), this uneven characteristic manifests itself by a chaotic curvature and glow of light emitted by the detonation front as observed from photographic experiments. This instability phenomenon cannot be appropriately explained by the hydrodynamic theory of detonation, which is thermal in nature and based on the formation of a single discontinuity across the entire shock front. Although not well understood, the collapse of the chemical reaction is attributed to the action of the rarefaction wave cooling the reaction zone below the temperature required for selfpropagation. The explosion products in the shock front where the explosive actually 25

44 detonated will act as a piston, compressing the unreacted explosive material ahead and subsequently, reactivating the detonation in the form of an adiabatic thermal explosion. Shvedov (1985) analyzed the ability of low density explosive mixtures to sustain detonation in large charges (i.e. the stability of the detonation front). Different explosive mixtures were evaluated, including a mix of ammonium nitrate and foamed polystyrene at a density of 0.2 g/cm 3. These mixtures were initiated under confined conditions (steel tubes) with diameters ranging between 34 and 42 mm and lengths between 3 and 5 meters. A low detonation velocity regime characterized by periodically bulged and ruptured sections along the entire length of the steel tube was observed. The location of these sections was in perfect synchrony with the measured VOD ( m/s at the bulged areas and m/s in the rupture sections). Shvedov concludes that the pulsating process consists of a low velocity detonation regime followed by an unstable combustion process and followed again by a low velocity regime, with the whole process repeating itself. Similar experimentation relevant to unstable propagations, this time using powdered trotyl of different particle sizes (between 0.2 mm and 1 mm) showed the shock process damping when larger particles were used. Although the reason was attributed to a behavior analogous to channel effect, where a precursor air shock interrupts propagation of the explosive shock, it is reasonable to speculate that the lower reaction rates associated with the larger particles will play a role on the observed behavior. 26

45 2.5.1 The Shock Behavior of Expanded Polystyrene (EPS) The unstable regimes produced by low density binary mixtures using EPS beads led researchers to believe that their role in the detonation process goes beyond that of simple fillers. As a consequence, additional studies were conducted to investigate the role that the expanded polymer itself plays when subjected to the action of a shock. In an attempt to understand the decomposition behavior of the polymer at various densities, Nifad ev et al (1992/a) experimentally focused on the phase transitions experienced during shock loading of EPS and the nature of the subsequent gasification wave. Experiments consisted of confining EPS granules in 1 to 1.5 m long, 35 mm diameter steel tubes and shocking them by means of an explosive charge. Different EPS densities and explosive weights were used. Furthermore, in order to separate products of detonation from igniting the expanded polymer, a 10 mm steel plate barrier was used in selected tests to separate the explosive charge from the EPS. Four different decomposition behaviors were observed: 1. Total gasification of the EPS accompanied by the formation of a thin sooty deposit on the steel tube walls 2. Agglomeration of the EPS granules into cm long lumps due to heating and breakup of the beads surface during compression. Results for EPS at g/cm 3 shocked with 20 g of ammonite in contact with a steel plate are shown in Figure

46 Figure 2.1: EPS at g/cm 3 subjected to shock from 20 g of ammonite through a steel plate (Nifad ev et al, 1992/a) 3. Gasification of the coarsest EPS granules with no major effect on the smaller sizes as shown in Figure 2.2. The EPS beads at a density of g/cm 3 were shocked through a steel plate with 20 g of ammonite Figure 2.2: Before and after effect of shock on EPS at a density of g/cm 3. Shock was provided by 20 g of ammonite through a steel plate barrier (Nifad ev et al, 1992/a) 4. Partial gasification of the polystyrene foam with the formation of solidified products of decomposition as shown in Figure 2.3 for an EPS density between g/cm 3. Shock was provided with 100 g of ammonite without the protective steel barrier. 28

47 Figure 2.3: Partial gasification of EPS granules at a density of 0.05 to 0.10 g/cm 3, subjected to explosive shock induced by 100 g of Ammonite without a steel plate barrier (Nifad ev et al, 1992/a) In similar experiments conducted on glass tubes instead of steel tubes to allow the use of photographic techniques, Nifad ev et al showed that ultra light EPS (~0.006 g/cm 3 ) gasify faster and produce a weaker luminescent zone than higher density (0.025 g/cm 3 ) smaller EPS granules, which gasify slower and produce a much brighter light band. This behavior was attributed to the fact that the lower density larger granules had a wider particle size range, the larger ones receiving the brunt of the compressive shock and gasifying first under lower shock intensities. The pulsating nature of the gasification wave was observed during these tests as luminescent bands separated by darker parallel bands. According to Nifad ev et al s findings, the brightness, uniformity, width and separation of these luminescence bands depended on the density, size distribution and strength of the initiating explosive. The closed cell arrangement of EPS contains a large quantity of air micropores which led 29

48 researchers to believe that they can act as hot spots upon compression by the shock wave, which in turn could lead to adiabatic thermal decomposition and heating of the granules to induce gasification. However, the photographic analysis contradicted such a line of thought, showing that the granules were subjected to intense heating without noticeable change in their volume. Experiments conducted with mixtures of 80% volume of EPS (at g/cm 3 ) and 20% volume of ammonium nitrate prills, which resulted in a bulk density of 0.20 g/cm 3, showed a similar pulsating pattern of the gasification front, this time with the luminescence bands being brighter, more uniform and closer than when no ammonium nitrate was involved. The series of experiments led Nifad ev et al to the following conclusions regarding the possible role and gasification process of expanded polystyrene: Shock decomposition of expanded polystyrene can occur in the detonation regime under certain conditions of initiation power and density Gasification of foamed polystyrene is highly dependent on the energy dissipation mechanism of the polymer, which in turn is a function of its compressibility and behavior under high pressures The ability of the expanded polystyrene to gasify explosively indicates its participation as an active component in a low density explosive mixture The ability to detonate low density explosive mixtures (<0.20 g/cm 3 ) depends on the density of the foamed polystyrene selected 30

49 2.5.2 Detonation Behavior of EPS and AN Mixtures Further research published by Nifad ev et al (1992/b) focused on the behavior of low density explosives consisting of a mixture of expanded polystyrene and ammonium nitrate. The objective was to analyze the role that EPS had on the initiation and propagation of detonation for mixtures at densities below 0.20 g/cm 3. The experiments proved the existence of a pulsating precursor wave moving ahead of the detonation wave. This precursor wave was attributed to the gasification of the polystyrene foam, with the detonation process itself taking place in the gas phase products of thermal decomposition of the mixture. The research established that the process of subjecting foam polystyrene to strong shocks and compressing it to densities of 0.03 g/cm 3 to 1.05 g/cm 3, was accompanied by a phase transition of the polymer from a solid to a gas state. Moreover, for the case of ultra low density polystyrene foam (0.005 g/cm 3 to g/cm 3 ), weaker shocks also attained a similar transformation, with the gasification front traveling distances of 2.5 m to 3 m in 35 mm diameter steel tubes. This led to the realization that, under certain conditions, the gasification is a self sustaining process that is associated with the liberation of heat in a fashion similar to a detonation wave. Based on the gasification behavior of foamed polystyrene, it was obvious that the expanded polymer played an important if not critical role on the detonation of lowdensity explosive mixtures containing such polymer as its diluting agent. Experiments evaluating ammonium nitrate/polystyrene foam mixtures at densities ranging between 31

50 g/cm 3 under confined conditions (steel tubes of 28 to 70 mm in diameter) concluded that the detonation of these low density mixtures consisted of three distinct stages: compression, heating and gasification of the expanded polystyrene as a result of the initiating explosive; heating and decomposition of the ammonium nitrate and interaction of the decomposition products of ammonium nitrate and foam polystyrene. Visual observation of the steel tubes showed a series of evenly distributed bulges along their lengths as well as an alternating sequence of light/dark bands in between the bulges, the latter ones being visible on both the inside and the outside of the tubes. Figure 2.4 shows the bulges produced in a 35 mm diameter steel tube by the detonation of an EPS ammonium nitrate mixture at a density of 0.07 g/cm 3. Also noticeable in the picture is the banding produced by a similar mixture detonating at a density of 0.09 g/cm 3. Bulges Bands Figure 2.4: Bulging and banding effect produced by low density mixtures of ammonium nitrate and expanded polystyrene (Nifad ev et al, 1992/b) 32

51 Photographic analysis of the experiments concluded that the banding is a result of the pulsating nature of the gasification wave while the bulges are attributed to the interaction of the ammonium nitrate and the products of decomposition of the polystyrene foam. The distance between bands corresponds to the oscillation length of the gasification wave while the one between bulges matches the length of the detonation wave, with both of them depending on the density of the explosive mixture. Moreover, according to the same author, the period (T) of oscillation of the gasification wave was 5 to 30 times smaller than the one corresponding to the detonation, thus, the frequency (1/T) of the detonation front results in much lower values than the gasification one. Following the tests, a mechanism of detonation of low density AN/EPS mixtures with densities below 0.20 g/cm 3 has been proposed by Nifad ev et al. It consists of the following stages: 1. The priming charge generates a compression wave which travels into the low density mixture, initiating the chemical reaction with a delay that corresponds to the time needed for the mixture to chemically react. As a consequence, the compression wave precedes the reaction wave by a certain delay. 2. The amplitude of this compression or thermal precursor wave is sufficient to heat the EPS and cause thermal decomposition, producing an explosive gasification or expansion of the EPS. This gasification zone contains products in a plasma state, producing high temperature turbulent motion jets. 33

52 3. These jets and turbulent behavior induce a precursor which is then supported by the expansion of the EPS, producing the fluctuating behavior that characterizes low density products 2.6 Chapter Conclusion One of the main objectives of this chapter was to review pressure concepts and discuss the importance that density has on this parameter and consequently on damage. General comments describing the advantages and disadvantages of using low density products to reduce borehole pressure and control damage are also discussed. Experimental research conducted in Russia on mixtures of explosives and expanded polystyrene beads, considered relevant for the present studies, has also been included in this chapter. For reference and thesis completion, a survey of past and present wall control products has been included in Appendix

53 Chapter 3. Rock Breakage and Wall Control Topics 3.1 Introduction It is not the objective of this thesis to detail the mechanisms of rock fragmentation and wall control techniques; however, since explosives play a major role in the final outcome, it would be appropriate for completion to review issues involved in the fragmentation process due to explosive loading as well as introduce a brief description of the techniques presently available to control rock damage. 3.2 Mechanisms of Rock Breakage At the present time, the mechanism of rock fragmentation by the action of explosive loading cannot be fully described by any single theory. Questions have been and still are being raised as to the exact mechanism responsible for breakage; whether it is the shock induced stress, the expansion of the detonation gases or more probably a combination of both, remains the subject of ongoing debate and controversy. A good summary has been presented by Winzer et al (1983), whose findings are selectively addressed during this section. There are just too many variables influencing fragmentation for a comprehensive theory to be able to describe the breakage mechanism, therefore, it comes as no surprise that contradicting theories regarding the mechanism of rock fragmentation and the role played by the field parameters affecting it have been postulated. 35

54 The main debate on the actual mechanism responsible for fracture and fragmentation started in the fifties with work presented by Obert and Duvall (1950), Hino (1956) and Duvall and Atchison (1957); all of them strong supporters of the stress wave being predominantly responsible for fracturing the rock. According to this school of thought, the predominant mechanism of failure was spalling: tensional stresses generated by the reflection of the initial compressive stress pulse at a free surface. High speed movies of blast practices however, suggested that spalling could not be the sole governing factor behind rock fragmentation, except perhaps at the vicinity of the free face. The stress based theory was soon challenged by findings of Fogelson et al (1959), Nicholls et al (1962), Langefors and Kihlstrom (1963) and Porter and Fairhurst (1970); who concluded that the useful energy contained in the stress wave was relatively small compared to the total available explosive energy, with the difference mostly stored in the expanding products of detonation. These researchers favored the quasi static gas pressurization as the preferred breakage mechanism. In their view, the predominant mechanism of rock breakage was due to the action of the gases pressurizing and extending the initial radial fractures until they intersected the free face. They found that the reflected tensile stresses were not a major factor in blasting operations since they were too weak for the large burdens typically used in operational environments. Experimentation conducted in Plexiglas and small rock models by Kutter and Fairhurst (1971) and Field and Ladegaard Pedersen (1971) concluded that both stress waves and gas pressurization contributed to rock fragmentation, with the latter being relevant only 36

55 after the rock has been preconditioned by the stress wave. According to their findings, the tensional stress field generated by the detonation of the explosive will first induce radial fractures at the borehole wall and precondition the rock mass. Following this preconditioning effect, the expanding gases will penetrate the radial cracks, pressurizing them and extending their lengths. The presence of a free surface will contribute by extending the gas pressurized radial fractures in a direction perpendicular to that surface. Later, in yet another apparent turn of events, a group of researchers favored gas pressurization as being the dominant factor in rock breakage. Hagan (1983) suggested that the strain wave was responsible for fracturing in the immediate vicinity of the borehole, while the remaining energy (named heave energy) was responsible for extending the initial fractures and inducing fragmentation by the longer wedging action of the gas pressure. In addition, the gases would also provide the energy for displacing the rock mass. Persson et al (1969) supported the same concept as Hagan. McHugh (1983) designed experiments to assess, independently of each other, the effects of gas pressure and stress upon the extension of radial fractures. His findings indicated that gas pressurization could extend fracture lengths 5 to 10 times; however, his experimental results cannot be confidently transferred to field conditions since they were conducted in Plexiglas models. Extending McHugh s approach to field experiments, Brinkman (1987) conducted a series of tests grouting metal pipes in boreholes, with and without stemming. Stemmed holes 37

56 will be affected by both shock and gases, while in un stemmed holes only the effect of shock will prevail, since gases were allowed to vent to the atmosphere. In his conclusions, Brinkman claimed the shock wave was the main factor responsible for the initial fracturing in the vicinity of the blasthole, providing the fracture network for gas penetration and fracture extension, which in turn was followed by final fracturing, fragmentation and mass displacement. Another line of thought arose from studies conducted by Barker et al (1978), Fourney and Barker (1979 a), Fourney et al (1979 b) and Holloway et al (1980) at the University of Maryland, using flawed and unflawed models made of Homalite, a transparent polymer similar to Plexiglas, as well as models made of rock. While recognizing the role that reflected stresses at the free surface have on extending radial fractures, their experiments emphasized the role that existing discontinuities play in generating multiple reflections in the rock mass and causing the creation of new fracture networks and, as a consequence, an overall smaller fragmentation size. Small bench and full scale production tests conducted by Fourney et al. (1983) and Winzer and Ritter (1979) concluded that pre existing fractures act as the source of new fractures and remain active well after detonation of the explosive, with new fractures developing at the free face at twice the time it takes for the P wave to cover the burden distance. Moreover, gas venting was observed to occur late in the event through already opened cracks, an indication that much of the fracture network is not pressurized; hence, gas does not contribute to the creation of new fractures. Also, in production scale tests, 38

57 fractures seemed to originate and propagate from existing bedding planes and joint systems, replicating the experiments conducted in homogeneous Homalite models. More recent studies, including fragmentation modeling, have helped clarify some of the processes involved in the dispute over which mechanism plays the predominant role in fragmentation. What is certain now is that no single mechanism is adequate to fully describe the fragmentation process for the typical conditions encountered in the field; neither stress nor gas pressure would be effective in producing fragmentation by itself. 3.3 Wall Control Techniques Wall control methods can be grouped into three broad categories: those relying on empirical approaches, those based on particle velocity as the parameter responsible for damage (PV damage criteria) and lately those using fracture mechanics and crack propagation concepts as a design tool; the latter is not yet established in the industry. All the previous methods strive for the same goal: to control damage. In order to achieve such a goal, a reduced explosive energy concentration at the final wall is of utmost importance. These can be obtained by changing drilling parameters, controlling the energy output of the explosive and altering the loading density of the explosive. Given the direct role explosives play on inducing damage, and in turn, the influence explosive density has on controlling its energy output, it seems appropriate to briefly describe selected wall control methods, including some innovative techniques that have been evaluated in the recent past. 39

58 3.3.1 Empirical Approaches Wall control techniques such as line drilling, presplitting, buffer, cushion and smooth blasting are, for the most part, empirical methods based on operational experience. A brief description of them follows Presplitting The presplitting technique consists in isolating the wall by creating an open crack along its perimeter before the production blast takes place. Presplit holes have smaller diameters, spacing and loads than those from the main blast. The presplit row is always fired prior to the production blast; however, it can be fired days before the actual drilling of the main blast takes place or just a few milliseconds ahead of the production holes. For the latter case, it is recommended to initiate the presplit row about 100 to 200 ms before the last row of the main production blast to ensure the crack barrier will be effective in reducing the vibrations arising from the production holes. For economical and operational reasons, the present trend in large mining operations is to use larger diameter holes in the presplit line, which consequently leads to a larger spacing between holes, the use of decoupled and/or decked charges or the loading of low density explosives. Unless air blast noise becomes an environmental issue, stemming the presplit boreholes is usually avoided in order to allow immediate venting of the gases, which in turn helps reduce unnecessary vibrations. Presplit blasts are conducted under conditions of infinite confinement (large burden) and high load per delay due to the simultaneous detonation 40

59 of its holes. As a result, they will generate higher vibration and noise levels than other methods. Due to the need for smaller diameters and closer spacing, presplit rows will require more drilling relative to production rows, thus the need for more small drilling equipment in order to keep up with production rates. Spacing of presplit rows can be designed by means of analytical approaches such as the one provided by Canmet (1977), unless the existence of rock discontinuities dictates otherwise. According to Canmet, spacing should be less than twice the spacing of major open joints. Holes are usually charged up to ¾ of their length and stemming is not required unless air blast represents a problem. In the presence of incompetent ground rock, unloaded guide holes can be placed between loaded holes to ensure proper splitting. Alternatively, empirical based formulae such as the ones suggested by Konya (1984) can also be use as a first approximation to a presplit design. The latter are as follows: Spacing 20 Equation 3.1 r b LoadingDensity 2 = Spacing Equation 3.2 In Equation 3.1, spacing and borehole radius (rb) can be expressed in any given unit of measurement, while in Equation 3.2 spacing is expressed in meters and the corresponding loading density in kilograms/meter. 41

60 Experimental values for presplit blast design have also been recommended by Du Pont de Nemours (1976) and Hustrulid (1999) Suggested Presplitting Mechanism Questions have been raised as to the exact mechanism involved in forming the presplit fracture. One school of thought sustains that the fracture is generated by the overlapping of tangential tensile stresses resulting from the collision of the stress waves in between two contiguous boreholes. This theory is difficult to sustain in practice given that for the wave to collide, initiation requirements need to be extremely precise, something that seldom happens with the technology commonly employed (blasting caps, detonating cord). However, the relatively recent introduction of electronic detonators allows for an accurate control of timing, therefore true simultaneous initiation of the presplit line can be achieved. A second school suggests that the radial fractures generated by the detonating explosive extended by the wedging action of the expanding products are the preferred fracture mechanism. This approach led to the development of a widely spread formula used for spacing calculation in presplitting (Canmet, 1977). This formula was derived from similarities with a pressurized cylinder having an infinite wall thickness and zero external pressure. The final design equation proposed by Canmet (1977) is: Pb + T S 2 rb Equation 3.3 T 42

61 where S is the maximum spacing required between consecutive holes for a split fracture to occur and T represents the dynamic tensile strength of the rock. Equation 3.3 shows the important role that borehole pressure and dynamic tensile strength of the rock play in determining the spacing required for splitting the rock. It is worth noting however, that the validity of this equation is limited to homogeneous rock conditions, which is not the case in most field operations where rock discontinuities will dictate, to a great extent, the outcome of the presplitting technique. In addition, given that the rock is subjected to a dynamic rather than a static load, for tensile failure to occur, the magnitude of the tensile stress should at least equal the dynamic tensile strength of the rock. Dynamic tensile strengths are several times larger than their corresponding static values. Rinehart (1965), Birkimer (1970) and Mohanty (1987) provide dynamic tensile strength results for different types of rock Buffer Blasting This method involves modifications of the last and occasionally the two last rows of the blast by reducing spacing, burden and explosive loads relative to the ones design for the main production blast. This technique is usually accompanied by presplitting and if possible, makes use of pre existing fractures or discontinuities. Spacing and burden typically range between ½ and ¾ of those corresponding to the main production blast, while powder factor is kept at around ½ of the one used for production holes. Buffer holes are fired after the production rows and the presplit line, if the latter is part of the 43

62 blast design. Buffer hole diameter and depth should remain the same as for production blastholes; however, if they are above a future bench crest, they should not be subdrilled. According to Chiappetta (1994), buffer loads can be designed on the following equation, which is based on cratering theory: d c 1/ 3 = A W Equation 3.4 where A is an empirical constant function of the rock nature, W is the weight (kg) of the top first 8 diameters of explosive column and dc represents the distance (m) from the center of mass of the explosive to the collar of the blasthole. Values for the constant A as suggested by Jackson (1981) ranged from 1.1 m/kg 1/3 for hard/brittle rocks to 1.8 m/kg 1/3 for soft/plastic rock conditions. Smaller values of A result in smaller values for dc, which in turn implies longer explosive columns for hard formations Cushion or Trim Blasting This technique is practiced in surface blasting and is similar to smooth blasting in underground operations; however, the holes are fired after the muck from the main production blast is removed. The objective of this technique is to trim or slash to the final excavation line while minimizing damage and maintaining stable slopes, hence the other names it is known for: trim or slash blasting. Cushion holes can be drilled before or after the production blast takes place. They are lightly loaded with cartridge charges stringed to each other with detonating cord and 44

63 distributed evenly along the explosive column. Stemming material such as sand, gravel or crushed stone is added to cushion the shock and protect the finished wall. A coupling ratio of at least 0.5 is recommended in order to maximize the cushioning effect, which is the reason why it is convenient to lay the charges against the wall closer to the main blast before adding the stemming material. At the collar, cushion holes should be left unloaded and fully stemmed. Typical cushion diameters range between 50 mm and 165 mm and are preferably initiated simultaneously or with minimum delays between them. Burden ranges between 60% and 80% of the burden used in the production blast, while recommended spacing is about 60% to 80% of the cushion burden. In the presence of weathered conditions at the top of the rock formation or of overall poor ground conditions, guide holes can be placed between the cushion holes to enhance results of this technique. These guide holes need to be drilled only to the depth where bad rock conditions exist (ISEE Blaster s Handbook, 1998). Cushion blasting frequently uses larger diameters than presplitting, therefore, less deviation is produced and consequently deeper holes can be drilled without sacrificing alignment. In addition and contrary to presplitting, cushion blasting permits the evaluation of results immediately after the blast takes place. The main disadvantage relates to the time demanding loading and stemming requirements, which delay production and increase costs. Design guidelines for Cushion blasting were suggested by different authors, including those by Hemphill (1981) and Canmet (1977). 45

64 Smooth Blasting Smooth blasting is a wall control technique used primarily in underground stopes, tunnels and headings. It is the underground equivalent of cushion blasting, the main differences being that smooth holes are stemmed only at the collar (not the entire length as for cushion holes) and that they are fired with the rest of the round, avoiding face clean up as is the case of cushion blasts where the main blast needs to be excavated first. Smooth holes are always fired last in the sequence, usually right after the lifters. Smooth blasting in heading rounds is accomplished by drilling the holes at the perimeter of the excavation and loading them lightly with well distributed explosive charges. Suggested values (Chiappetta, 1994) for hole spacing for the smooth row are around 80% of the burden, that is S/B<0.8. A burden to spacing ratio of 1.5:1 is recommended as a starting point, which then is adjusted as needed, to account for particular rock conditions (ISEE Blaster s Handbook, 1998). Recommended Smooth blasting parameters (Burden, spacing and charge concentration) are summarized by Bhandari (1997) Air Decking The initial concepts of air decking as applied to mining were proposed by Melnikov et al (1972) in Russia. The technique consists of loading a relatively short column of explosive and creating an air chamber above this column by means of inflatable gas bags, stemming plugs or any device capable of supporting the weight of the stemming material placed above. Upon detonation of the explosive charge, an air shock will propagate upwards in the chamber until it reaches the bottom of the stemming material. 46

65 At this point the shock is reflected back into the detonation products creating a source of secondary stresses which will enhance the stress field and further extend the fracture network before gas pressurization takes place. Depending on the length of the air chamber, the process is capable of repeating itself, creating multiple loadings that will improve fracturing. Air decking provides for the expansion of the detonation gases in the borehole, thus, a reduction of the borehole pressure. However, by inducing multiple stress loadings in the rock, the duration for which the loads are applied is increased and consequently, the resulting fragmentation enhanced. Air decking can be considered as an energy accumulator, storing energy and releasing stress waves that will reduce damage arising from excessive crushing. Laboratory research conducted by Fourney et al (1981) using Plexiglas blocks demonstrated that upon reflection on the stemming material, the stress field is enhanced and the duration which it acts upon the borehole wall is increased by a factor of 2 to 5 times. Moreover, experimental observations also indicated that the highest stress field (other than the one in the vicinity of the explosive) was occurring at the stemming/airdeck interface and into the actual stemming itself. This is an indication that an air deck located at the top of the borehole should induce good fragmentation at the collar. Air deck length plays a critical role on the fragmentation results. Moxon et al (1991) extended Fourney s observations and proposed the following equation to estimate the top column air deck length: 47

66 P Pl P d = l L deck 1/ 3 /W Equation 3.5 where Pd is the detonation pressure, Pl the limiting pressure required to induce fragmentation and W the explosive weight. When applied to ANFO (Pd ~ 2700 MPa) on a rock with limiting pressure Pl = 100 MPa, Equation 3.5 can be used to solve for the airdeck length. This results in: L deck 1/ 3 2 1/ 3 = 2.6 W = 2.6 ( π re Lt ρe ) Equation 3.6 where re is the radius of the explosive charge, Lt is the total length (charge and air deck) and ρe the explosive density. Equation 3.6 will allow the estimation of a critical air deck length beyond which little additional fragmentation is obtained and the degree of fracturing will start to decrease. According to Moxon et al, the critical air deck length as estimated by the previous formula, should be reduced in the presence of a competent rock and increased 30% to 50% for softer rock formations. Compared to top column air decks, the stress fields created by mid column air decks will be further enhanced due to the reinforcing action of the shock from the bottom explosive column. Therefore, for a given air gap size, mid column air decks will produce more fragmentation in the vicinity of the air gap than a top column air deck. 48

67 3.3.2 Particle Velocity Damage Criteria Design Methods Two similar design approaches can be placed within this group; the Swedish approach as developed by Holmberg and Persson (1978) and the one proposed by Hustrulid (1999) from the Colorado School of Mines or CSM for short. Both design methodologies rely on particle velocity as the parameter that is most representative of damage. A brief description of differences and similarities of both methods is presented next Swedish Approach Holmberg and Persson (1978) proposed an empirical model relating peak particle velocity (PPV) to damage for a given explosive unit load (q in kg/m) and rock type. The model calculates the PPV induced by an explosive column at a distance R from the longitudinal axis of the charge; the results are then compared to a particle velocity damage criterion in order to establish the associated damage. As illustrated in Figure 3.1, the explosive column is divided into elemental lengths (ΔZ), each having a weight W = q * ΔZ, with q representing the load per unit length. These elemental cylinders can be considered as spherical charges, thus, the general attenuation relationship between particle velocity, explosive weight and distance for spherical charges will apply, that is: α W PPV = K Equation 3.7 β R 49

68 where K, α and β are field constants related to the rock and the explosive; W is the elemental weight; and R the distance from the center of mass of one element to the point of measurement. By integrating Equation 3.7 over the length of the charge, the total PPV contribution at the point of interest (r0, z0) is given by Equation 3.8: PPV H z = K q 2 2 [( r ) + ( z z ) ] β 0 2 α 0 0 α Equation 3.8 R zo W = q*δz ro (z zo) [ro 2 + (z zo) 2 ] 1/2 z H ΔZ Z Figure 3.1: Notation used for the calculation of particle velocity using Holmberg s approach Table 3.1 shows critical vibration levels associated with damage for different rock types as provided by Holmberg and Persson. These values define a particle velocity damage 50

69 criterion; thus, for a given rock type, vibration levels above the critical PPV will induce damage to the rock. Table 3.1: Peak particle velocity damage thresholds Rock Type Critical PPV (mm/sec) Hard rock/strong joints 1000 Medium hard rocks/no weak joints Soft rocks/weak joints <400 Based on extended blast field data generated by the US Bureau of Mines in the early 1970 s, Lundberg et al (1978) determined the constants K, α and β of the attenuation relation by performing a linear fit to the experimental data in a log PPV vs. log R/W 0.43 scale. Based on Lundberg et al s work, Holmberg and Persson assigned the following values to these constants under hard rock conditions: K = 700 mm/s α = 0.7 β = 1.5 This engineering model is often used as a predictive tool to assess extension of damage; however, it suffers from several shortcomings. The model does not account for a finite VOD, thus, vibration peaks arrive at the point of interest simultaneously. In addition, the direction of arrival is not considered. The model also ignores the effect of reflections at free faces and does not consider the effect of loading density of explosive (decoupling, decks) nor the type of explosive used 51

70 Therefore, the Holmberg and Persson method neglects the differences in arrival times of the vibration wavelets traveling from each discrete element to the measuring point. Moreover, it assumes that the peak amplitude generated by the elemental weights do not depend on the direction of travel. The effect of decoupling on PPV was evaluated by Atchison et al (1964) who established a correlation between the peak particle velocity generated by a fully coupled blasthole (PPVc) and a decoupled blasthole (PPVd). This is given by: PPV d 1.5 = PPVc ( re / rb ) Equation 3.9 where r / r represents the ratio of explosive to borehole diameters. The previous e b equation can be used to account for the effect of decoupling in Holmberg s equation CSM Approach To overcome Holmberg s shortcomings, Hustrulid (1997) proposed a predictive model based on theoretical work conducted by Favreau(1969) and later used by Harries (1983). Favreau (1969) described the particle velocity (PV) history at any given point in space, generated by a single spherical charge detonating in a homogeneous, isotropic infinite medium. His solution takes into account the change in borehole pressure from the explosion pressure (Pe) down to the equilibrium or semi static pressure (Peq) reached as a consequence of the expansion of the hole. 52

71 Favreau s solution was extended by Harries (1983) to produce an analytical model for cylindrical charges that models the explosive column as a stack of equivalent spherical charges. In order to keep both the stack of spherical charges and the cylindrical charge the same length, the radius b of these equivalent spheres was selected such that its volume equals the volume of a cylinder whose length is the diameter of the sphere. If b is the radius of the spheres and rc the radius of the cylindrical charge, for the previous to occur b = rc. Building on work by Favreau and Harries, Hustrulid proposed a PV predictive equation that not only includes Peq, the equilibrium borehole pressure, but a damping factor that accounts for the energy losses due to inelastic mechanisms, in other words, for attenuation effects beyond the usual geometric spreading. The final expression follows: PPV = r P c eq e I ( R b) / ρ c R r Equation 3.10 where Peq is the equilibrium borehole pressure, b the radius of the equivalent spheres, I is the rock damping factor, c the elastic wave velocity in the rock and R the distance from the charge to the point of interest. In Hustrulid s equation, the maximum PV value (i.e. PPV) is generated by the sphere closest to the gauge point, the others having no influence; thus, unlike Holmberg and Persson s equation, the charge length does not influence the PPV level. 53

72 By taking into account the geometry of the blast, the effect of rock type and timing considerations upon the peak particle velocity, the model showed that the latter does not increase when the charge length is increased beyond 16 blasthole diameters. It is worth noting that in addition to those factors affecting borehole pressure (i.e. explosive type, loading practice, etc.), the damage radius around a borehole is strongly influenced by both the in situ strength characteristics of the rock and the nature, frequency and orientation of geological features. In addition, borehole diameter and timing between holes also influence the resulting damage radius: with the remaining conditions kept constant, the larger the diameter the greater the damage. Similarly, with other conditions unchanged, simultaneous firing of holes produces less damage than delayed firing Innovative Contributions to Wall Control In addition to the methods described before, there have been a number of other contributions towards the creation of a clean fractured wall worth mentioning. The most relevant are discussed in the next paragraphs. Slotted explosive cartridges (Wei et al, 1987), consist of an explosive formulation encased in a plastic tube where two slots have been cut parallel to the longitudinal axis of the explosive tube. The wall thickness of the slotted tube and the width of the actual slot ranged between 2 mm and 5 mm. The slot enhances fracture creation in a given direction by focusing the energy of the reaction products into the rock. 54

73 B Gel (Lownds, 1999), is a non explosive product developed to substitute water as the shock transfer medium of choice in dimensional stone applications. The noncombustible, non explosive gel is claimed to absorb a large part of the shock energy delivered by the explosive, most typically detonating cord, reducing undesirable micro fractures and increasing the quality of a rock slab The attenuation effect that foamed materials have on the shock energy of an explosive has also been evaluated (Torrance et al, 1987). These materials include shaving cream and fire fighting foams as well as some particulate foam obtained by blending a liquid foam and talc. In addition, solid foams consisting of a two part polyurethane resin as well as expanded polystyrene beads and vermiculite particles were also assessed Specially designed linear shaped charges constructed of aluminum and filled with extruded high explosive (Composition B) were evaluated during different research projects (Rustan, 1983; Bjarnholt et al, 1983). When applied in granite, this technique generated notches between 15 to 20 mm deep when a standoff of 35 mm was used. Although the technique proved quite successful, it was costly to manufacture and implement Water jet cutting technology, used to generate longitudinal notches along the borehole (Hoshino et al, 1980). These notches were cut after drilling the blastholes. They were oriented in the direction of split and required at least three up and down passes, at a rate of 13 cm/s, to produce a 20 mm deep notch. 55

74 Experimental double and single pass drilling and notching mechanisms capable of generating a longitudinal notch to promote stress concentration and fracture propagation in the direction of the shear line has been also evaluated by Fourney et al (1977) and Holloway et al (1987). One of the difficulties experienced during development trials was the tendency of the notching tool to rotate, which made it difficult to keep the notches aligned as drilling progressed. Recent analytical and numerical fracture mechanics studies conducted by the Julius Kruttschnitt Mineral Research Center in Australia (Djordjevic et al, 2006) evaluated the effect of blasthole notching in minimizing damage. Results concluded that decoupled charges on notched boreholes minimized damage and reduced the intensity of the gas pressure required to propagate the crack, however, the direction of propagation was uncontrolled and the length of the useful cracks were drastically reduced. According to their study, the best outcome maximizing length and direction of propagation was obtained by means of notched blastholes in combination with a decoupled low VOD (i.e. low density) explosive charge. 3.4 Chapter Conclusions Different schools of thought regarding fracture mechanisms have been outlined; existing wall control techniques, including empirical approaches and innovative methods, described and the engineering models currently available to predict damage and design blasts presented. An issue that stands out is the key role that density plays on detonation, explosion and borehole pressures, therefore on reducing and controlling 56

75 rock damage. In addition, the use of low density bulk loaded explosive products stands out as an important contribution to current field methods used to reduce explosive pressure, since not only do they address damage but they also improve loading times, labor requirements and in some instances fragmentation results. 57

76 Chapter 4. Development of the Low Density Reactive Agent 4.1 Introduction This is one of the core chapters of the present research and deals with the actual development of the novel low density explosive mixture. As mentioned in an earlier chapter, in the context of this research thesis and to differentiate it from low density explosives in general, the new product will be referred to as Low Density Reactive Agent or its acronym LDRA. The reason behind this naming will become apparent in the coming paragraphs. One of the key factors considered when developing the LDRA was focussed in controlling the segregation inherent in mixtures diluted with bulking agents, in particular for diluted ANFO dry mixtures. In more instances than not, this effort has been rewarded with limited success due to the natural tendency of the lighter particles to segregate as a consequence of the extremely different bulk densities among ingredients. Diluting ANFO with fillers like polystyrene, vermiculite or perlite will be a good case example. If we evaluate the issue of segregation from a different perspective, looking not at whether it is going to occur but at the level at which it will occur, two possible scenarios arise. First, the degree of segregation upon mixing and loading is such that an uneven density distribution along the length of the borehole will exist, resulting in an inefficient use of the explosive energy. Second, the degree of segregation is such that the explosive 58

77 column fails to initiate or propagate reliably. From the two alternatives mentioned, it is obvious that the latter cannot be afforded and every effort should be done to prevent this from happening. Starting from the premise that segregation cannot be eradicated in dry ANFO mixtures, focus turned to eliminating the risk of detonation and propagation failure arising from either, a high dilution or severe segregation. To address such an objective, it was decided to provide a surface treatment to a bulking agent to allow adhesion of an energetic ingredient that will make the mixture detonable while still maintaining its bulking characteristics. In other words: to develop a reactive bulking agent capable of sustaining stable detonation at extremely low densities. The reactive nature of the LDRA at densities comparable to those of typical bulking agents becomes then the most important and distinctive characteristic of the product. To achieve this objective, the density of the selected bulking agent should be low enough for the resulting LDRA mixture to maintain its bulking characteristics once the explosive ingredient is added. Detonation of the LDRA product should be a self propagating and stable event whether or not it is mixed with other explosive compositions such as standard ANFO prills or emulsion explosives. As will be described in the next section, the extremely low density of expanded polystyrene was the most important factor behind its selection as the bulking ingredient of choice. On the other hand, the selection of ammonium nitrate as the energetic ingredient for the LDRA was a straight forward choice since its price, availability, 59

78 handling and safety characteristics cannot be challenged by any other oxygen providing chemical and much less by any high explosive compound. With the previous concept in mind, if the need comes to mix LDRA and ANFO, the reactive nature of the former will render the probability of initiation and or propagation failure of the resulting mixture negligible regardless of the degree of dilution sought or the level of segregation produced upon mixing or loading. In addition, the LDRA ingredients and preparation methods need to be simple enough to present themselves as an attractive economic alternative to existing wall control products. In summary, several characteristics were sought during the developmental stage of the LDRA. These include a product with a reactive nature at low densities, amenable to bulk loading operations, containing ingredients that are safe to handle, inexpensive to purchase and readily available in the industry and a manufacturing process that does not require costly equipment. In addition, the LDRA ingredients should be easy to mix and chemically compatible with typical commercial explosive products, and preferably have a favorable classification for transportation and packaging purposes. 60

79 4.2 Selection of Bulking Agents Due to their low density nature, all bulking agents suffered from one major drawback: the high incidence of transportation costs on its price, however, consideration was given to products with potential to improve the commercial viability of the LDRA. Factors such as high loading rates, safe handling and transport characteristics, segregating behavior, availability and price were considered. Different bulking agents were considered for providing the low density structure to the LDRA, including those composed of absorbent and non absorbent material. From all of them, expanded polystyrene was the selected bulking agent and will be addressed in more detail within this section. The remaining alternatives, although briefly introduced in the next paragraph, are explained in more detail in Appendix 4.1. Nanogel, consists of micron range translucent, irregularly shaped aerogel particles prepared by treating amorphous silica gel with an organo silicon compound. Expancel, are small spherical plastic particles consisting of a polymer shell encapsulating a gas. Cellulose Granules (K Sorb), is a particulate product manufactured from recycled cellulose capable of retaining more than five times its weight in liquid. UltraFoam, is a state of the art reticulated vitreous carbon rigid foam obtained by pyrolysis of a polymeric material. Perlite, refers to a siliceous rock containing interstitial water that when subjected to a thermal shock softens and expands several times its original volume, becoming very light and absorbent. 61

80 4.2.1 Expanded Polystyrene (EPS) Beads The manufacturing of expandable polystyrene is presently a one step process consisting of the suspension polymerization of a styrene monomer in water simultaneous with the impregnation of the suspension with a foaming agent. The foaming agent consists of a low boiling temperature liquid (25 C 45 C), usually of the pentane family, which will impregnate and remain entrapped within the unexpanded polystyrene beads, as long as it is not subjected to temperatures above their boiling point. The actual expansion of the unexpanded beads is carried out by the application of heat in the form of water vapor, which softens the polymeric structure of the beads at the same time that it gasifies the low boiling point liquid. As a result, a substantial increase in volume, i.e., density reduction, will occur along with the formation of the closed cell arrangement that characterizes the expanded polystyrene beads. Final EPS density will depend on the amount of foaming agent present in the unexpanded bead, the initial size of the beads, the presence of certain additives, the time the beads are subjected to vapor heat as well as the temperature of the vapor itself. Typical densities range between 0.01 g/cm 3 and 0.05 g/cm 3, with diameters ranging between 1 10 mm. The air filled polyhedral closed cell arrangement within the bead is as small as 45 μm in diameter and 1 μm in wall thickness. The air in these closed cells accounts for 96% to 98% of the total volume of the bead, with the polymer walls accounting for the remaining part. 62

81 EPS is neither absorbent nor hygroscopic and although impervious to liquid water, it is permeable to water vapor. Low temperatures have no adverse effects on EPS; however, at temperatures above 100 C it will start to soften and experience damage to its structure. EPS is a comparatively inert product not affected by mildly acid or alkaline materials. However, it is soluble in many organic solvents such as acetone, ethyl acetate, xylene, paint thinners, carbon tetrachloride and turpentine as well as saturated hydrocarbons like kerosene, gasoline, fuel oil, diesel oil and others. Moreover, certain commonly used paints, adhesives, wood preservatives, pasting agents, etc. may also damage EPS. In some cases, even vapor from these substances can be harmful to its physical structure. The detrimental effect of the saturated hydrocarbons on the physical integrity of the EPS beads is one of the most important factors to consider when dealing with polystyrene/anfo mixes. Given its extremely low density, worldwide availability and low volumetric cost, expanded polystyrene has been used extensively as filler material to dilute ANFO. However, segregation and handling issues, like its tendency to cling due to static buildup and its low density transport characteristics, tend to outweigh potential benefits of an otherwise fairly successful mixture. An important step forward to making EPS the bulking agent of choice for diluting ANFO would be the introduction of inexpensive process technology to allow for on site expansion of the beads. This will make the use of EPS very attractive since it will reduce the costs inherent in transporting a product whose volume is composed of more than 90% air. 63

82 From the above choices, perlite and expanded polystyrene (EPS) were selected for preliminary trials. Low density explosive mixtures were manufactured and detonated using either one of the two choices; however, the final decision was taken in favour of EPS for three basic reasons. The first reason is that the density of EPS can be chosen to be an order of magnitude lower than the one corresponding to Perlite, providing a wider margin for developing an LDRA with low density characteristics. The second reason relates to the difficulties associated with providing surface treatment to absorbent materials such as perlite, vermiculite and cellulose granules, which will permit adhesion of the energetic ingredient. Third and final reason has to do with the greater potential that EPS has of being expanded on site due to process requirements. This is an important consideration given the high incidence that transportation costs have on the price of low density products in general. 4.3 Developmental Stages of the LDRA This section describes the experimental steps that led to the development of the final LDRA mixture. The initial interest on low density explosives containing expanded polystyrene as bulking agent was triggered by Schvedov et al (1984) with the introduction of technology capable of expanding the polystyrene beads while simultaneously applying a saline layer of ammonium nitrate crystals. 64

83 As opposed to current technology where expansion of the polystyrene beads is achieved by application of vapor heat followed by a drying and settling stage, Russian technology made use of a 5 to 10% weight of ammonium nitrate solution preheated to a temperature ranging between 100 C and 120 C. The unexpanded polystyrene beads are first fed into the hot solution tank via an auger screw or similar system. Upon contact with the hot solution, the polystyrene beads will expand as a result of the combined action of vaporization of the entrained pentane and softening of the polystyrene structure itself. As expansion of the polystyrene beads takes place, they will float to the surface where they are skimmed out and covered with a layer of the hot saline solution. Once this layer cools and dries, ammonium nitrate crystals will precipitate and loosely adhere to the surface of the expanded beads. Critical to this technology was not only the quantity but also the strength of the adhesion of these crystals to the surface of the expanded beads. In order to validate its viability, the Russian approach was duplicated in the laboratory using ammonium nitrate solutions at different concentrations. Following is a summary of the preliminary research steps that led to the development of the final LDRA composition Laboratory Tests with Hot Solutions of Ammonium Nitrate The very first step for duplicating the Russian process was to obtain unexpanded polystyrene beads. Various manufacturers including BASF, StyroChem and Kumho were approached and samples of unexpanded beads having different blowing agent contents were delivered for testing. 65

84 The unsaturated hot solution of ammonium nitrate was prepared at a higher concentration than the 5% 10% used by Russian experimentalists to ensure enough crystals would be present upon cooling and drying. A small glass reactor, equipped with a stirrer and a heating device, was used for dissolving the ammonium nitrate into solution. The temperature of the solution was set at 85 C, high enough to induce vaporization of the blowing agent and softening of the polystyrene structure, both of which are necessary for a proper expansion of the bead. The solubility of ammonium nitrate in water at this temperature is in the order of 87% (87 g of salt in 100 g of solution), as reported by Wang (1994). Unsaturated solutions in the order of 70% were prepared, initially by dissolving ammonium nitrate explosive grade prills and later using high purity ammonium nitrate Stengel crystals, in an attempt to reduce the impurities generated by the surfactant present in the prills. A small sample of unexpanded polystyrene beads was placed in the reactor and allowed to sink to the bottom of the aqueous solution. After a few seconds in contact with the hot solution, the expansion process got under way and the beads rose to the surface where they were manually skimmed and let dry. Visual inspection of the expanded beads showed the ammonium nitrate crystals loosely adhered to their surface. The density of the mixture was in the order of 0.04 g/cm 3. With careful handling to prevent excessive segregation, samples were analyzed to determine the amount of salt adhered to the surface. This was achieved by dissolving the polystyrene beads in toluene and weighing the remaining salt after filtering and 66

85 drying. Results indicated salt adhesion in the order of 50% by weight. However, the adhesion strength was weak, with most of the crystals falling off easily upon gentle shaking. It became apparent that a way to improve the strength of adhesion was required; otherwise, excessive segregation of the ammonium nitrate crystals upon activities such as handling, transportation and loading could result in an unreliable product. Several avenues were investigated to improve adhesion strength, these included wetting expanded polystyrene in a 10% solution of hydrochloric acid to induce surface ionization prior to placing them in the hot solution of ammonium nitrate, wetting expanded polystyrene in surfactant solutions prior to immersion in the ammonium nitrate solution. Three different surfactants were used: sodium dioctyl sulfosuccinate, sodium diamyl sulfosuccinate and sodium dodecyl sulfate. In addition, previous experiments were duplicated but dissolving pure ammonium nitrate crystals (Stengel crystals) rather than explosive grade prills, in an attempt to evaluate the effect that impurities may have on adhesion strength. Adhesion strength did not show a noticeable improvement using the previous concepts. Segregation of the ammonium nitrate crystals from the surface of the expanded polystyrene beads was still present upon gentle shaking of the mixture, particularly once the solution was dry and the wet layer had disappeared. In addition, the above experiments involved a series of steps that made the laboratory scale manufacturing process time consuming; much more so if extrapolated for a larger, 67

86 commercial scale manufacturing. It was evident that a simpler approach to strengthened adhesion and reduce crystal segregation from the polystyrene beads was required for the product to have a chance of succeeding at an industrial scale Experiments Using Mineral Oil as Surfactant As experienced throughout the previous tests, handling and manipulation of expanded polystyrene beads tend to produce static charges that make them difficult to work with. To get rid of the static clinging, it is standard in the industry to apply a thin coat of mineral oil to the surface of the beads. The extremely high surface affinity that exists between oil and polystyrene requires a small oil percentage (~5% by weight) for that static activity to disappear and the expanded polystyrene beads to settle. More importantly, the coat of oil applied to the expanded beads will not evaporate and dry out but remains indefinitely adhered onto their surfaces, thus providing a wet layer for the ammonium nitrate crystals to hold to. This simple discovery turned out to be a breakthrough in the manufacturing of the final LDRA mixture; however, care should be taken to ensure paraffin or vegetable based oils are used since polystyrene is not chemically compatible with naphtene based products such as gas oil, fuel oil and kerosene. Other lubricant or cooking oils, including motor oils, chain oil, sunflower oil and corn oil, are appropriate alternatives to mineral oil. Mineral oil was readily available and selected for much of the experimentation conducted throughout this research. 68

87 When compared to the laboratory approach simulating the Russian experience, wetting the EPS with mineral oil will have several advantages. To start, the expansion of the polystyrene beads would not be part of the preparation process, thus, beads expanded elsewhere (preferably on site) could be used. There is no need for hot solutions of ammonium nitrate, since crystals can be deposited over the layer of mineral oil, which have a strong surface affinity with polystyrene and does not evaporate, thus, providing a permanent wetted surface for the adherence of crystals. Moreover, the use of mineral oil precludes the use of surfactants to precondition the polystyrene surface, thus, no need for filtering and drying would be necessary. In addition, oil is chemically compatible with both polystyrene and ammonium nitrate crystals, inexpensive and readily available in the industry. In an attempt to evaluate the influence that oil content has on the adhesion of ammonium nitrate fines to the surface of the polystyrene beads, a few simple laboratory experiments were conducted by wetting the beads with a given weight % of mineral oil and then mixing them with ammonium nitrate fines. Adhesion of the powder was measured by weight difference, after sieving the loose fines and using toluene to dissolve the polystyrene. Two different expanded polystyrene beads were used during these experiments: 1 mm and 2 mm diameter beads. The decision behind this selection was based on work conducted by Heltzen (1980), where he concluded that segregation between ANFO prills and EPS beads was reduced when similar particle size diameters were mixed. Given that 69

88 one of the applications foreseen for the LDRA is as a diluting agent of ANFO, polystyrene beads with diameters similar to those of ANFO prills were consequently chosen for evaluation. Results presented in Table 4.1 show an increase in adhesion, and consequently mixture density, as the percentage of applied mineral oil increases; however, this seems to reach a plateau beyond which a further increase in oil content does not translate into an increase in powder adhesion. In other words, salt adhesion may have reached its maximum for the given bead diameter. Table 4.1: Laboratory results showing the effect of oil content on adhesion of ammonium nitrate fines and mixture density for different EPS bead diameters EPS Diameter EPS Density Oil Content Crystal Adhesion Density Mix Specific Surface Constant (mm) (g/cm 3 ) (% wt) (% wt) (g/cm 3 ) (cm 2 /g) K It is worth noting that the trend between bead diameter and crystal adhesion presented in Table 4.1 is apparent in nature. The maximum adhesion value should be a sole function of the specific surface (SS in cm 2 /g) of the beads, which in turn is an inverse function of two parameters, bead diameter and bead density; in other words: 2 1 Adhesion SS( cm / g) = Equation ρ( g / cm ) φ( cm) 70

89 Therefore, an approximately constant relationship (K) should be expected between specific surface (SS) and measured adhesion percentage. Although the few tests conducted were not sufficient to establish a statistically significant relationship, calculations shown in the last column of Table 4.1 seemed to indicate a trend towards a constant K value for both bead diameters as oil content is increased. Two preliminary conclusions are obtained from the latter tests. First, there is a strong indication that the 2 mm diameter, g/cm 3 density expanded beads would reach an adhesion plateau at about 60% to 65% ammonium nitrate, corresponding to a mixture density in the order of 0.05 g/cm 3. Therefore, if higher mixture densities are desired to ensure product reliability, the extra fines that need to be added will not be adhered but remain entrapped within the bead s voids. Second, a more qualitative conclusion, clearly improved adhesion strength was observed as a consequence of the oil film coating the beads. This was noticed from the reduced segregation observed during handling and shaking of samples. At this stage of the research it was decided to conduct an initial set of field tests at the Queen s University explosives test site using a low density mixture prepared by wetting EPS beads with mineral oil and mixing them with ammonium nitrate powder. The main objective of this initial phase was to observe whether the formulation would react when mixed with standard ANFO prills at various volumetric ratios and the conditions at which those reactions occur. Continuous VOD testing was conducted as performance 71

90 indicators of the reaction. Two similar low density mixtures (LDE) were prepared at a target density of 0.05 g/cm 3. These were: LDE #A: consisting of 2 mm EPS beads wetted with mineral oil (30% weight relative to EPS) and ammonium nitrate powder to a bulk density of 0.05 g/cm 3 LDE #B: consisting of 1 mm EPS beads wetted with mineral oil (30% weight relative to EPS) and ammonium nitrate powder also to a bulk density of 0.05 g/cm 3 The previous low density mixtures were in turn mixed with standard ANFO prills at decreasing densities. The initial charge consisted of a mix at 70/30 volume of LDE # A and ANFO prills loaded in a 100 mm diameter steel tube at a density of 0.28 g/cm 3. All charges were initiated using a 220 g pentolite primer. Table 4.2 next summarizes the experimental results. Table 4.2: Experiments conducted using a preliminary composition of a LDE mixture consisting on oil wetted EPS beads and ammonium nitrate powder Test # LDE Volume ratio LDE/ANFO Density (g/cm 3 ) Diameter (mm) (inch) VOD (m/s) 1 A 70/ A 80/ B 90/ B 100/ B 100/ Despite evident segregation observed between the LDE and ANFO, all the charges tested detonated. Most noticeable were the ones where no ANFO prills were used, which detonated at an extremely low density in relatively small diameters. The low 72

91 pressures generated during these tests did not properly short the VOD probes; hence, no clean electrical output was produced, which in turn resulted in VOD records difficult to interpret. The poor uniformity of the mixtures due to segregation would explain the poor VOD density trend observed in the results. It became apparent that a different approach needed to be implemented in future tests if more reliable VOD records were to be obtained. This will be presented in a later chapter of the thesis. Following these initial sets of experiments, further detonation testing was conducted using the same mixing approach to validate the results; however, these proved inconsistent, with some experiments failing to detonate altogether. The cause for these failures was attributed to a loss of sensitivity of the low density mixture as a consequence of the high hygroscopic characteristic of the ground crystals of ammonium nitrate used. Regardless of the reason, mixture reliability proved unacceptable and further investigation was warranted. At this stage of the research there were two areas targeted for improvement. First, improving adhesion strength to ensure segregation of crystals is minimized and second, improving moisture protection of the ammonium nitrate crystals in order to reduce caking and agglomeration and increase sensitivity to detonation. The larger 2 mm polystyrene beads were selected for the remainder experiments since their lower bulk density would result in lower LDRA densities and their larger diameter in less segregation when mixed with standard ANFO prills. 73

92 With the previous objectives in mind, a set of laboratory tests was conducted to look at the influence of oil viscosity on crystal adhesion by wetting the 2 mm beads with a light mineral oil and heavier chain saw oil. Depending on the oil content applied, results indicated a 5% to 15% increase of crystal adhesion when the higher viscosity oil was used. However, higher viscosity oils do not disperse as easily as the lighter ones, prompting a look for ways to improve adhesion strength without increasing viscosity. A tackifying agent available in the industry proved to be such an alternative. This additive is a solution of a copolymer of ethylene and propylene which is easily soluble in petroleum and synthetic lubricants. It is typically used to increase adherence of oils without actually increasing their viscosity. Two similar products were tested; the first one is offered under the name Vanlube TK 100, manufactured by R. T. Vanderbilt Inc. based in the United States and the second one marketed by the name HiTEC 152, manufactured by Ethyl Corporation in the United Kingdom. An oil solution containing 0.5% to 5% of the tackifying polymer in oil is the recommended range suggested by the manufacturers, beyond which no substantial increase in oil tackiness will be realized. Evaluation was conducted on the 2 mm diameter beads wetted with different contents of mineral oil tackifyer solution, in turn containing different proportions of tackifying agent. No difference in performance between the two commercial products was observed. To further improve anti caking characteristics, increase ammonium nitrate sensitivity and adhesion strength, it was decided to mix the ground AN crystals with the same 74

93 solution of mineral oil and tackifying agent at proportions corresponding to an oxygen balance composition. For the remainder of this thesis, the mix of ammonium nitrate fines and mineral oil tackifying solution will be referred to as ANMO fines or simply ANMO. As shown in Table 4.3, an even adhesion strength of ANMO fines, independent of oil content, was observed as a consequence of the tackifying agent. The adhesion value itself showed a slight improvement when compared to results of previous tests using just oil or other surfactants. Table 4.3: Adhesion of ANMO fines for different tackifying and wetting oil contents Mineral Oil on EPS (% weight) Tackifyer Content (% weight) Adhesion of fines (% weight) Density (g/cm 3 ) The protective layer provided by wetting the EPS particles with mineral oil should be sufficient to prevent chemical attack from fuel oil or diesel oil in the eventuality that these fuels were selected to substitute mineral oil for mixing with the ammonium nitrate fines. However, this would represent a risk not worth taking since, at the very least, storage life of the mixture would be reduced. Therefore, mineral or similar lubricant oils are favored over diesel oil or fuel oil for mixing with the AN fines. 75

94 4.3.3 Selection of Ammonium Nitrate Fines Much of the laboratory testing conducted used either crushed ammonium nitrate prills or crushed ammonium nitrate Stengel crystals, the latter being high purity crystals. Observations during mixture preparation brought up the long standing issue of the hygroscopic behavior of ammonium nitrate and its tendency to cake. This tendency was readily observed soon after crushing the ammonium nitrate crystals, while the crushed explosive grade prills would last longer due to the anti caking agent present on its surface. An improvement to this issue was sought in order to increase life, sensitivity and overall reliability of the final low density explosive mixture. Nitrochem, an AN manufacturing plant, was approached for consultations. It was learned that during the manufacturing process of the ammonium nitrate prills, once the anti caking surfactants had been added, there is a certain amount of fines generated that cannot be recycled into the stream due to the surfactant treatment applied. It is worth noting that, as it was learned later, the carbon content of these fines due to the presence of the surfactant was 0.4%, which is higher than the 0.2% limit required for the ammonium nitrate prills to classify as an oxidizing agent in Canada. Having 0.4% carbon content will classify the product as a high explosive rather than an oxidizer, thus losing the advantages related to handling, storing and transporting the product. It was decided to test these fines for three reasons: first, no grinding of AN prills or crystals would be required; second, the fines are better protected from humidity and 76

95 caking; and third, being a non recyclable process stream, the fines are considered a reject and could be acquired at a favorable price if industrial quantities were required. Upon visual observation and handling, the fines did not present any signs of caking and remained free flowing despite the fact that they were stockpiled in warehouses at the manufacturing plant for long periods of time. Results of a sieve analysis conducted on three small samples (~250 g) of these fines are illustrated in Table 4.4. Table 4.4: Sieve analysis of ammonium nitrate fines as supplied by Nitrochem Tyler Mesh Sieve Aperture Sample 1 Sample 2 Sample 3 # (microns) Cum % Pass Cum % Pass Cum % Pass The cumulative % by weight passing graph is illustrated in Figure 4.1, where it can be seen that over 85% of the size distribution falls within 833 and 104 microns (Tyler #20 and Tyler #150 sieves respectively). 77

96 AN Fines (Nitrochem Inc) 120 Cum Weight Passing (%) Sieve Aperture (microns) Sample1 Sample 2 Sample 3 Figure 4.1: Cumulative weight retained for ammonium nitrate fines provided by Nitrochem It was decided to conduct a few laboratory experiments to qualitatively assess the effect that the ammonium nitrate grain size has on segregation/adhesion and ease of mixing. Ammonium nitrate fines of different grain size distributions were mixed with the mineral oil solution to a balanced formulation (i.e. balanced ANMO fines) in order to optimize mixture sensitivity and increase reliability. Three samples from Nitrochem were sieved using upper cut off limits of 417 μm (Tyler #35 sieve), 590 μm (Tyler # 28 sieve) and 833 μm (Tyler # 20 sieve) respectively. An oilsoluble blue dye was added to the mineral oil solution to provide a contrast effect with the ammonium nitrate fines and help visualize quality and ease of mixing. Results were as expected, the smaller the ammonium nitrate particle size, the more difficult the mixing with the mineral oil solution becomes. Moreover, when the smaller 417 μm mesh (Tyler # 35 sieve) was used as the upper cut off for the fines, not only did it 78

97 take longer to mix them with the oil, but once mixed, the resulting ANMO fines were no longer behaving as free flowing particles. The fines would tend to clump and stick to each other, making the mixing process with the polystyrene beads more difficult. On the other hand, the larger the particle size of ANMO fines, the higher the tendency of the grains to segregate from the surface or the void spaces of the expanded polystyrene beads. This effect was best illustrated by a drop in density from 0.14 g/cm 3 to 0.08 g/cm 3 when the smaller 417 μm mesh (Tyler # 35 sieve) was changed to the larger 833 μm mesh (Tyler # 20 sieve). It became apparent that a compromise should be reached to settle these two opposing behaviors by selecting a particle size having an upper limit that would reduce segregation of ANMO fines and a lower limit to prevent excessive ultra fine material that will prolong mixing time with the mineral oil solution. From them both, segregation was considered to have priority over ease of mixing, since the latter can be controlled by increasing the mixing time. Upon several observations and density measurements, it was concluded that for the 2 mm diameter polystyrene beads and the ammonium nitrate fines as supplied by Nitrochem, an upper limit of 589 μm (Tyler # 28 sieve) and a lower limit of 104 μm (Tyler # 150 sieve) proved adequate to achieve the mentioned goals. However, if a different source of fines were used, such as grinding ammonium nitrate prills, care should be given to ensure particles were evenly distributed throughout the suggested 79

98 range to avoid a high percentage of ultra fine particles that will lengthen the mixing process. 4.4 LDRA: Final Mixture Following on the previous findings, a second experimental phase was set at the Queen s testing range to evaluate the detonation behaviour of the LDRA mixture, this time introducing three basic improvements relative to the initial experimental phase implemented earlier. These were: 1. The use of ammonium nitrate fines sourced from the reject stream of the manufacturing process of AN prills, as opposed to the ground crystals or crushed prills used during the initial phase. The surfactant present in these free flowing fines increases their anti caking characteristics, reduces mixing time with mineral oil and extends overall life of the mixture 2. The use of oxygen balanced ANMO fines instead of just ammonium nitrate fines, in order to increase sensitivity to initiation and stability of propagation while reducing the hygroscopic effect and extending the life of the LDRA mix 3. The use of a tackifyer additive in the oil in order to increase adhesion and cohesive strength of ANMO fines without increasing viscosity The implementation of the above changes in the trials was successful, generating reliable initiation and stable propagation of the LDRA at densities above 0.10 g/cm 3 ; however, unstable detonation regimes, manifesting themselves in the form of bulges along the 80

99 steel tube, were observed at densities below the mentioned value. Performance results will be discussed in more detail in the next chapter. From the series of laboratory experiments and the experience gained during field testing, several useful comments and recommendations regarding the LDRA final mixture, including ingredient composition, preparation stages and mixture density, are described in more detail in Appendix 4.2. Nevertheless, two relevant topics will be addressed in this particular section: oxygen balance and water resistance Oxygen Balance The ingredient composition proposed for the final LDRA results in a mixture with a markedly negative oxygen balance (OB), in other words, a fuel rich mixture that will tend to generate excessive carbonaceous products including CO, CO2 and methane gas as well as elemental carbon. This characteristic will preclude the use of the product in underground mining operations unless proper ventilation conditions and environmental regulations allow for such product to be applied. As a baseline, the OB of an LDRA mixture at 0.15 g/cm 3, which contains approximately 84% of ammonium nitrate, 6% of oil solution and 10% of expanded polystyrene by weight, is in the order of 345 g O per kg of mixture. For this proposed composition, there is no room for substantial improvements to the OB other than substituting the polystyrene itself, which for all practical purposes will result in a new explosive composition. Reducing the oil content of the ANMO fines will have some effect on OB; however, it was decided to mix ANMO to a balanced composition to ensure optimum 81

100 sensitivity and propagation stability. These two important explosive properties cannot be guaranteed if a poor quality fuel oxidizer mix is produced by fuelling the ammonium nitrate fines with the large polystyrene particles. The influence of polystyrene in the OB of the composition is just too large for the simpler, most obvious changes to have a practical effect. Reducing mineral oil content on the beads, or even in the ammonium nitrate fines, will improve the OB but to a negligible extent, however, it could greatly affect initiation and performance of the explosive mix. The addition of sodium nitrate, potassium nitrate, sodium perchlorate or similar oxygen provider salts will also improve the balance but marginally due to the small amounts that these salts can be used to replace ammonium nitrate. As an example, if we were to substitute 20% of the ammonium nitrate by sodium nitrate, whose balance in oxygen by weight more than doubles the former, the OB of the resulting composition will be around 295 g O per kg of mixture. It is worth noting that the above changes will affect performance and cost of the explosive product and in some cases, such as with ammonium perchlorate, it could also affect sensitivity and in turn raise environmental and safety concerns. The partial substitution of expanded polystyrene for low density inert materials such as perlite or vermiculite is also an alternative available to improve OB; however, the extent to which this can be done without the mixture failing to detonate should be given utmost consideration. As an example, if the LDRA used as baseline was to be diluted to 82

101 a 50/50 volume ratio with an inert ingredient such as perlite, the corresponding OB for this mixture will be in the order of 210 g O per kg of mix, still a fairly negative value. In conclusion: the LDRA composition as developed does not lend itself for underground applications. Although improvements can be incorporated into the composition in the form of addition of richer oxygen providers or substitution of fuels by inert ingredients, these cannot be done to the proportions needed for balancing the ingredients. Moreover, eventual changes introduced in the LDRA composition will affect the performance behaviour of the resulting mixture to a point that could lead to initiation and or propagation failures Water Resistance Like standard ANFO prills, the LDRA product should be treated as a dry blasting agent with no resistance to water. There is, however, an important difference between both products in the presence of a blasthole containing water. Due to its high solubility, a large volume of ANFO prills will soak water, sink to the bottom and dissolve in a short time span. The LDRA on the other hand, will float on top of the water surface, producing a plug which will act as a barrier and protect the remaining column length, however, a slow desensitization process is expected to occur as water rises through the ANMO fines by capillary action. 83

102 4.5 Potential Improvements for Field Applications Low density explosives formulated around any bulking agent will suffer from a major disadvantage as far as their manufacturing cost is concerned; that is, the strong influence that transportation has on the price of these low density consumables. Discussions on a viable approach to address the transportation problem, environmental issues originating from the expansion process of polystyrene beads and considerations regarding mechanized loading alternatives are discussed in Appendix Chapter Summary The present chapter details the steps involved in the development of the LDRA, from inception to final product, listing the various low density materials considered as potential bulking agents and discussing the characteristics of expanded polystyrene as the alternative of choice. Oxygen balance and water resistance characteristics of the final LDRA mix are also discussed. In the appendices associated with the present chapter, further information on the various bulking agents considered is provided, along with general suggestions regarding the final LDRA explosive. The latter include ingredients and mixture preparation as well as those conditions limiting the use of the product. In addition, recommendations regarding polystyrene expansion issues and mechanized loading alternatives, which may have an important impact in the viability of the LDRA of becoming a commercial product, are also addressed in the appendix section. 84

103 Chapter 5. Characterization of the LDRA 5.1 Introduction This chapter focuses on the experimental evaluation of the low density reactive agent (LDRA) as developed in the previous section. Particular experimental emphasis was placed on characterizing those properties considered relevant to wall control, the main application foreseen for the LDRA mixture. Velocity of detonation, detonation pressure and explosion pressure tests comprised the core of the experimental research. VOD testing was primarily conducted to evaluate the effect of diameter, primer and confinement in an attempt to better understand explosive behavior at density levels that have been shown to generate unstable detonations. In addition, given the important role detonation and explosion pressures play in the performance of explosive products designed for wall control, a sustained experimental effort was also focused on these two parameters. Experiments to observe the stability of propagation in longer column charges, the response of the LDRA in a decking configuration and the ability of the LDRA to initiate and be initiated by a column of ANFO were also evaluated via VOD tests. In addition, mixtures of the LDRA with ammonium nitrate prills as well as ANFO prills at different volume ratios were evaluated. The low pressure amplitudes characterizing the LDRA required modifications to existing techniques for recording VOD, while on the other hand, they presented the 85

104 opportunity to investigate the full history of the explosion pressure by designing experiments of a non destructive nature that allowed the repetition of tests under different loading scenarios. In addition, density measurements were conducted on the LDRA in an attempt to quantify density gradients upon column loading as well as density changes upon vibration and tapping. Similar experimentation was also conducted on mixes of LDRA ANFO in an attempt to qualitatively assess mixture segregation upon mixing and loading. 5.2 Velocity of Detonation Experiments The heterogeneous nature of the LDRA product, its tendency to settle and increase density upon shaking and or tapping, the different primer weights utilized throughout the experiments, the batch manufacturing nature of the product added to the uncertainty of interpreting some experimental records were all contributing factors to the scatter and lack of reproducibility observed on experimental results. Scatter of experimental data in absolute terms for low VOD, low pressure regimes, such as the present case, tends to be large relative to high density homogeneous products; nevertheless, the importance of this scatter has little or no practical implications in a mining scenario. In other words, assuming 50% error, a typical emulsion explosive rated at 40 Kbar and delivering 60 Kbar will induce substantial damage. On the other hand, 86

105 the damage induced by a low density product rated at 2 Kbar but delivering 3 Kbar will have little or no significance as far as damage is concerned. Only VOD data generated by the LDRA as finally developed were considered for analysis, that is 2 mm polystyrene beads wetted with 15% by weight mineral oiltackyfier solution (MO) then mixed with ANMO fines, in turn consisting of ammonium nitrate fines (AN) mixed to an oxygen balance concentration with the same MO solution. The fines were rejects from a local AN manufacturing plant. Ingredients were mixed to obtain a density of 0.15 g/cm 3. Tests in which density itself was the variable of interest were analyzed separately. The experimental records corresponding to the VOD measurements have been included in Appendix VOD Measuring Technique The high resistance Ni Cr wire technique, a continuous VOD recording method, was used in almost all the experiments, and only a few readings were obtained from discontinuous methods where the time span required for the detonation front to travel a known distance was recorded and the resulting average velocity calculated. The VOD sensors were constructed by threading a high resistance Ni Cr wire through a 1.6 mm diameter, 90 cm long brass tube and crimping both at one end to establish electrical connection, with the brass tube acting as a return lead. The initial experiments using brass sensors generated VOD traces that were discontinuous and electrically noisy due to the intermittent opening/closing of the 87

106 contact, making the records difficult to interpret as shown in Figure 5.1. This particular experiment was conducted on LDRA at a density of 0.12 g/cm 3 confined in a 75 mm steel pipe. VOD Trace using Brass Sensor m/s Distance (m) m/s 1430 m/s Time (ms) Figure 5.1: VOD record using a brass sensor on LDRA at 0.12 g/cm 3 in 75 mm steel pipe On many occasions after the experiment, it was possible to recover the brass sensor as a single piece, which although twisted and burned was not effectively crushed. This was an indication that the brass tubes were able to withstand the crushing pressure of the explosive, establishing an intermittent electrical contact with the Ni Chrome wire, and therefore producing poor quality records that were difficult to interpret. In order to reduce bias in the interpretation of VOD traces, brass tubes were later replaced by smaller diameter (0.8 mm), much softer aluminum tubes. The relative dimensions of the aluminum and brass tubes used for the VOD sensors are illustrated in Figure

107 Figure 5.2: Aluminum and brass tubes used for VOD sensors Construction of aluminum sensors proved time consuming due to the easiness by which they kinked or collapsed when slight pressure was applied during assembly, preventing threading of the Ni Cr wire within the tube. To avoid damaging the aluminum tubes, a regular brass tube was taped alongside to provide structural rigidity. As illustrated in Figure 5.3, aluminum sensors greatly improved the quality of VOD traces. 1.0 VOD Trace using Aluminum Sensor Distance (m) VOD = 1893 m/s Time (ms) Figure 5.3: VOD trace with an aluminum sensor. LDRA at 0.12 g/cm 3 in 75 mm steel pipe 89

108 5.2.2 VOD Diameter Relationship The relationship between velocity of detonation and the diameter of an explosive charge is a consequence of energy dissipation to the side of the explosive column. These energy losses are inversely proportional to the diameter, in other words, the larger the diameter the smaller the dissipative losses relative to the total energy at the wave front. Therefore, as the diameter of the explosive charge is reduced, the VOD will decrease until it reaches a point at which the dissipative forces overcome the generative forces and the detonation front fails to propagate. This point is known as critical or failure diameter. On the other hand, as the diameter of the charge is increased, the VOD will increase until it reaches a plateau where it becomes independent of diameter and steady state conditions prevail; in other words, it is a state representative of stable chemical reaction for the given ingredient composition at infinite diameter. This typical VODdiameter trend was observed in the LDRA product. Due to environmental restrictions at the Queen s testing facility limiting the size of the explosive charge, the largest experiment conducted consisted of a 200 mm diameter, 1.5 m long steel pipe containing about 7 kg of LDRA. The core of the experiments however, was conducted in 50 mm and 75 mm diameter, 1 m long steel pipes. Figure 5.4 illustrates VOD as a function of charge diameter for the LDRA at 0.15 g/cm 3 under steel confinement. 90

109 Diameter Effect 0.15 g/cc, steel) VOD (m/s) Diameter (mm) Figure 5.4: VOD Diameter relationship. LDRA at 0.15 g/cm 3 under steel confinement Figure 5.5 graphs the average VOD as a function of diameter, showing the error bars corresponding to the minimum and maximum uncertainty values. As previously mentioned, dispersion of VOD and pressure data is to be expected from heterogeneous mixtures such as the LDRA, where up to a 20% increase in density can be induced just by tapping the loaded tubes. The curve eye fitted to the average data indicates a steady state velocity regime is reached at a VOD in the order of 2100 m/s to 2200 m/s, which compares favorably with the ideal velocity of 2480 m/s as calculated by the Cheetah thermodynamic code. The critical diameter of the LDRA at the target density of 0.15 g/cm 3 and under steel confinement was in the order of 25 mm. As will be seen later, the failure diameter proved to be extremely sensitive to the degree of confinement of the LDRA charge. 91

110 Diameter Effect (Average VOD) VOD (m/s) Diameter (mm) Figure 5.5: Average VOD as a function of diameter for LDRA at 0.15 g/cm 3 In order to predict the experimental velocity of detonation at infinite diameter, the above velocity data can be presented in the form of VOD versus the reciprocal diameter (1/φ). This will result in a linear relationship between the two parameters, as illustrated in Figure 5.6. The ideal VOD is then inferred by extrapolating the regression line to intersect the VOD axis. An ideal VOD of about 2100 m/s is predicted, a velocity that closely resembles the steady state detonation regime indicated by the asymptotical trend seen in Figure

111 VOD at Infinite Diameter (0.15 g/cc, steel) VOD intercept ~ 2100 m/s VOD (m/s) Inverse Diameter (1/mm) Figure 5.6: Ideal VOD inferred from experimental data for LDRA at 0.15/gcc under steel confinement Experimental results indicate that under steel confinement, the LDRA at 0.15 g/cm 3 approaches ideal behavior rather quickly for such a heterogeneous mix. For example, the average VOD at 75 mm diameter for the LDRA was calculated at ~1900 m/s, thus: 1900/2100 ~ 90% of ideal velocity is developed. Under a similar testing scenario, a standard ANFO at 0.84 g/cm 3 reports a VOD of 3200 m/s and an ideal VOD of 4600 m/s, therefore: 3200/4600 ~70% of ideality is developed. This characteristic should prove advantageous for bottom priming, where maximum performance of the explosive is desirable Effect of Confinement The trend observed between the velocity of detonation and the degree of confinement of any explosive charge is explained by the reaction zone length, which although in 93

112 detonation theory is treated as a discontinuity having zero thickness, has in reality a finite dimension. Experimental studies (Cooper et al, 1996) indicated that these reaction zone lengths varied from a few hundredth of a millimeter for high density military explosives, to several centimeters for lower density commercial blasting agents. In addition, studies also showed that the reaction zone length appears to increase with increasing confinement and decreasing density. Since the LDRA is a low density particulate product with a relatively poor fuel oxidizer contact, a large reaction zone length is to be expected, in particular if conditions of high confinement prevail. Confinement played a major role on the sensitivity to initiation and propagation of the LDRA charges. Failure diameter and steady state VOD regimes were substantially different with changes in the degree of confinement of the explosive. This is shown in Figure 5.7 where an increase in steady state velocity and a decrease in failure diameter (from 75 to 50 to 25 mm) are observed as the degree of confinement increases from paper to cardboard to steel. It is worth mentioning that the LDRA charges confined in 50 mm diameter paper tubes did not initiate in spite of the 220 g primer used. 94

113 VOD (m/s) Confinement Effect (0.15 g/cc) steel cardboard papertube Diameter (mm) Figure 5.7: Effect of confinement on stable VOD regime and critical diameter For comparison purposes, a series of tests conducted elsewhere (Sudweeks, 2000) on a LDRA mixture at a density of 0.27 g/cm 3 under unconfined conditions (Sch 80 cardboard tubes) is presented in Figure 5.8. Comparing the unconfined data recorded at a density of 0.27 g/cm 3 with the confined experiments conducted at the target density of 0.15 g/cm 3, a relative measure of the effect that confinement has on explosive performance becomes apparent. 95

114 g/cc- (SCH 80 cardboard tubes) VOD (m/s) Diameter (mm) Figure 5.8: VOD of LDRA at 0.27 g/cm 3 confined in heavy duty cardboard tubes Data appear to indicate that reducing charge confinement from steel to cardboard practically compensates for the increase in density from 0.15 g/cm 3 to 0.27 g/cm 3, as indicated by the little effect observed in their respective steady state velocities, which are considered good indicators of the energy released at the detonation front Effect of Density As mentioned in previous chapters, both the velocity of detonation and particularly the detonation pressure are strongly dependent on the density of the unreacted explosive. There is a linear dependency between density and ideal VOD for a given explosive where as density increases so does the velocity of detonation. Experimental results for the LDRA are illustrated in Figure 5.9. The velocity data have been plotted separately for each of the two main diameters tested, since this parameter has an influence of its own on the VOD. For clarity purposes, average VOD values are 96

115 included in the graph for those conditions tested the most: 50 mm and 75 mm diameter under steel confinement Density Effect (steel confinement) VOD (m/s) mm 50 mm Density (g/cc) Figure 5.9: Density vs. VOD for 50 mm and 75 mm confined LDRA charges In spite of the observed scatter, a linear regression to the experimental data shows the expected increasing trend of VOD with increasing density for both charge diameters. The difference in performance between the two diameters is attributed to larger energy losses occurring in smaller charges Initiation Behavior of the LDRA General observations regarding the behavior of the LDRA under various initiation scenarios, including primer weight and confinement conditions, will be addressed in this section. 97

116 The LDRA proved insensitive to blasting cap initiation when confined in thin paper tubes regardless of the charge diameter used. Moreover, the product failed to detonate in 50 mm diameter paper tubes in spite of using a 220 g pentolite primer. Under this particular confinement condition, a minimum primer weight of 40 g of emulsion and a minimum diameter of 75 mm were required for proper initiation. An intermediate degree of confinement was obtained by using thick wall cardboard and PVC plastic tubes. The initiation behavior of the LDRA under these conditions proved somewhat similar to the paper tube confinement in that it remained insensitive to blasting cap initiation, regardless of charge diameter. Experiments in 75 mm diameter cardboard tubes showed primer requirements for initiation in the range of 20 g to 40 g emulsion primer. In addition, under this same confinement condition, in 50 mm diameter cardboard tubes, the LDRA detonated with a 220 g pentolite booster, although indications were that a smaller primer weight would have been more than sufficient. Worth noting is the behavior exhibited by the LDRA under steel confinement conditions, where the low density product at the target density of 0.15 g/cm 3 proved sensitive to blasting cap initiation at diameters of 25 mm, although it failed to detonate in 75 mm or larger diameters under equal conditions of confinement and priming. One would expect that under equal conditions of initiation and confinement, the product should also detonate under larger diameters, which was not the case for the LDRA. 98

117 This apparent contradictory behavior can be explained by Katsabanis (2002), where he used numerical modeling to examine, within other parameters, the effect that confinement has on priming an emulsion explosive. His findings concluded that confinement reflected the shock generated by the primer, increasing its amplitude and duration and consequently the decomposition rate of the explosive within the annular space. This shock interaction is responsible for creating high pressures zones (hot spots) within the explosive mass and promoting lateral initiation of the explosive mass. Therefore, increasing the diameter from 50 mm to 75 mm as was done for the LDRA, will reduce the amplitude of the reflected shocks and decrease the chances for lateral initiation of the explosive. Moreover, the previous model is in agreement with the concept of energy fluence (Cooper et al, 1996), which refers to the energy per unit area imparted to an explosive by an initiating shock. As shown in Equation 5.1, energy fluence is a function of the initial density ρo, the pressure P, the shock duration t and the shock velocity U imparted to the unreacted explosive by the initiating shock. E f 2 P t = ρ U 0 Equation 5.1 There is a critical threshold value of energy fluence (Ecr) below which the explosive will fail to initiate. These critical values are characteristic of each explosive: the more sensitive to shock an explosive is, the lower the critical energy becomes. Therefore, an 99

118 increase in charge diameter will decrease the amplitude and duration of the reflected shock and consequently reduce the energy fluence below the critical value necessary for initiation of the LDRA. Another set of experiments evaluating the initiation behavior of the LDRA involved using detonating cord. For some particular applications, such as ornamental stone quarries where diameters in the order of 32 mm are common, detonating cords could be used to top initiate the LDRA and avoid using the more expensive electric blasting caps. Experimental observations showed that the LDRA could be initiated in 25 to 38 mm diameter steel pipes by inserting a 15 cm length of detonating cord having a minimum PETN core load of 16 g/m. Velocities of detonation recorded from these experiments are shown in Table 5.1. Table 5.1: Initiation of LDRA at 0.15 g/cm 3 with detonating cord Diameter 15 cm ends Core Load VOD (mm) (#) (g/m) (m/s) 25 1 x Failed 32 1 x Failed 38 1 x x x As gathered from the previous table, a detonating cord having a core load of 10.6 g/m and less did not initiate the LDRA under the stated setup conditions that closely resemble the small diameters and highly competent rock typical of dimensional stone quarries. At 16 g/m core load, the initiation energy provided by the 15 cm long 100

119 detonating cord end appeared enough to detonate the LDRA, however, if the recorded velocity of detonation is compared to the ones obtained for higher core loads, indications are that initiation with this core load may have been marginal. Finally, detonating cords having core loads of 32 g/m and above seemed excessive, since no improvement was observed in the recorded velocities of detonation Priming Recommendations for the LDRA From the experimental observations regarding priming of the LDRA, the following summary is presented: Under confined conditions, the LDRA at 0.15 g/cm 3 can be reliably initiated with a 40 g primer in diameters ranging between 25 mm and 200 mm, which should be no challenge for the 150 g to 900 g primers typically used in the industry In the presence of good competent rock conditions, a blasting cap proved sufficient to initiate the LDRA in small diameter blastholes ranging from 25 mm to 50 mm, however, the use of 3 g to 8 g primers is considered a safer approach For 75 mm to 125 mm diameter in competent rocks, a minimum primer of 3 g to 8 g is a requirement for reliable initiation; however, 40 g or higher is recommended For 125 mm diameter and beyond, in competent rock, 40 g primers or larger depending on diameter are recommended For diameters between 75 mm to 125 mm under soft and or heavily fractured rock resembling unconfined conditions, a 40 g to 220 g primer should prove sufficient Beyond 125 mm diameter, 220 g primers are recommended 101

120 For the typical small diameters used in dimensional stone quarrying, short ends of detonating cord having a core load of 16 g/m or over proved reliable to top initiate a column of LDRA Propagation Stability of the LDRA in Longer Charges As mentioned in previous chapters, the experiments conducted on the LDRA at densities less than 0.10 g/cm 3 showed indications of unstable detonations in the form of periodic bulging of the pipe walls. This behavior has been identified in gaseous explosions and low density explosive products of similar nature to the LDRA. In order to ensure that the propagation front at the target density of 0.15 g/cm 3 will remain stable for longer distances, two experiments were conducted in 3.20 m long steel pipes: one in a 50 mm diameter tube and the other in a 75 mm tube. To record the VOD along the entire length of these long charges, two 1.80 meter long aluminum sensors were required on each test. The possibility of using coaxial resistance cable sensors, the available technology to measure VOD in boreholes, was discarded since the low crushing pressures generated by the LDRA would result in records difficult to interpret. The LDRA charges were initiated with a 220 g pentolite primer to ensure proper detonation. Experimental results corresponding to the two sensors placed in each tube are presented in Table

121 Table 5.2: VOD for LDRA at 0.15 g/cm 3 in 3.20 meter long steel pipes Diameter VOD Probe 1 VOD Probe 2 Average VOD (mm) (m/s) (m/s) (m/s) In spite of the discrepancy in VOD observed between the diameters, a stable detonation regime is maintained throughout the entire length of the explosive charge. Figure 5.10 illustrates the experimental set up and the results after the detonation event took place. Figure 5.10: Before and after events of VOD experiments conducted on 3.20 m long, 50 mm and 75 mm diameter steel tubes loaded with LDRA at 0.15 g/cm 3 The stability of the propagation front is better visualized by plotting on the same graph the records corresponding to the two probes positioned in the tube. This is illustrated in Figure 5.11, which corresponds to the 75 mm experiment. 103

122 g/cc- 75-mm, 3.2 m long steel tube 2.5 Distance (m) Probe 1 = 1971 m/s Probe 2 = 1969 m/s Time (ms) Figure 5.11: Composite traces for the two aluminum probes used to measure VOD in the 3.20 meter long, 75 mm steel tube loaded with LDRA at 0.15 g/cm 3 As observed in the above figure, the slopes corresponding to each of the two probes contained in the same pipe are quite parallel, an indication of the constancy of the VOD along the entire length of the tube. Worth noting is the primer effect on VOD, which can be observed as a steeper slope at the beginning of the first trace. The previous experiment presented irrefutable evidence that the LDRA at the target density of 0.15 g/cm 3 sustains stable detonation away from the primer influence, making it a valid alternative for use in quarrying, benching and any long hole application Behavior of LDRA in Decking Configurations One of the issues that needed to be investigated was the capability of the LDRA to initiate a column of regular ANFO. Two reasons justify this need. The first one has to do with the possibility of using the LDRA as an intermediate deck, to provide a continuous but variable density explosive column. The second reason has to do with safety and 104

123 considers the possibility of total segregation of a LDRA/ANFO mixture during mechanized loading, generating a LDRA deck that could result in a partially detonated blasthole if it does not provide the energy necessary to initiate ANFO. With this objective in mind, a velocity of detonation experiment was conducted in a 1.20 meter long, 75 mm diameter steel pipe, with about half its length loaded with LDRA at 0.15 g/cm 3 and the remaining with ANFO at 0.85 g/cm 3. A 220 g pentolite primer was placed within the LDRA section for initiation. The resulting VOD record is presented in Figure ANFO Initiation Test Distance (m) LDRA = 1905 m/s ANFO = 3192 m/s Time (ms) Figure 5.12: VOD record of a 75 mm diameter steel tube loaded with similar column lengths of LDRA at 0.15 g/cm 3 and ANFO at 0.85 g/cm 3 The previous experiment proved that the energy delivered by the LDRA at a density of 0.15 g/cm 3 is sufficient to initiate a column of regular ANFO prills. The ability to transmit detonation between both explosives opens the opportunity for using the LDRA 105

124 in decking configurations, such as in long holes, eliminating the need for priming individual decks. Further testing conducted elsewhere (Sudweeks, 2000) evaluated the behavior of LDRA at a density of 0.27 g/cm 3 for decking applications. Three experiments were conducted using a 3 deck configuration, as illustrated in Figure In all three tests the upper and lower decks were loaded with a LDRA/ANFO blend. Two of these tests used a 50/50 volume blend (0.52 g/cm 3 ) while the remaining experiment used a 70/30 blend (0.42 g/cm 3 ). Moreover, two of the experiments had the LDRA (0.27 g/cm 3 ) placed in the middle deck while for the third experiment, the LDRA was replaced with straight expanded polystyrene beads to assess differences in behavior and evaluate inter deck propagation. The experiments were conducted in 125 mm diameter, 3 m long PVC plastic pipes, divided in three decks measuring 1.0 m long each. Explosive charges were initiated from the top deck. 106

125 Test 1 Test 2 Test 3 LDRA/ANFO 50/50 vol LDRA/ANFO 70/30 vol LDRA/ANFO 50/50 vol LDRA 0.27 g/cm 3 LDRA 0.27 g/cm 3 Polystyrene beads LDRA/ANFO 50/50 vol LDRA/ANFO 70/30 vol LDRA/ANFO 50/50 vol Figure 5.13: Test configuration used for the evaluation of the LDRA in decking applications Table 5.3 summarizes the VOD results obtained during these tests. Table 5.3: Evaluation of LDRA in a decking configuration VOD (m/s) Deck Test 1 Test 2 Test 3 top middle Failed bottom Failed The following conclusions are obtained from the above experiments: The LDRA/ANFO blends at both volume ratios provided enough strength to initiate the LDRA at 0.27 g/cm 3 under unconfined conditions The LDRA at 0.27 g/cm 3 can in turn initiate either of the two LDRA/ANFO blends A 1 m long expanded polystyrene deck will attenuate the shock wave originated from a 50/50 volume LDRA/ANFO deck to the point at which it will not initiate a 107

126 similar LDRA/ANFO deck. Therefore, a disruption of the detonation wave leading to failure could occur if polystyrene/anfo mixtures are used and segregation occurs Mixtures of LDRA and ANFO When it comes to tonnage consumption, a product with the low density characteristic of the LDRA will have niche applications when utilized alone, with foreseen uses directed mostly at the implementation of control blasting techniques (presplit, buffer, decking) and dimensional stone operations. A much larger consumption tonnage however, is envisaged when the LDRA is treated as a diluting agent for ANFO rather than as a stand alone product. The mixing of both will open the door for large diameter production blastholes; therefore, it was considered important to gather basic information on the behavior of such mixes. As mentioned in previous chapters, expanded polystyrene (EPS) beads have been used to reduce the density of ANFO with varying degrees of success. Most of the problems arise from segregation issues and limitations to the degree of dilution that can be attained without risking detonation failure. In order to evaluate eventual differences between the LDRA and EPS when used as diluting agents of standard ANFO, a series of experiments was conducted on LDRA/ANFO and EPS/ANFO mixtures to compare their relative performance on an equal density basis. The first set of experiments compared the performance of the two mixtures at densities varying from 0.15 g/cm 3 up to 0.50 g/cm 3. The required proportions of ANFO at 108

127 0.85 g/cm 3, LDRA at 0.15 g/cm 3 and EPS at g/cm 3 were mixed to reach the selected densities, and then loaded into 50 mm and 75 mm diameter steel pipes and initiated with a 220 g primer. Comparison of the two mixtures could only be made at the 75 mm diameter since the VOD records corresponding to the EPS/ANFO mix loaded in 50 mm tubes were unreliable to interpret, even at the higher densities and despite using aluminum probes. These consistently poor quality records were a direct consequence of segregation between the polystyrene beads and ANFO prills. Figure 5.14 compares the traces produced by mixtures of ANFO diluted with LDRA and EPS beads respectively in 50 mm diameter charges. Records clearly show the degraded quality of the trace obtained when EPS rather than LDRA is used as bulking agent of ANFO in spite of the higher density of the EPS/ANFO mix. At lower densities, the quality of the EPS/ANFO records was shown to deteriorate further. On the other hand, good quality VOD records were obtained when the LDRA instead of EPS was used as the diluting agent of ANFO g/cc. 50-mm steel 0.40 g/cc. 50-mm steel Distance (m) LDRA/ANFO at 0.30 g/cm 3 Distance (m) EPS/ANFO at 0.40 g/cm Time (ms) Figure 5.14: Comparison of VOD traces generated by LDRA/ANFO and EPS/ANFO mixes Time (ms) 109

128 Figure 5.15 illustrates the velocities of detonation recorded in the 75 mm diameter steel tubes for mixtures of LDRA/ANFO and EPS/ANFO at various densities. For comparison purposes, the ideal velocities of detonation obtained from the thermo chemical code Cheetah for a LDRA/ANFO mixture has been included in the graph ANFO mixtures Ideal VOD (m/s) ANFO/LDRA ANFO/EPS Density (g/cm (g/cc) 3 ) Figure 5.15: VOD tests on mixes of ANFO/LDRA and ANFO/EPS at different densities The above figure illustrates the considerable difference in performance between the two mixtures at any given density. Since ingredient composition and density are similar for both mixtures, the performance difference can be attributed to the poor quality of the mix as a consequence of segregation of the EPS from the ANFO particles. The synergistic effect provided by the reactive nature of the LDRA will compensate for the eventual segregation and contribute to an increase in performance. The substantial difference in velocity observed between predicted (ideal) and experimental values can be attributed to both the EOS used by the computer code to 110

129 calculate ideal velocity and the non ideal behavior of the LDRA under the conditions in which it was actually used. A similar set of experiments conducted elsewhere (Sudweeks, 2000) under unconfined conditions (cardboard tubes) is included for completeness. Mixtures of LDRA/ANFO and EPS /ANFO at volume ratios of 50/50 and 70/30, with resulting densities of about 0.50 g/cm 3 and 0.40 g/cm 3, were tested at various charge diameters. Results are summarized in Table 5.4. Table 5.4: VOD of LDRA/ANFO and EPS/ANFO at different diameters and mix ratios (Sudweeks, 2000) LDRA/ANFO EPS/ANFO Volume ratio 50/50 70/30 50/50 70/30 Mixture density 0.52 g/cm g/cm g/cm g/cm 3 Diameter (mm) VOD (m/s) Detonated (n/a) Failed Failed Detonated (n/a) Detonated (n/a) Compared on the basis of equal density, the ANFO mixtures diluted with LDRA outperformed the ones diluted with EPS in regards to the steady state velocity. Except for the unexpected high velocity recorded at the 75 mm diameter, the usual velocitydiameter and velocity density trends are present. Worth noticing are the detonation failures of the 75 mm diameter EPS/ANFO mixtures at the two densities tested as well as what appears to be a low order detonation for the 100 mm diameter charge (1500 m/s at 111

130 0.44 g/cm 3 ). Again, segregation of the EPS from the ANFO is partly accountable for this behavior Mixtures of LDRA with Ammonium Nitrate (AN) Prills Expanded polystyrene beads (EPS) consist of long hydrocarbon polymer chains and as such, provide fuel to the chemical reaction and influence the oxygen balance of the mixture. Therefore, EPS/ANFO mixtures will be fuel rich, negatively balanced compositions, the degree of which will depend on the percentage of EPS beads present in the mix. Similarly, when LDRA at 0.15 g/cm 3 instead of straight EPS is used to dilute ANFO prills, no balanced compositions are possible with these two ingredients. The substitution of ANFO prills by AN prills is a feasible approach for improving oxygen balance and was considered for evaluation. A few VOD experiments were set to evaluate performance of LDRA/AN mixtures at various densities in 75 mm diameter steel tubes. Careful loading was required to decrease density gradients along the tube, however, segregation was unavoidable during loading. Test results are summarized in Table 5.5. Table 5.5: VOD of LDRA/AN mixes at various densities LDRA/AN Mixtures Density VOD (g/cm 3 ) (m/s)

131 From the few experiments conducted, there is no indication of the expected VOD density trend, that is, an increase in VOD as density increases. No significant errors can be attributed to interpretation of the velocity records, which as Figure 5.16 illustrates, were of good quality for both densities tested (0.25 and 0.35 g/cm 3 ). ρ = 0.25 g/cm 3 ρ = 0.35 g/cm 3 Figure 5.16: Quality of VOD traces for two experiments conducted on AN/LDRA mixtures This lack of trend can be attributed to a poor fuel oxidizer intimacy and a poor quality mix due to segregation, both of which will increase as density is increased with the addition of AN prills. These factors will lead to a strong non ideal behavior of the mixture; hence, the chemical reactions occurring in the expansion phase will not contribute to the VOD. As the weight percentage of AN is increased, more AN prills will stop contributing to an increase in VOD and eventually a reverse trend could occur. As a result, the products of detonation may generate gases of an oxygen rich and an oxygen deprived composition simultaneously; that is to say nitrogen oxides and carbonaceous gases will form. 113

132 5.3 Segregation Tests on LDRA and LDRA/ANFO Mixes In an attempt to gain an understanding of the effect that segregation has on density, laboratory tests were designed to evaluate, in a qualitative manner, segregation of the ammonium nitrate/mineral oil fines (ANMO fines) within the LDRA, density changes due to settling induced by tapping/vibration of the LDRA as well as density gradients as a result of gravity loading LDRA and diluted mixtures of LDRA and ANFO Segregation Tests on the LDRA At the target density of 0.15 g/cm 3, the ANMO fines comprising the LDRA composition proved more than sufficient to cover the surface of the polystyrene beads, with the excess fines accumulating between the interstices of the polystyrene beads. In order to evaluate segregation of these fines, LDRA samples were subjected to vibration and results observed for any indication of segregation of the ANMO fines from the LDRA mixture. For this objective, a 150 mm diameter plexiglas tube sealed at one end with a plexiglas plate was filled with the LDRA sample, placed on top of a sieve shaker and subjected to a high frequency vibration for a period of about five minutes. A blue dye, soluble in mineral oil, was used to provide contrast between the fines and the polystyrene beads and help visualize segregation. Observation of the vibrated samples revealed the presence of stratification or layering of the fines against the plexiglas wall, but these were not pronounced and it could have been the result of static charges generated on the plexiglas surface. More importantly, little segregation of fines was observed through the 114

133 plexiglas plate at the bottom of the tube, an indication that in spite of the induced vibration, the ANMO fines will tend to remain adhered to the EPS beads and within the voids provided by the polystyrene particles. Following on the previous experiments, a LDRA sample was then subjected to vibration for a full 24 hour period, again without noticeable segregation beyond the one observed in the 5 minute vibration tests. The photographs depicted in Figure 5.17 show a little accumulation of fines around the periphery at the bottom end of the plexiglas tube after removal of the back plate. The layering on the side view is hardly discernible in the photograph. ANMO fines Figure 5.17: Bottom and side view of a 150 mm plexiglas tube loaded with LDRA after being subjected to a 24 hour vibration test Another set of experiments conducted on the LDRA was led by the noticeable settling observed when the product was subjected to tapping or vibration and the effect this produced on density. It is well known that loading rate will also affect density and in the particular case of LDRA/ANFO mixtures, density will also be affected by the segregation 115

134 arising from the different free fall velocities of the ingredients, the latter enhanced by the out flowing air stream that is generated as the mixture is gravity fed into the blastholes. Given the difficulties associated with full scale trials, it was decided to conduct a few laboratory tests to quantify the effect of tapping on LDRA density, in an attempt to gain an understanding of the density changes that are to be expected. With this objective in mind, a batch of LDRA at 0.15 g/cm 3 was prepared and a 2000 cm 3 graduated cylinder used to calculate the untapped sample density. The graduated cylinder was then tapped until particle settling was unnoticeable, with the tapped density then calculated from the drop in column height. The process was repeated six times, each one with a new sample from the batch, with each pair of untapped/tapped measurements corresponding to the same sample. Results are presented in Figure Effect of Taping on LDRA Density 0.17 Density (g/cm (g/cc) 3 ) % 17% Untapped Tapped Sample # Figure 5.18: Effect of tapping on LDRA density 116

135 From the above graph we observed that upon tapping, particles will settle and the density of the LDRA will increase no more than 17%. Moreover, the density difference from samples of the same batch of explosive, whether they are tapped or untapped, is in the order of 4%. It is worth noting that the targeted density for the LDRA (0.15 g/cm 3 ) falls between the recorded tapped and untapped densities For higher density explosives, a 17% density change will result in a significant difference in product performance and hence on its ability to control fragmentation and reduce damage; however, at the low densities characterizing the LDRA, this percentage change will have no practical consequences. For example, a 17% density increase would bring a 0.85 g/cm 3 ANFO to a density of almost 1.0 g/cm 3, which is quite considerable and will carry practical implications as far as fragmentation and damage are concerned. However, if the same percentage change is applied to the LDRA, the resulting density of 0.17 g/cm 3 will not produce significant differences in fragmentation and damage. An additional set of experiments was aimed at measuring the variation of density with column length; in other words, density gradients along the length of a tube loaded with LDRA. For this purpose, a 75 mm and a 125 mm diameter cardboard tube was divided into several sections of equal length by means of blast gates of the appropriate diameter, as shown in Figure With the blast gates fully opened, the LDRA was freely poured into the tubes. Once the tubes were fully loaded, the blast gates were closed in order to separate the tube into the several sections from which individual density measurements were taken. 117

136 Figure 5.19: Test setup to measure density gradients Density measurement results for each of the two tube diameters tested are presented in Figures 5.20 and 5.21 next g/cc. Density Gradients in 75-mm tubes Density (g/cm Density Density (g/cm (g/cc) 3 ) 3 ) Trial 1 Trial 2 Trial 3 Trial 4 Trial 5 Trial Tube Section Figure 5.20: Density readings of LDRA taken from four sections of a 75 mm tube 118

137 g/cc. Density Gradients in 125-mm tubes 0.17 Density (g/cm (g/cc) 3 ) Trial 1 Trial 2 Trial 3 Trial 4 Trial Tube Section Figure 5.21: Density readings of LDRA taken from four sections of a 125 mm tube The maximum density gradient measured between any two sections of a tube for any individual experiment was 17%, which on average is reduced to about 11%. This 17% value coincides with the settling experiments mentioned before. Diameter appears to have no significant effect on density gradients. From the previous observations, which do not account for the effect of column height, it would be safe to expect density differences less than 20% for an explosive column loaded with LDRA at the target density of 0.15 g/cm 3. Remembering that pressure is a cube function of density, this 20% difference in density will translate into pressure differences as high as (1.20) 3 = 72%, a percentage that should be considered acceptable for the LDRA explosive Segregation Tests on LDRA/ANFO Mixes Similar experiments, this time with mixtures of LDRA and ANFO prills, were conducted to qualitatively evaluate the effect of segregation on density gradient when loading a 119

138 mixture containing these two very different ingredients. Results were not expected to be representative of bulk loading into larger and longer boreholes, where the different settling velocities of the two main particles and the influence of the turbulent medium they create upon free falling cannot be accounted for in small scale laboratory experiments Nevertheless, experiments proceeded with the expectation of extending some laboratory observations to field conditions. The same blast gate arrangement used to measure density gradient for the LDRA was utilized. Tests were conducted by loading a LDRA/ANFO mixture at a nominal density of 0.50 g/cm 3 into the 75 mm diameter cardboard tube, the latter divided by means of blast gates into seven sections from which individual density readings were taken. For comparison purposes, tests were also conducted on a mixture having the same nominal density but using expanded polystyrene (EPS) beads wetted with mineral oil instead of the LDRA as the diluting agent of ANFO. Figures 5.22 and 5.23 show the results obtained for both these mixtures g/cc (76-mm tubes) 0.7 Density (g/cm 3 ) Density (g/cc) Test 1 Test 2 Test 3 Test Tube Section Figure 5.22: Density gradient in mixtures of LDRA/ANFO 120

139 g/cc (76-mm tubes) 0.7 Density (g/cm 3 ) Density (g/cc) Test 1 Test 2 Test Tube Section Figure 5.23: Density gradient in mixtures of EPS/ANFO For the LDRA/ANFO tests, most of the density readings fluctuate within the 0.40 g/cm 3 to 0.60 g/cm 3 range, dropping considerably mostly at the bottom section of the tube (section #7). This drop in density at the bottom section can be explained by the behavior of the mixture at the point of discharge. A double layer bed with a larger volume of the lighter material on top will form and fall first until the discharge becomes uniform. As seen in Figure 5.23, this tendency becomes more pronounced as the difference in density between the ingredients becomes larger, as would be the case when ANFO is diluted with EPS beads instead of LDRA The few experimental results indicate that wetted EPS beads tend to segregate more from ANFO prills than LDRA does. This observation, explained by the lower density of the EPS beads with respect to the LDRA, could only be magnified if dryer EPS beads are used, particularly in larger diameter and longer holes, where turbulent flow and settling velocities become a critical issue. 121

140 5.4 Pressure Measurements on the LDRA The importance that detonation and explosion pressure has on controlling damage has been reported in previous chapters. Experimental approaches to measure high pressure transient events include X ray techniques, the standard aquarium test and the use of pressure sensors. This section describes the work performed to measure the detonation and explosion pressures generated by the LDRA explosive, with particular emphasis on pressure sensors. Although X ray methods were not used during this research, a few comments regarding the measurement of detonation pressure by means of this technique are included in Appendix 5.2. In addition, the same appendix describes with more detail the fundamental concepts behind the standard aquarium test used to infer detonation pressure. This standard aquarium method was used as reference in the attempt to develop a modified experimental technique that will be described next Modified Aquarium Technique The same working principle as detailed for the standard aquarium technique described in Appendix 5.2 was used to investigate the potential of a modified aquarium test offering a much simpler experimental approach than photographic methods do. The modified Aquarium technique made use of the same small diameter aluminum tubes used for VOD measurements, with the expectation that they will produce clean signals and reliable records at the relatively low pressure amplitudes generated by both the LDRA and the shock it transfers to water. 122

141 Figure 5.24 shows the actual experimental setup alongside a diagram clarifying its principal features. The initial test was conducted with ANFO at 0.82 g/cm 3 loaded in a 75 mm diameter steel pipe and initiated with a 220 g pentolite primer. Results of this particular experiment are illustrated in Figure Primer LDRA Thin water seal Water container Aluminum probe Figure 5.24: Photograph and schematic arrangement of the modified aquarium test 1.4 Modified Aquarium ANFO 0.82 g/cc in 75-mm steel 1.2 Distance (m) VOD anfo = 2834 m/s Shock in w ater = 2705 m/s Time (ms) Figure 5.25: Record obtained from the modified aquarium method showing a constant VOD on the explosive section and a decaying shock velocity in water 123

142 Although the record obtained with ANFO shows a clear change in velocity as the detonation shock enters the water and travels through the inert fluid, it was difficult to infer the shock velocity at the explosive water interface; that is, the point of transition from one medium to the other. A linear fit conducted on the experimental record at either side of the interface produced velocities of 2834 m/s for the ANFO and 2705 m/s for the shock in water at the point of entry. Using the mismatch equation presented before, a detonation pressure of 14 Kbar is inferred, a value that compares reasonably well with the value of 16 Kbar calculated from the equation typically used to approximate detonation pressures (i.e. P = ρ D 2 4 ). However, as the pressure generated by the test explosive decreases, as is the case of the LDRA, so does the quality of the record and consequently the uncertainty in interpreting the results. This is shown in Figure 5.26, where the transition point, which could not be clearly identified from the record, was approximated from the positioning of the probe. The poor quality of the decaying shock following the interface would preclude proper evaluation of the record. 124

143 Modified Aquarium 0.15 g/cc- 50 mm steel Distance (m) approximate transition point VOD = 1872 m/s Time (ms) Figure 5.26: Modified aquarium test conducted on LDRA at 0.15 g/cm 3 Worth noting from the above two graphs is that while the ANFO test consumed about 70 cm of the aluminum probe, the LDRA crushed only 40 cm, most of which corresponds to the section of probe positioned within the explosive. In other words, not enough pressure was generated by the LDRA in the water to crush the aluminum sensor and generate a proper voltage signal. In an attempt to eliminate the ambiguity presented by the transition point from explosive to water, it was decided to replace the single aluminum probe with two probes instead, one positioned within the explosive and the other one within the water and contacting the interface as illustrated in Figure

144 Primer LDRA Thin water seal Water container VOD Aluminum probes Figure 5.27: Schematic of the experimental arrangement used to infer detonation pressure by measuring explosive VOD and shock velocity in water via two aluminum probes For comparison purposes, the traces corresponding to the detonation front and the shock in water are plotted together in Figure The shock in water generated a steplike and short trace (~10 cm), preventing proper interpretation of the velocity and leading in most cases to shock velocities (Usw) lower than the Cow of water; thus, negative particle velocities (uw) and pressures (Pw) would be obtained. 600 Modified Aquarium 0.15 g/cc (2 Al probes) Distance (mm) LDRA = 1806 m/s Shock w ater~1350 m/s Time (microsec) (μs) Figure 5.28: Records produced by aluminum probes placed in the LDRA and in water in order to measure VOD and shock velocity respectively 126

145 It was evident from the previous experimentation that the thin aluminum sensors were not capable of generating clean signals when subjected to the low pressures induced in water by the LDRA. In conclusion, the modified Aquarium method seemed to generate reasonable results for high density explosives, producing a sharper transition and a fair quality record of the shock in water. However, the quality of the attenuated shock in water is poor and does not allow a proper application of the method when lower density products such as the LDRA are used. The experimental effort was discontinued for the remainder of the thesis in order to concentrate on established techniques; however, further study in this area is recommended for future research Pressure Sensors Experimental measurement of detonation pressure by means of pressure gages is not an easy task to accomplish. Explosive discontinuities, temperature effects, shock wave effects (electromagnetic frequency pulses), embedment material of the gage, etc., have been responsible of contributing to ambiguous measurement results. There has been, however, a variety of piezo resistive and piezo electric sensors that have been used to measure detonation pressure including manganin gages, carbon gages, tourmaline gages, quartz crystals, lithium niobate crystals, polyvinylidene fluoride (PVDF) film gages and carbon composition resistors (CCR). Selection of the appropriate gage will depend not only on the pressure range to be measured, but also on cost, the latter an important consideration given the destructive nature of the experiments. 127

146 From the above choices, two fairly different gages were selected: the piezo resistive carbon composition resistors and the PVDF piezo electric film gages. The former were used as the principal gage for this investigation, given their low cost, availability and particularly their acceptance as a reliable sensor for harsh non homogeneous environments such as the LDRA. On the other hand, PVDF film gages, which are more than two orders of magnitude more expensive than carbon composition resistors, were selectively used to investigate their performance and reliability when subjected to products with the particular characteristics of the LDRA PVDF Film Sensors The PVDF piezoelectric film sensors are self powered gages, having a stress range of over 400 Kbar. They are unobtrusive in nature, have an extremely fast response (nanosecond range resolution) and their output is stress rate dependent, all of which makes them ideal for measuring many high pressure dynamic events. However, they are fairly expensive gages requiring careful assembly, fast recording instrumentation and high quality connecting lines to reduce signal degradation and generate reproducible output. PVDF gages consist essentially of a polymeric film that has been mechanically stretched to a 25 μm thickness and poled to a certain remnant polarization by the application of high voltage. With a carefully controlled manufacturing process, the charge induced by polarization responds to the application of stress in a well defined manner. PVDF gages from two US manufacturers were used during these experiments: Dynasen Inc. based in 128

147 California and Ktech Corporation based in New Mexico. The main characteristics of the PVDF gages, corresponding to each manufacturer, are shown in Table 5.6. Table 5.6: Characteristics of PVDF sensors used during experimentation PVDF Gage Characteristics Dynasen KTech Sensing element size (mm) 6.35 x x 5 Active area (cm 2 ) Thickness (μm) Encapsulating material Kapton n/a Remnant polarization (μc/cm 2 ) n/a 9.0 to 9.3 Mechanical stretching Uniaxial Biaxial Figure 5.29 depicts a PVDF gage from Ktech Corp. showing the sputtered Pt/Au electrodes deposited on opposite sides of the polymeric film. The polarized area (i.e. sensing area) of the PVDF sensor is located where the two electrodes of the gage cross each other. This sensing area comes in different dimensions (1 mm 2 to 5 mm 2 ), where the greater the area the larger the voltage output produced by the sensor for a given pressure input. For the low detonation pressures generated by the LDRA, a sensor with a medium to large sensing area was selected. PVDF film Sputtered gold and platinum electrodes Polarized sensing area Figure 5.29: PVDF film gage manufactured by Ktech Corp. 129

148 PVDF sensors can be used in two different recording modes: charge mode and current mode. The charge mode requires either a hardware integrator, consisting of a resistor and capacitor (RC circuit) housed within a brass unit suited for direct connection to the BNC terminal at the scope. Alternatively, similar RC circuits can be soldered directly at the sensor s leads, which brings the benefit of eliminating the capacitance of the connecting coaxial cable as a source of signal degradation.. The capacitor (C) in the circuit stores the charge (ΔQ) flowing from the gage, which is then observed in the scope as a voltage drop (ΔV) according to ΔV = ΔQ/C, from where the charge ΔQ can then be determined. Figure 5.30 depicts the hardware integrator acquired from Dynasen Inc., a model which came fitted with a 50 Ω resistor and a 100 ηf capacitor. 50 Ω 0.1 or 0.01 pf Figure 5.30: Dynasen s RC hardware integrator for charge mode recording with PVDF gages Figure 5.31 illustrates the alternative approach for charge mode recording, whereby the RC circuit has been connected to the coaxial cable at the PVDF sensor end. 130

149 RG 174 coaxial cable 50 Ω resistor 100 ηf capacitor PVDF leads Figure 5.31: Charge mode recording with the RC circuit connected at the PVDF leads In the current mode recording, the charge (ΔQ) flows from the sensing element through a shunt resistance (R) placed between the PVDF electrodes and is viewed as a current according to i = ΔV/R, with the latter also written as i = ΔQ/Δt = ΔV/R. In order to infer Q, numerical integration of the current time data is required as per Q = i * dt. In other words, the voltage drop (ΔV) across the known resistance (R) provides the current, which in turn is representative of the charge (Q) flowing from the gage as a consequence of the applied pressure. Compared to the charge mode, the current mode has the advantage of requiring numerical rather than hardware integration, thus preventing spurious voltages from being integrated into the output signal. Charge mode, on the other hand, has the advantage of generating outputs that are a direct representation of the pressure pulse waveform. Regardless of the recording mode used, the resulting voltage drop will be a direct function of the charge (ΔQ), which can be converted into pressure by means of 131

150 established ΔQ P calibration curves. The basic calibration equations provided by Dynasen (charge mode) and KTech (current mode) respectively are presented in the next two equations: P( kbar) Q = 5.8 Q Q Equation 5.2 P ( kbar) = Q Q Q Q Q Q Equation where Q( μ C / cm ) is the charge per unit area generated by the event. Worth mentioning is that in Dynasen s calibration equation, Q is affected by a temperature compensation factor, which for brevity has not been included. Both recording modes were attempted during these experimental trials and although the charge mode pressure output is straightforward, the fact that it integrates any existing electrical noise into the signal made the current mode recording the preferred mode of operation Assembly and Recording Instrumentation The PVDF gages, as received from the manufacturer, were mounted on a 32 mm diameter, 90 mm long plexiglas rod by applying a thin layer of a 5 minute epoxy glue. Care was taken to prevent the formation of air bubbles during the gluing process. The plexiglas rod was machined to enable a smooth folding of the sensing element at one end as well as to protect the soldered leads. The whole assembly (PVDF Plexiglas rod) 132

151 was then embedded in a liquid epoxy resin placed within a larger cylindrical mold and let to harden. The sensing end of the assembly was later covered with a thin aluminized Mylar film to prevent electromagnetic frequency (EMF) pulses from interfering with the digital recording. Figure 5.32 illustrates a PVDF sensor assembled for charge mode recording (i.e. soldered RC circuit) and connected to an RG 174 coaxial cable, showing the sensing area of the PVDF gage folded on the impacting end of the plexiglas rod. The assembly was later placed within a cylindrical mold and encapsulated in epoxy resin for further protection. Soldered R C circuit Machined Plexiglas rod Sensing area RG 174 pig tail cable Figure 5.32: PVDF gage assembly set for charge mode recording Current mode recording was conducted by soldering a 1Ω resistor across the leads of the gage so the recorded voltage is viewed as current. Numerical integration of the resulting current time data required to obtain ΔQ was implemented either in a spreadsheet or by 133

152 means of PlotData, software developed at Sandia National Laboratories (Wackerbarth et al, 1992) for the evaluation of PVFD responses under current mode recording. Figure 5.33 is a screen plot generated by PlotData, where the Y axis refers to the amplitude of the shock pressure (Kbar) as read by the sensor, while the X axis represents time in microseconds. This particular experiment corresponds to a 50 mm Sch 80 steel pipe loaded with LDRA at 0.15 g/cm 3. Both, the current trace recorded by the oscilloscope and the corresponding numerical integration representing the pressure as observed by the PVDF gage are shown. The actual detonation pressure is then inferred from the shock amplitude by means of the reflected Hugoniot method as described later in Appendix 5.6. Current trace Pressure trace Figure 5.33: PlotData screen plot showing results on LDRA test using current mode configuration. Both the current trace and the corresponding pressure profile are shown Recording equipment consisted of a Tektronix digital storage oscilloscope (DSO) capable of a 5 GHz sampling rate. Most measurements were taken at sampling rates of 134

153 1 GHz or higher. The short length of RG 174 coaxial cable leading from the PVDF assembly was initially connected to a 30 m long RG 58 coaxial cable extending to the actual DSO recording unit. The latter cable was later changed to a better quality coaxial (RG 8) to reduce degradation of the output signal Experimental Analysis and Results The configuration of the PVDF experiments was modified as testing progressed in order to accommodate suggested improvements. The initial tests were conducted in charge mode, using Dynasen s hardware integrator connected directly to the DSO recording unit, although this setup produced suspicious voltage output and was discontinued. Following suggestions from Sandia National Laboratory (Anderson et al., 2001), the hardware integrator was connected in close proximity to the PVDF gage rather than at the DSO unit, in an attempt to reduce signal degradation. Later changes included the substitution of the hardware integrator by an RC circuit soldered directly at the gage leads (Figure 5.31) as well as the replacement of the RG 58 for a higher quality RG 8 coaxial cable, the latter shown to reduce signal degradation by a factor of three. Final changes involved the switch from charge to current mode by replacing the RC circuit for a 1 ohm current viewing resistor (CVR) soldered between the leads of the PVDF. Figure 5.34 illustrates typical pressure profiles generated by the PVDF gages under charge and current recording modes respectively. Table 5.7 summarizes the results of all PVDF pressure tests conducted on LDRA under both recording modes. PVDF pressure traces are included in Appendix

154 PVDF # 6 (Charge mode, RC circuit soldered at leads) PVDF # 14 (Current mode, 1 Ohm CVR) Pressure (Kbar) Charge mode Pressure (Kbar) Current mode Time (u sec) Time (u sec) Figure 5.34: Comparison of charge and current mode records for LDRA in 50 mm steel pipes Table 5.7: Detonation pressure results obtained from PVDF gages on LDRA at 0.15 g/cm 3 Test Dynasen KTech Pressure (Kbar) Pressure (Kbar) Pressure (Kbar) PVDF # Charge Mode Current Mode Current Mode Comments (*) HI at scope (*) HI at scope (*) HI at scope (*) HI in line by gage HI in line by gage RC soldered at leads CVR soldered at leads CVR soldered at leads CVR soldered at leads (*) CVR soldered at leads CVR soldered at leads CVR soldered at leads CVR soldered at leads CVR soldered at leads CVR soldered at leads CVR soldered at leads CVR soldered at leads CVR soldered at leads Average HI = Hardware Integrator; RC = RC circuit; CVR = current viewing resistor (*) excluded due to unreliable output 136

155 The pressures recorded by the PVDF sensors proved higher than empirical approximations and theoretical codes for the LDRA at 0.15 g/cm 3. The detonation pressure, as calculated by Cheetah using the BKW equation of state (EOS) for the LDRA, was 3.4 Kbar. The pressure difference can be partly attributed to factors beyond the experimental approach, of particular importance being the equation of state for detonation products used by the thermo chemical codes being inappropriate to predict the performance of heterogeneous mixtures at densities almost an order of magnitude lower than typical explosive products. On the other hand, there are issues related directly to the experiments that may have also contributed to the observed high pressures. One such issue includes the fact that neither density nor VOD were measured for each individual experiment; an average VOD value and the target LDRA density were used instead. It is feasible that during loading, the density of the LDRA in contact with the PVDF element was higher than the target density used for the calculations and as a result, a higher VOD would also be produced. PVDF sensors may not represent the best alternative to measure detonation pressures in low density, highly heterogeneous regimes characterized by uneven detonation fronts, so other methods needed to be investigated. 137

156 Carbon Composition Resistors Carbon composition resistors (CCR) provide a relatively inexpensive alternative to evaluate dynamic and semi static pressures. CCR are rugged, simple to work with and produce reliable results in environments where no other pressure transducer has been successfully applied. Research conducted by Ginsberg et al (1991) proved their feasibility in difficult environments as found in multidimensional flows produced by heterogeneous, composite or particulate reactive materials such as large particle gunpropellants. Given the particulate characteristics of the LDRA mixture, using CCR should represent a suitable alternative to measure detonation pressure in this explosive. According to Ginsberg et al, carbon composition resistors have a response time in the microsecond range (as opposed to the nanosecond range displayed by the PVDF gages) and are claimed to respond reliably up to a maximum pressure of 50 Kbar for the case of the 470 ohm nominal resistor and of 160 Kbar for a nominal resistance of 4700 ohm Assembly and Calibration Equation Selection Various experimental arrangements have been used by researchers to calibrate the pressure response of carbon composition resistors; these responses being measured in terms of relative resistance and relative conductance (inverse resistance) change. Ginsberg et al (1991) used gas gun and aquarium experiments to produce calibration equations, Hollenberg (1986) used shock wave experiments in water for his calibration experiments while Wieland (1987 and 1993) proposed a calibration constant relating pressure and relative conductance change that was claimed to work well for pressure amplitudes below 1.0 Kbar. Austing et al (1991) calibrated the resistors by subjecting 138

157 them to an explosive source whose CJ pressure was estimated using the experimentally measured VOD along with empirical relationships available in the literature. The resistors used during Austing s experiments were encapsulated in polymer blocks and attenuator plates were positioned between the explosive source and the resistor gages; therefore, impedance mismatch techniques were used to determine the amplitude of the detonation pulse. The same brand and type of carbon composition resistors used by the previous authors was selected for the detonation pressure experiments conducted on the LDRA, that is, resistors manufactured by Allen Bradley having a nominal resistance of 470 Ω and 1/8 W power. Following suggestions proposed by Ginsberg et al (1991) and illustrated in Figure 5.35, the resistors leads were rugged and extended by crimping them within small diameter brass tubes; the whole assembly was then wrapped in polyethylene shrink tubing for insulation and added protection. Different protective configurations were tested throughout the trials, including embedding the resistors in wax, gel or within water filled containers. In the latter cases, the pressures read by the sensors corresponded to those acting on the transfer media; thus, shock impedance mismatching methods were required to infer the incident detonation pressure. 139

158 Resistor s leads CCR Brass tube (1.6 mm ID) Polyethylene shrink tubing Figure 5.35: Setup used to protect the carbon composition resistors (CCR) The two best known calibration equations are those developed by Wieland (1987) and Ginsberg et al (1991) and are presented next. Wieland: Ro R P( Kbar) = R Equation 5.4 for (0 < P < 1.0 Kbar) Ginsberg et al: P R R ( ) R o Ro ( GPa) ( ) ( ) 8.40 e Ro R = Equation 5.5 for (0 < P < 50 Kbar) In the above equations, Ro and R refer to the original and the instantaneous resistances respectively, expressed in ohms. As mentioned before, the C J condition computed by Cheetah for the LDRA at 0.15 g/cm 3 produced a detonation pressure of 3.4 Kbar. However, when the resistor gage 140

159 was embedded within wax, water or a hard polymer material, as it was in most cases, the resulting shock pressure was expected to approximately double due to the impedance mismatch between the different media. Thus, maximum shock pressure amplitudes of 8 Kbar could be expected during the detonation test of the LDRA. The calibration equation proposed by Wieland (1987) during his cross borehole stress measurement research was claimed to produce reliable results for pressure amplitudes ranging between 0.1 and 1.0 Kbar, well below the pressure range of interest. Therefore, Ginsberg s relationship, claimed to perform for amplitudes up to 50 Kbar, was selected to determine the detonation pressure of the LDRA mixture Detonation Pressure Results The majority of the experiments conducted with the carbon composition resistors were done under confined conditions in 51 mm diameter steel tubes. As mentioned before, the resistors were protected with polyethylene shrink tubing and embedded in wax, epoxy or water filled plastic containers, then positioned in the steel tube with the connecting leads folded down stream of the detonation flow. From the various configurations evaluated, the one with the resistors embedded in a water filled bottle produced the cleanest and most reliable output. Figure 5.36 illustrates the shock pressure recorded by the sensor in water and the corresponding detonation pressure of the LDRA determined from shock Hugoniot reflections are shown. 141

160 LDRA Detonation Pressure Test 1 8 Pressure (Kbar) Water LDRA Time (μsec) Figure 5.36: Pressure generated by a CCR placed in water and shocked with LDRA at 0.15 g/cm 3. Detonation pressure of the LDRA is then determined from this record Figure 5.37 illustrates the detonation pressures (peak amplitudes) recorded for all the configurations tested; that is, embedded in wax, gel and water, for LDRA at 0.15 g/cm 3 confined in 50 mm steel tubes as recorded by carbon composition resistors Corresponding experimental records are presented in Appendix Detonation Pressure with CCR 0.15 g/cc in 51-mm steel Pressure (Kbar) Test # in Wax in Water in Gel Figure 5.37: Summary of detonation pressure experiments conducted in 51 mm diameter steel tubes using carbon composition resistors under different configurations 142

161 As seen in the previous graph, greater scatter is observed when the resistors are embedded in wax, while more reproducible results are obtained if gel or water is used instead. These larger dispersion values can be attributed to the tendency of wax to entrap air and/or crack upon cooling, providing a non homogeneous transfer media that will affect pressure output. In addition, much of the experimentation using wax as transfer media was conducted on different explosive batches at the initial stages of development of the LDRA. Considering all the experimental data recorded (i.e. wax, gel and water), an average detonation pressure of 3.72 Kbar is obtained. However, when only the experiments having the resistor embedded in water are considered, the average pressure becomes 3.60 Kbar. Both these averages fall within a reasonable range of the ideal detonation pressure as calculated by theoretical codes (3.4 Kbar). Since diameter and confinement were shown to play an important role on the velocity of detonation of the LDRA, a few experiments were conducted to observe pressure behaviour under unconfined conditions by loading cardboard tubes of different diameters with LDRA at the target density. Table 5.8 summarizes results of these experiments, where water was used as the transfer medium. Appendix 5.5 includes the corresponding pressure records. 143

162 Table 5.8: Average detonation pressures of LDRA under different loading conditions. Detonation Pressure (Kbar) Unconfined Cardboard Tubes Diameter 50 mm 75 mm 87 mm 100 mm Gage # Average In view of the scarce unconfined data, no conclusive statement can be given, although a definite trend between detonation pressure and charge diameter for the unconfined experiments is clearly revealed. Moreover, this trend appears to merge with the average detonation pressure recorded from the larger number of confined experiments conducted in 50 mm confined charges. The effect of confinement upon pressure is most clearly revealed when average detonation pressures under cardboard and steel confinement for the same charge diameter are compared, where as confinement increases, pressure increases from 2.6 to 3.6 Kbar Novel Explosion Pressure Measurement The biggest limitation entrained in the equations estimating pressure relates to the constant values assumed for the coefficient of adiabatic expansion (γ) as the products sequentially expand from detonation pressure to explosion pressure to borehole pressure. As mentioned in earlier chapters, this coefficient does not remain constant during the expansion process and the values typically assigned to the products 144

163 expanding from detonation to explosion pressure (γ=3) and from explosion to borehole pressure (γ=1.2) are not necessarily representative of the LDRA product. The experimental arrangement set to measure detonation pressure with the carbon composition resistors allowed for the recording of the peak values and in some instances, for part of the release wave, but it would not last to capture the full history of the products of detonation as they expand to borehole pressure and atmospheric values. The resistors in most cases would record just a few tens of microseconds before damage occurs. In addition, the Sch 40 steel tubes containing the explosive charge would rupture and release the detonation gases prematurely. The low pressures generated by the LDRA opened a window of opportunity to design experiments where the products of detonation could be contained in the steel pipe for a longer time to allow recording of the expanding products as they vent to the atmosphere. In order to record longer times, the whole experimental setup needs to survive the detonation process; these include the steel tube, the resistors, their connecting leads and the coaxial cable to the recording unit. A 51 mm steel tube with a thick 19 mm wall proved strong enough to withstand the shock and gas pressures generated by the LDRA without rupturing, allowing for repetitive tests to be conducted with the same unit. Furthermore, to contain the gases for a longer time within the tube, a steel plug machined to house the resistors and their leads was threaded at one end of the tube as shown in Figure

164 Steel plug housing resistors 19 mm thick wall steel tube Thin thread Gas channel Threaded steel cover Pressure chamber Resistor Threaded steel plug body Epoxy Connecting cable Figure 5.38: Threaded thick wall steel tube and steel plug used during the explosion pressure experiments. Cross section of assembly showing pressure chamber and resistor placement Several design changes were implemented throughout the testing stage before producing satisfactory results. The following issues proved relevant for defining the testing arrangement finally adopted: Upon gas pressurization, the steel plug will expand, damaging the threads and locking it in the tube. To enable reuse of the plug, its front section was extended and left unthreaded, thus providing protection to the threaded section while allowing some tolerance for expansion 146

165 A coarse rather than thin thread on the steel plug unit proved stronger to withstand explosion pressure, thus facilitating retrieval from the tube and improving system survival The carbon resistors and connecting leads were embedded in epoxy, which would act as the pressure transfer medium while providing additional protection from shock, heat and the impingement of expanding gases For further protection from shock and particle impingement, the chamber housing the carbon resistors was covered with a threaded plug. Pressure build up within the chamber was allowed via a 6 mm orifice drilled in the plug cover A stainless steel cable (designed for thermocouple applications) was initially chosen to extend the carbon resistors leads within the steel housing and to the connecting coaxial cable. To prevent the resistors and steel cable assembly from being displaced by the action of gas pressure, stainless steel Swagelok tube fittings were required The previous steel cable assembly was later replaced by a high tear strength duplex cable, which proved strong enough to withstand the crushing gas pressure generated within the steel plug. In addition, the hard epoxy embedding the resistors proved sufficient to hold the cable assembly in place with no need for special fittings. This cable assembly maintained the same level of survivability and proved much easier to replace than steel cables Figure 5.39 shows the main parts of the measuring system: the coarse threaded steel plug housing the carbon resistors and the corresponding cover, a stainless steel and duplex wire cable soldered to the resistors already embedded in epoxy and the 147

166 Swagelok fittings used to prevent the resistor steel cable assembly from being blown out by the gas pressure. Steel cover Threaded steel plug Rubber O ring Epoxy embedded resistor Stainless steel cable Duplex wire cable Swagelok tube fittings Figure 5.39: Components of the steel plug to house the resistor. A stainless steel thermocouple cable and a duplex wire cable connected to the CCR embedded in epoxy are also shown Development of a Calibration Equation The explosion pressure amplitudes expected for the LDRA fall between 1 Kbar and 3 Kbar. The pressure records generated by the proposed experimental arrangement were initially estimated by means of the two best known calibration equations, that is, the one presented by Wieland (Equation 5.4) and the one given by Ginsberg et al (Equation 5.5). Pressures calculated with Ginsberg s equation were within the expected range; however, the exponential nature of the equation produced negative pressure values during much of the rise when applied to the LDRA data. In addition, Ginsberg s equation is best suited for higher pressure 148

167 ranges than the explosion pressure expected from the LDRA. Wieland s equation on the other hand, claims reproducible results for pressures up to 1.0 Kbar, thus, for lower pressure amplitudes than expected from the LDRA. Given the ambiguity of the equations for the pressure range of interest, it was decided to make use of existing pressure data from aquarium experiments conducted by Katsabanis (1997) in order to generate a calibration equation deemed more suitable for this particular pressure range. Figure 5.40 shows the test setup used by Katsabanis, where the shock generated by an explosive donor is transferred into water through a 12 mm thick plexiglas plate. The velocity of the shock as it travels in water is continuously recorded via a streak camera. Knowledge of the shock velocity and available shock Hugoniot data of water allows the attenuation relationship (i.e. pressure versus distance of travel) to be calculated. Streak camera slit Aquarium Plexiglas plate Distance of Travel Explosive donor Pressure Mini booster Figure 5.40: Aquarium arrangement used by Katsabanis (1997) 149

168 Analysis of streak camera records obtained during Katsabanis experiments resulted in the following equation relating pressure (P) and distance (d): P( GPa) ( d ) = d e Equation 5.6 where P(GPa) is the shock pressure and d (mm) is the distance of travel measured from the explosive plexiglas interface to the point of interest; thus, the 12 mm plexiglas thickness needed to be accounted for when estimating the shock pressures in water. As illustrated in Figure 5.41, duplicating the previous aquarium arrangement, this time placing carbon composition resistors at measured distances from the bottom of the plexiglas plate, data relating resistance change ( R) and distance (d) can be established. Water Container Carbon composition resistors Plexiglas plate RDX explosive pellet and initiator Support Figure 5.41: Aquarium test used for developing calibration equation for low pressure regimes 150

169 Using the ( R d) correlation generated from the experiments along with Equation 5.6 (Pd), a calibration equation relating pressure (P) and resistance change ( R) can then be established via a non linear fit to the (P d) data. The resulting equation follows: P( MPa) = (( R R Equation R) / R0 ) (( R0 R) / 0 ) where P is given in MPa and R0 and R refer to the original and instantaneous resistance respectively, both expressed in ohms. The application range of this equation is limited to pressures below 300 MPa (3 Kbar). Given the variety of test arrangements and recording instrumentation under which existing calibration equations were developed, the use of Equation 5.7 has the added advantage that resistance change data were gathered using sensor configurations and recording instrumentation similar to those used during the actual explosion pressure experiments Validation of Experimental Approach In an attempt to test the experimental arrangement and evaluate Equation 5.7, detonating cord of various strengths was used during a series of initial trials. As shown in Figure 5.42, the detonating cord was centered within the steel tube and initiated with an electric blasting cap placed at the sensor s end in order to hold the gas pressure for a longer time before venting takes over. 151

170 Embedded Resistor Explosion chamber Thick wall steel tube Steel plug assembly Blasting cap and external trigger wire Detonating cord Figure 5.42: Experimental setup used to measure explosion pressure of detonating cord The resistance change recorded from the above experimental setup was then evaluated by means of Equation 5.7 to determine the corresponding borehole pressure. Figure 5.43 illustrates results of typical pressure traces that have been superimposed for comparison and clarity. 0.8 Borehole Pressure with Detonating Cord Pressure (Kbar) g/ m 32 g/m 64 g/m 96 g/m 128 g/m Time (ms) Figure 5.43: Borehole pressures generated by detonating cord of different strengths, evaluated with the experimental arrangement and calibration equation developed at Queen s 152

171 Experimental records clearly show the arrival of a shock front in the form of pressure spikes, followed by a relatively slow pressure build up produced by the detonation gases entering and pressurizing the resistor s chamber; the event then slowly venting to atmospheric conditions. Following the same line of thought as Nie (1999), the resulting peak amplitudes are interpreted as the semi static borehole and/or explosion pressure generated by the products of detonation impacting against the steel tube s wall. In an attempt to validate the proposed arrangement, the borehole pressure generated by the detonating cord was estimated by using the Canmet (1977) analytical equation, which is rewritten next for convenience. ρ D φ φ 2 e e 2 γ e 2 γ P b = ( ) = Pe ( ) Equation φb φb where Pb (Kbar) is the borehole pressure, Pe (Kbar) the explosion pressure, ρe (g/cm 3 ) the explosive density, D (m/s) the velocity of detonation, φe/φb the ratio between the explosive and borehole diameters and γ the coefficient of adiabatic expansion for the detonation products at the expanded state. As mentioned in earlier chapters, the value of γ does not remain constant but changes throughout the expansion process, decreasing with pressure as venting to atmospheric conditions progresses. Typically, γ values at venting conditions ranging between 1.1 and 1.5 have been used to estimate pressures. More recently, Cunningham (2006) reports 153

172 typical γ values ranging between 1.3 and 2. For the present validation exercise a value of γ = 1.4 was selected for both the LDRA and the detonating cord. In order to determine the borehole pressure generated by the detonating cord, either the density or the diameter of the PETN core needs to be known. Laboratory tests conducted by Meng et al (2005) in addition to information he obtained from Dyno Nobel s detonating cord manufacturing plant in Chile, produced PETN core densities close to 1.40 g/cm 3. Moreover, this value is in line with the value used by Sanchidrian et al (2002) during their modeling work. In the author s opinion however, this density appeared excessive and upon consultation with other scientists in the field, a more conservative value of 1.3 g/cm 3 was finally adopted for the calculations. At this density, the ideal velocity of detonation as determined by Cheetah was 6600 m/s, value that was also used for the calculations. Once the density and core load of the detonating cord are known, the diameter of the PETN core required by Equation 5.8 can be determined by means of the following expression: Coreload φ ( ) e = Equation 5.9 π ρ where diameter (φe) is given in cm, core load in g/cm and density (ρ) in g/cm 3. Figure 5.44 compares borehole pressures as calculated by the Canmet equation versus those obtained from the experimental arrangement applying the calibration equation 154

173 previously developed. Two sets of experimental data are plotted along the Canmet equation, one corresponding to the tests conducted in 0.9 m long steel tubes and the other to the tests in 2.4 m long tubes. The latter was to ensure enough pressure build up time was provided before venting to atmospheric conditions is reached. Figure 5.44: Validation of experimental setup used to measure borehole pressures with carbon composition resistors: experimental results versus output estimated by Canmet s equation As observed from Figure 5.44 above, pressure amplitudes for both sets of experiments on detonating cord followed a similar trend. As expected, increasing charge length from 0.9 m to 2.40 m produced higher borehole pressures, with the latter resembling closer the amplitudes predicted by Equation 5.8. It is worth noting that the gap between predicted and experimental results would have narrow further if a decreasing rather than constant gamma value was included in the calculations. There are a number of factors that have an important influence on the results generated by Canmet s equation. These include the uncertainties of the values used for density and or diameter of the detonating cord as well as the value assigned to the adiabatic 155

174 exponent for the detonation products as they first expand to the original volume (explosion pressure) and then to the volume of the containment (borehole pressure) Explosion Pressure Experiments with LDRA The same experimental setup and calibration equation (Equation 5.7) was used for the bulk loaded LDRA mixture. The fully coupled experiments set to measure explosion pressure were initially conducted with short 30 cm long LDRA charges. This length was later increased in a stepwise fashion to a maximum of 240 cm. Examples of typical explosion pressure records for a 30, 90 and 240 cm long LDRA charge have been superimposed for comparison in Figure Amplitudes, rise times and duration of the semi static pressure pulse are seen to be an increasing function of column length. Similar to the detonating cord tests, rise times in the millisecond as opposed to the microsecond range are present. Test results indicated that pressure amplitudes tended to level at charge lengths around 150 cm. Explosion Pressure: Fully coupled LDRA Pressure (Kbar) Time (ms) 30 cm 90 cm 240 cm Figure 5.45: Composite of explosion pressure records generated by a fully coupled LDRA 156

175 Figure 5.46 summarizes the average explosion pressure of the LDRA at the target density of 0.15 g/cm 3 as a function of charge length. Error bars depicting scatter of experimental data corresponding to each charge length have been included. A rapid initial increasing trend is observed, with pressure amplitudes later converging to a plateau at about 1.90 Kbar. The latter value is interpreted as the explosion pressure of the LDRA mixture, a value that in turn agrees reasonably well with the accepted ratio between detonation and explosion pressure when conditions of total confinement prevail, that is: the explosion pressure being half the amplitude of the detonation pressure g/cc. Average Explosion Pressures Pressure (Kbar) LDRA Charge Length (cm) Figure 5.46: Explosion pressures generated by the LDRA as a function of charge length As for the other LDRA parameters previously measured (VOD, detonation pressure), a considerable scatter of experimental results is observed for the explosion pressure. Several factors can be attributed to explaining such scatter, including the density differences from batch to batch as well as between tests from the same batch. 157

176 When compared to the output from thermo chemical codes such as Cheetah, the experimentally measured detonation and explosion pressures seem to indicate ideal reaction of the LDRA. However, the equations of state for the reaction products used by such codes are semi empirical in nature, that is to say they have been calibrated to experimental results for particular explosive compositions, in most cases comprising condensed, homogeneous, high density explosive mixtures. The LDRA is a particulate, highly heterogeneous low density mixture to which the usual equations of state will not necessarily apply. Thus, the reaction of the LDRA is not necessarily complete as the experimental results seem to indicate. Explosion pressure is an important explosive parameter frequently used for blast design calculations and is usually estimated by simple empirical formulae based on explosive density and experimental VOD. This approach works well for high density homogeneous explosives where detonation reaction approaches ideal conditions, thus, experimental and theoretical VODs tend to converge. However, this is not the case for non ideal products, such as the particulate LDRA mixture. Cunningham (2006) proposed a corrective approach to extend these empirical formulae to non ideal explosives. Cunningham s proposition is based on two sound principles: that pressure is governed by energy and that VOD depends only on the chemical reaction taking place within the reaction zone (i.e. the detonation head). For ideal explosives, where chemical reactions are completed within the detonation head, no corrective action would be required, since 158

177 experimental VOD and ideal VOD tend to match. Non ideal explosives on the other hand, are characterized by an incomplete chemical reaction within the detonation head, as a result of which the measured VOD will be lower than the ideal value. This incomplete reaction however, does not stop; it goes to completion behind the detonation front without contributing to the VOD. As a consequence, the energy that is still being delivered behind the reaction zone and its effect on pressure are not accounted for if the experimental VOD value is used to evaluate explosion pressure. To remedy the above shortcoming and extend the use of the explosion pressure formula to non ideal explosives, Cunningham proposed using the ideal rather than the measured velocity of detonation regardless of whether the explosive behaves ideally or not. The experimental VOD for LDRA at 0.15 g/cm 3 was measured at about 1900 m/s, thus, an explosion pressure of 0.67 Kbar can be calculated from Equation 5.8. The ideal VOD produced by Cheetah is 2480 m/s, in turn resulting in a 1.15 Kbar explosion pressure. The latter value lies closer to the average pressure generated during LDRA testing. In other words, the explosion pressure measurements conducted on the LDRA appears to validate Cunningham s suggested approach, since the experimental pressures of the LDRA are best approximated by using ideal rather than measured VOD values. 5.5 Chapter Summary This chapter describes the experimental evaluation of the LDRA as well as its mixtures with ammonium nitrate and ANFO prills under laboratory conditions. Velocity of detonation was the parameter typically used to evaluate product performance under 159

178 different conditions of confinement, diameter and initiation. Given the influence that the detonation and explosion/borehole pressures of an explosive have on controlling damage, a substantial effort was placed at measuring these two parameters on the LDRA using different experimental techniques. In addition, segregation of the ammonium nitrate fines from the LDRA as well as the LDRA itself from mixtures with ammonium nitrate prills were qualitatively evaluated to gain an understanding of the behavior to be expected under field scenarios. Velocity of detonation experiments on the LDRA show the usual increasing trend with charge diameter in spite of the scatter expected from these heterogeneous mixtures. Laboratory scale experiments indicated that reaction to stable detonation is reached fairly quickly under confined conditions. The LDRA at the target density of 0.15 g/cm 3 under steel confinement proved sensitive to blasting cap initiation, with a critical diameter in the order of 25 mm to 37 mm. This, however, increased substantially when confinement was reduced from steel to cardboard. Subjecting samples of LDRA at the target density of 0.15 g/cm 3 to shaking and vibration resulted in negligible segregation of the ammonium nitrate mineral oil fines, which remain adhered and in between voids of the wetted polystyrene beads. At densities of 0.20 g/cm 3 and above, however, a certain accumulation of fines in the form of layering and at the bottom becomes more noticeable. Mixtures of the LDRA with either standard ANFO or ammonium nitrate prills are expected to segregate given, their very different densities. The reactive nature of the LDRA, however, was shown to synergize the 160

179 mixture and increase its performance relative to mixtures of ANFO and straight polystyrene beads. In addition, the reactive nature of the LDRA will prevent propagation failures of its mixtures with ANFO, regardless of the level of segregation attained, something that cannot be guaranteed for mixtures of ANFO and polystyrene beads alone. Measurement of detonation and explosion pressures demanded much of the effort of this research stage. Attempts to develop a simpler and novel approach to determine detonation pressure proved unsuccessful for low density mixtures, although it remains to be proven for higher density homogenous explosives. The use of PVDF sensors produced reasonable records, however, their amplitudes were suspiciously high, an indication that they may not be the technology best suited to measure dynamic events of heterogeneous mixtures such as the LDRA. The expensive unit cost of the sensors would add up to preclude them as the system of choice. Carbon composition resistors became the gages of choice for measuring both detonation and explosion/borehole pressures. From the various configurations tested, the one producing the most consistent output was by using water as the shock transfer medium embedding the sensor. Average experimental pressures were in the order of 3.6 Kbar, slightly higher than the predicted by ideal codes using semi empirical equations of state, the latter being particularly suited for condensed high density explosives. Measurement of explosion pressure required the design of an explosive chamber capable of surviving the passage of the initial shock wave and holding the products of 161

180 detonation without rupturing during a relatively long period of time, until venting of the gases to the atmosphere occurs. In addition, the designed system allowed the carbon composition sensors and corresponding cable connections to survive throughout the explosive event. The average explosion pressure for the LDRA at 0.15 g/cm 3 was measured around 1.9 Kbar, a value that falls reasonably close to the expected value of half the detonation pressure. 162

181 Chapter 6. Field Application of the LDRA 6.1 Introduction The opportunity to evaluate the novel LDRA explosive at a semi industrial scale within a mining environment was sought throughout the developmental and characterization stage of the product; the effort finally paying off with the implementation of a field project at Chuquicamata Mine, in northern Chile. Figure 6.1 below shows a panoramic view of the open pit mine. Figure 6.1: Overview of Chuquicamata Open Pit Mine looking north Chuquicamata has a long standing slope stability problem that is affecting long term planning, including the transition to underground panel caving methods to extract what presently is a geological resource estimated at 1500 million tons of Cu Mo ore. The open pit operation is scheduled to end production by the year 2013 and to achieve such a milestone, mine advancement is of utmost importance. In turn, to fulfill this advancement, implementation of a blast induced damage control program is imperative 163

182 to produce an adequate berm width and ensure that bench and overall mine slope angles comply with design. These issues are of critical importance and great economic significance in the performance of the operation. The opportunity to introduce a novel low density explosive concept into their blasting practice was considered beneficial for the operation as a whole and captured the interest of management. These benefits included allocating the bigger more efficient drilling equipment used for production blasting to drill larger diameter buffer holes, dedicating the less efficient small diameter equipment to presplit rows only. Implementing this new explosive technology for double bench blasting was also seen as a potential benefit for the operation. In addition, the option of substituting the string charges presently used in presplit for a product amenable to mechanized bulk loading was deemed advantageous from a labor requirement point of view. However, the factor considered to have the greatest economic impact and that would ensure fulfillment of the long term plans, is the elimination of the second stage presently required to blast the full expansion width. Reducing this blasting operation to a single full width blast while maintaining the slope angles will speed up production advancement and ensure compliance with the strategic mine plan. 6.2 Typical Blast Design Applied in Chuquicamata Currently, most mine expansions are conducted in a two phase approach: a production stage, whose goal is to obtain an adequate rock fragmentation, and a controlled blast phase, aimed at reducing damage and increasing slope stability. 164

183 The former consists of m wide benches using 311 mm or 350 mm drill holes while the control blasts consist of m wide benches having a presplit row and two buffer rows usually drilled in 165 mm diameter holes. For some special blast areas, such as ramps, the blast design incorporates larger 270 mm diameter holes for the buffer rows. In addition, to further reduce the possibility of blast induced damage from production blasts, the last production row in some expansion areas is drilled with half its normal spacing and loaded with half its normal charge. Such action will spread energy in a larger number of boreholes but the actual distribution of energy in each borehole will be inadequate from both damage and fragmentation standpoints due to the presence of short concentrated charges and long collars. A typical production pattern using 311 mm holes in mineralized areas is 8 m x 9 m, while larger patterns of about 10 m x 12 m are common in waste material. Production blast patterns for the 350 mm diameter drill used in some waste expansions is 8 m x 16 m. Presplit holes are spaced 1.5 m to 2.0 m apart and inclined 75 degrees from the horizontal. Bench heights range from 17 to 22 meters. Enaex, the largest explosive manufacturer in Chile, is under contract to supply ANFO and heavy ANFO, the two most common explosive products used in production and buffer rows at Chuquicamata Mine. Two types of heavy ANFO products are used: Blendex, a regular ANFO/emulsion mixture and Emultex, a microballoon sensitized explosive. Heavy ANFO densities are selected according to field requirements by adjusting ANFO emulsion ratios. Presplit charges consist of Enaline, an Enaex emulsion 165

184 product packaged in 12 m long polyethylene sleeves having a 10 g/m detonating cord attached alongside. The emulsion product is used in either 38 mm or 32 mm diameters, which are folded into 20 kg boxes for packaging and transport to the blast area. Typical powder factors used in production blasts range from 240 g/ton to 330 g/ton, in buffer holes from 220 to 280 g/ton and in presplit holes from 0.5 kg/m 2 to 0.8 kg/m 2. Figure 6.2 illustrates a plan view of a typical control blast pattern using 311 mm diameter production holes and 165 mm diameter blastholes for both the two buffer rows and the presplit row. 1.5 m LP 6 m 6 m 3 m 6 m 8 m 8 m 16 m Figure 6.2: Plan view of a typical blast pattern used in Chuquicamata Mine Worth noting from the above figure is the reduced spacing used in the last production row which, along with the two buffer rows and the presplit line, present a clear view of the effort placed in controlling excessive damage. 166

185 Initiation of the large diameter production holes is accomplished via 650 g conical shaped high explosive primers connected to a 600 ms down the hole non electric delay detonator. Typically, 17 ms non electric surface delays are used between holes on a given row and 42 ms delays between actual rows. Initiation of the 165 mm diameter buffer holes was achieved with 450 g primers while the presplit rows were initiated from the detonating cord attached to the Enaline explosive. The presplit row is detonated a few days ahead of the production and buffer holes but not always before drilling has taken place. 6.3 Proposed Objectives of the Field Project In broad terms, the objectives pursued for this field project were to evaluate the feasibility of manufacturing the low density product at a semi industrial scale, to evaluate the detonation performance of the LDRA and its mixes with ANFO under typical field scenarios, to analyze the LDRA as a tool to control damage and to gather practical input on the feasibility of using mechanized loading equipment. The field project called for the manufacturing of 6000 kg of LDRA explosive at a nominal density of 0.15 g/cm 3 to be used in a series of experiments, either on its own or as a diluting agent of standard ANFO prills. The LDRA and its mixes were to be loaded in production and presplit boreholes in selected areas of the operation, always in the last row of holes. 167

186 Explosive performance was evaluated by measuring the velocity of detonation using a continuous VOD monitor. The usefulness of the LDRA explosive in reducing damage was assessed by means of near field triaxial vibration sensors. In addition, gas pressure sensors were instrumented in an attempt to investigate possible damage as a consequence of detonation gases penetrating the final wall. Successful results of these initial field experiments would lead in the near future to a second phase, where the appropriate design changes would be implemented in a full expansion blast. 6.4 Large Scale Manufacturing of the LDRA Manufacturing of the six metric tons of LDRA was conducted within the Prillex America facility, the ammonium nitrate plant that Enaex operates in Puerto Mejillones, located a few kilometers north of the city of Antofagasta in northern Chile and about 300 km west of the Chuquicamata Mine. The original plan calling for the manufacturing of LDRA at a density of 0.15 g/cm 3 had to be changed due to the supply of expanded polystyrene beads that did not meet the specifications. Although the polystyrene beads complied with the requested bulk density of ~18 kg/m 3, their diameters proved too large for the ammonium nitrate fines to properly adhere, producing excessive segregation of the fines and a poor quality mixture that was deemed unacceptable for the trials. 168

187 Enaex had available on site a large supply of expanded polystyrene beads used for their own explosive mixtures. Although the bulk density of these beads (25 kg/m 3 ) was higher than desired, their diameters fell within technical specifications and as a result a much better quality mixture was obtained. A compromise needed to be reached in order to obtain a LDRA product with similar performance characteristics (~2000 m/s) as the original LDRA at 0.15 g/cm 3 using the higher density polystyrene beads. In order to achieve this goal, some adjustments to the ingredients proportions were necessary. After a few experiments assessing the uniformity of the mixture and using the velocity of detonation as a performance indicator, it was finally decided to conduct the field experiments using the LDRA mixture at a density of 0.20 g/cm 3 rather than the standard 0.15 g/cm 3 originally sought. During the research stage of the LDRA at Queen s University, the ammonium nitrate fines supplied by Nitrochem were a residual product generated by the manufacturing process of the ammonium nitrate prills; thus, there was no need to grind prills to produce the fines required by the LDRA. At the Prillex America plant, however, the prilling process was such that the fines generated during production were recycled back into the manufacturing stream. As a consequence, a grinding stage of AN prills was needed in order to obtain the ammonium nitrate fines required for the LDRA. In addition, the fines generated from the grinding stage needed to be sized in order to obtain a range within technical specifications. 169

188 6.4.1 Manufacturing Process of LDRA at Enaex Facilities The manufacturing site for the LDRA was improvised around the Nitro Carbon Nitrate (NCN) mixing facility within Enaex s Prillex America plant. Figure 6.3 shows a view of this facility, which is equipped to store and handle 1 ton AN prill bags and contains, within other equipment, a grinder and a stainless steel explosion proof mixer. Figure 6.3: NCN mixing facilities within Enaex s plant The AN maxi bags (~1 ton capacity) were positioned atop a holding frame and discharged into a bin feeding the grinder by means of a rotating screw system. Figure 6.4 illustrate the AN prills discharge and the grinder used for the manufacturing of the AN fines. Figure 6.4: Bin container for the AN prills and the grinder used to produce AN fines 170

189 In order to classify the fines generated by the grinder to the required size range, a double mesh vibrating screen unit, containing a 600 μm and 300 μm sieve aperture, was acquired for the duration of the manufacturing stage. The reason for the upper range size was to exclude larger particles that would tend to segregate from the LDRA mixture, and as a consequence, achieve poor adhesion to the polystyrene beads. The smaller mesh size on the other hand, was used to limit the amount of ultra fine particles that will tend to lump and lengthen the mixing time with the mineral oil solution. It was later observed that the AN fines, as produced from the grinding stage, did not contain a large amount of ultra fine material. Consequently, the decision was taken to eliminate the smaller aperture sieve and replace it with a second 600 μm sieve. Thus, the mesh area was increased by a factor of two and the throughput of the sieving stage greatly improved. Velocity of detonation tests recorded no significant difference in explosive performance when these changes were implemented. Figure 6.5 shows the vibrating screen used during these field trials. Once sieved, the AN fines were then packed in ~40 kg bags and stored for later mixing stages. Figure 6.5: Vibrating screen used to classify the AN fines 171

190 The steps comprising the manufacturing process needed to be organized in such a manner that appropriate use of the available equipment and human resources were realized. In order to reach a reasonable production schedule, the following issues were considered: 1. Grinding stage: The high production capacity of the grinder proved more than sufficient to supply the required quantities of AN fines 2. Classification stage: The vibrating screen was quickly identified as the bottleneck of the process and the problem was readily addressed by substituting the smaller aperture mesh. The unit was manned and operated non stop during the shift, in order to feed AN fines to the mixing stage 3. Mixing stage: This step was carried out in the stainless steel mixer and was, in turn, divided into the following tasks: a. Mixing of AN fines and mineral oil solution to obtain ANFO fines, which were then bagged and stored for later processing. The batch of ANFO fines that could be mixed at any given time in the mixer was ~120 kg b. Wetting of the polystyrene beads with mineral oil solution. This stage was readily achieved given the high surface affinity between mineral oil and polystyrene. Due to the extremely low bulk density of the expanded polystyrene, only ~5 kg of the beads could be loaded into the stainless steel mixer at any given time c. Mixing of the ANFO fines and the wetted polystyrene beads to produce the LDRA mixture. This final step was done immediately after wetting the 172

191 polystyrene beads by adding the required quantity of ANFO fines until the desired density was achieved. Approximately 40 kg of ANFO fines were added to the 5 kg of expanded polystyrene beads in the mixer to produce a batch of ~45 kg of LDRA, which was then discharged into bags of about 10 kg each. Figure 6.6 shows the stainless steel mixer used to mix the AN fines and the polystyrene beads with the mineral oil solution in separate stages and later to mix them together to obtain the LDRA explosive. For safety issues, the electric motor driving the mixing unit was placed outside the shed. Figure 6.6: Stainless steel rotating mixer used during manufacturing of the LDRA Figure 6.7 illustrates the LDRA mixture loaded into 10 kg bags, already labeled and placed over a wooden pallet ready for handling to Enaex s temporary storage facilities at their Prillex America plant in Mejillones. These bags were later shipped and stored at Enaex s explosive s magazine, located in the vicinity of Chuquicamata Mine, in wait for the field trials to take place. 173

192 Figure 6.7: LDRA 10 kg bags ready for storage Quality Control at Manufacturing Plant As for any other explosive product, the manufacturing process of the LDRA underwent two basic quality control (QC) tests: Velocity of detonation, to control explosive performance, and density. VOD measurements were conducted at nearby facilities, following the standard procedures set by Enaex. Figure 6.8 shows a typical experimental setup for a VOD test. Figure 6.8: Quality control test conducted on the LDRA at Enaex s testing grounds 174

193 The VOD experiments were done on every batch and consisted of point to point measurements obtained by inducing a voltage in two open circuit wires placed 15 cm apart in a 1 m long, 150 mm cardboard tube. The passage of the detonation front will shorten the wires and generate a voltage pulse that is recorded with a precision timer. The average VOD was then calculated from this transient time and the known distance separating the wires. The LDRA explosive was initiated with a 450 g primer, which in turn was initiated with a safety fuse cord. The wires were positioned at the opposite end of the initiation point to avoid priming effects. VOD readings for the LDRA mix prepared at 0.20 g/cm 3 averaged 2100 m/s. Density measurements were taken on random samples obtained from the LDRA bags and conducted at the QC laboratories located at the plant. This operation was carried out on a daily basis and for every batch of LDRA being produced. Average densities from the QC readings ranged between 0.18 g/cm 3 and 0.22 g/cm 3, which resemble the variations experienced during the characterization stage. Appendix 6.1 includes some of the spreadsheets containing QC results for both density and VOD. The maximum, minimum and average density and velocity of detonation values obtained for each batch have been highlighted. A descriptive statistical analysis of the 42 data points generated an average value of g/cm 3 with a standard deviation of and a 95% confidence limit of

194 6.4.3 Manufacturing of LDRA/ANFO Mixtures Due to its relatively low density, the LDRA was mainly sought for presplit and/or small diameter buffer rows where a low energy fragmentation is desired. However, for higher density requirements, the LDRA can be mixed with ANFO prills to the desired density, that is, to any value ranging between 0.82 g/cm 3 and 0.20 g/cm 3, without risking propagation failure due to segregation. It was considered important to evaluate the performance of selected ANFO/LDRA mixtures and learn about eventual problems associated with mixing, segregation and loading of these very heterogeneous products. Excluding the batch used for the first field test, where mixing was done on site, the remaining LDRA/ANFO mixtures were prepared at Enaex s plant in Mejillones and then transported to the mine. The density of the mixture was calculated from the densities (ρ) of its two components and their corresponding volume percentages (V). The following expression was used for that purpose: ρ1 V1 + ρ 2 V2 ρ mix = Equation As an example, the initial batch of LDRA/ANFO mixed at a 70/30 volume ratio with corresponding densities of 0.20 g/cm 3 and 0.82 g/cm 3 would result in a mixture density of ~0.39 g/cm 3. Expressed on a weight basis, this same density is obtained at approximately a 35/65 weight ratio. Once the required volumetric (or weight) ratio was provided to Enaex, the mix of LDRA and ANFO was conducted in the stainless steel mixer and discharged into bags. In order 176

195 to prevent any possible chemical attack of fuel oil on polystyrene that could shorten the life of the mixture, mineral oil, as opposed to fuel oil, was used to manufacture the ANFO. Segregation was an unavoidable fact due to the large density difference between the two main ingredients of the mix, and as a result, considerable variation in density was recorded during the quality control tests taken from random samples of the same batch. For example, the 80/20 LDRA/ANFO mix, whose calculated density was 0.31 g/cm 3, had measured values averaging 0.25 g/cm 3. This was an expected behavior, as laboratory experiments conducted at Queen s University and presented in earlier chapters had indicated. 6.5 Field Experiments at Chuquicamata Mine A total of nine experiments were conducted as part of the LDRA field application in Chuquicamata. These tests can be divided into the following two groups: 1. LDRA experiments: consisted of monitoring six blasts using LDRA and/or its mixtures with ANFO. Except for the first experiment, where the LDRA and its mixes were tested on a stand alone experiment, the remaining field experiments were either part of the last row of a production blast or part of a presplit row 2. Baseline experiments: primarily focused on measuring vibration levels induced by the explosive products currently used in the standard blasting practices at the mine, 177

196 with the objective of establishing a damage baseline to which to compare the vibration levels generated by the LDRA Description of Field Experiments As mentioned earlier in this chapter, field experiments were instrumented to monitor the velocity of detonation, used as an indicator of the explosive s performance under field conditions; the Peak Particle Velocity (PPV) of the vibration, in order to evaluate the explosive as a potential tool to control vibration and damage, and the effect of the detonation gases behind the last row, to assess its eventual contribution to damage on the final wall. Conventional and high speed digital video cameras were also available for qualitative evaluation of the blast (i.e. initiation sequence, stemming ejection, gases, etc.). Post blast visual inspection of the area was regularly conducted to confirm proper initiation of the holes and to look for signs of excessive damage. In addition, an assessment program implemented by the mine under the name of Full Control provided for an unbiased comparison of the quality of the wall in some of the expansions where the LDRA product was applied. Results of this program will be addressed in later sections of this chapter. Appendix 6.2 includes relevant geotechnical information of the different expansion areas where the experiments took place. 178

197 Experiments involving the LDRA mixture Overall, six field experiments were conducted at various areas of the Chuquicamata Mine using LDRA at 0.20 g/cm 3 or its mixtures with ANFO. A brief description of these particular tests, including available blast design information, is presented next: LDRA Test 1: The objective of this very first test was to evaluate initiation and propagation behavior of the LDRA and two of its mixtures with ANFO under the typical diameter and initiation conditions used at the mine and compare the levels of surface damage produced by mixtures having different loading densities. Test 1 conducted in Expansion 42 East in waste rock was the only stand alone experiment that was not part of an operational blast. It consisted of two sets of three holes each for a total of six blastholes. One set consisted of 165 mm diameter holes, normally used for buffer rows, while the other set was drilled at 311 mm, the typical diameter used for production holes. In turn, each set was loaded with 100% LDRA (0.20 g/cm 3 ), 70/30 LDRA/ANFO (0.37 g/cm 3 ) and 50/50 LDRA/ANFO (0.49 g/cm 3 ) respectively, all ratios expressed on a volume basis. All six blastholes in this particular experiment were monitored for VOD. Near field vibration and gas pressure sensors were placed behind each group of blastholes. Conventional digital video and photography were taken during this experiment. Figure 6.9 shows, in mine coordinates, the location of the blast and recording holes for this particular test, while Table 6.1 summarizes the corresponding loading data. 179

198 GS 1 Geo NORTH (m) GS 2 Geo 2 Geophones Gas Sensors 165 mm (6 1/2 in) 311 mm (12 1/4 in) EAST (m) Figure 6.9: Blast and monitoring hole coordinates corresponding to LDRA Test 1 Table 6.1: Blasthole loading features for LDRA Test 1 Hole Coordinates Density Φ Hole Length Column Length Stem Delay Unit load Total Mix Nº East North g/cm 3 mm m m m ms kg/m kg

199 LDRA Test 2: The objective of this experiment was to evaluate a 70/30 LDRA/ANFO mix (volume ratio) as a possible alternative to ANFO and or Blendex (heavy ANFO) used in their simulated buffer row in Expansion 47 West, where this test took place. The calculated density of the LDRA mixture was 0.37 g/cm 3. The experiment consisted of loading the LDRA/ANFO mix into seven holes of the last row of a production blast. Standard blasting practice in the area called for the last row to be drilled with the same 350 mm bit as for production holes, but at half its spacing, and charged with approximately half its load, all in an attempt to simulate a buffer row to control excessive damage. A total of 2520 kg of LDRA/ANFO mixture was applied in this test. Blasthole lengths averaged 18 m with about 9 m stemming. VOD was monitored in two blastholes. Two vibration and two gas pressure sensors were used in the experiment. In addition, a high speed digital video camera, as well as conventional digital video and photography, were used. Figure 6.10 shows the blast area with the location of the LDRA blastholes and monitoring instrumentation. 181

200 Gas and Vibration sensors ANFO LDRA Gas sensor PPV sensor Figure 6.10: Blast and monitoring holes coordinates corresponding to LDRA Test 2 LDRA Test 3: The objective of this test was to simulate a buffer row using 311 mm diameter holes loaded with a LDRA/ANFO mix instead of the usual control practice for this expansion area, usually consisting of two 165 mm diameter buffer rows loaded with Blendex (heavy ANFO) mixtures as for Baseline Test 2. The experiment was set along the ramp for Expansion 42 East, in waste rock. It comprised a row with seven blastholes loaded with the LDRA/ANFO 80/20 mixture (volume ratio) having an estimated density in the order of 0.31 g/cm 3. Hole lengths ranged between 15 m and 21 m deep, spaced about 5 m to 7 m apart. Five out of the seven blastholes were monitored for VOD. In addition, one vibration and two gas pressure sensors were placed at different distances behind the test row as 182

201 shown in Figure Original plans called for this test row to be the last of the blast, although, due to unforeseen operational issues regarding drill availability, it ended as a middle row of a much larger blast. Unfortunately, unexpected system problems in recording instruments prevented from gathering VOD, vibration and gas pressure events. Nevertheless, post blast survey of the area found no indications of malfunction of the LDRA/ANFO mixture. Geo PS m 6 m 6 m PS 1 6 m 3 m 5.8 m Spacing 7.2 m 5 m 7.1 m 6.8 m 5.2 m 5.3 m Depth 15.2 m 16 m 17.4 m 20 m 18.5 m 18.4 m 21.4 m LDRA/ANFO 80/20 Vol Blasthole initiated 200 ms after last production row Figure 6.11: Diagram of blast and monitoring holes location corresponding to LDRA Test 3 LDRA Test 4: The main objective of this test was to evaluate the behaviour of the LDRA at a density of 0.20 g/cm 3 in 165 mm diameter presplit holes, as an alternative to Enaline, the standard cartridge product used for presplit. Test 4 was conducted in Expansion 40 East, in mineralized rock. It consisted of loading a section of 10 presplit holes (out of a total 27 holes) with 100% LDRA (0.20 g/cm 3 ), with the remaining 17 holes loaded with the Enaline, the standard 36 mm cartridge product use for presplitting. Blastholes were 15 m long, spaced 2 m apart and inclined 75º to the horizontal. Average column length was about 9.5 m and no stemming was used. 183

202 Electronic detonators were used to ensure simultaneous initiation of LDRA blastholes. The first LDRA blasthole was initiated 10 ms prior to the remaining holes, to allow recording the VOD of two blastholes, considering the single channel capability of the portable VOD recorder. This experiment was instrumented with two VOD holes, one geophone (placed on the surface due to the unavailability of a measuring hole) and one gas pressure sensor, in addition to conventional digital video and photography cameras. Figure 6.12 illustrates the location of the blastholes and monitoring devices. Explosive Conventional (Enaline) LDRA at 0.20 g/cm 3 Pressure Sensor (PS) Geophone Geo PS Figure 6.12: Presplit and monitoring holes coordinates corresponding to LDRA Test 4 Table 6.2 next, depicts relevant loading information for the LDRA blastholes corresponding to the presplit blast of LDRA Test

203 Table 6.2: Blast design data for LDRA Test 4 Hole Coordinates Diameter Hole Length Column Length Stem Unit load Total LDRA Nº East North [mm] [m] [m] [m] [kg/m] [kg] LDRA Test 5: As with LDRA Test 4, this experiment evaluated the behaviour of the LDRA at 0.20 g/cm 3 in a standard presplit blast. LDRA Test 5 was set in Expansion 41 East, in waste rock. The presplit blast contained a total of 71 holes, with a section of 24 holes loaded with LDRA and the remaining 47 with Enaline. Blasthole length was about 20 m, spacing 1.5 m and inclination 75º to the horizontal. Column length was in the order of 12 m and no stemming was used. In order to reduce the possibility of excessive damage arising from the simultaneous detonation of the presplit, the 24 hole section loaded with LDRA was divided in three groups of eight blastholes each. Electronic detonators were used to achieve simultaneity within each group, to delay one group from the other and to provide a 20 ms delay separating the first hole from the remaining ones in order to allow recording of the VOD in two consecutive blastholes. 185

204 This experiment was instrumented by measuring the VOD in the first two blastholes of the preslpit row loaded with the LDRA and by recording vibration levels with one geophone and gas pressure with two sensors placed behind the preslpit row, as indicated in Figure In addition, conventional digital video and photography was used. Explosive Conventional (Enaline) LDRA at 0.20 g/cm 3 Geophone PS 2 Pressure Sensors (PS) Geo PS 1 Figure 6.13: Presplit and recording holes corresponding coordinates to LDRA Test 5 Table 6.3 illustrates relevant blast loading information corresponding to the 24 LDRA blastholes of LDRA Test

205 Table 6.3: Blast design data for LDRA Test 5 using LDRA at 0.20 g/cm 3 Hole Coordinates Diameter Hole Length Column Length Stem Unit Load Total LDRA Nº East North [mm] [m] [m] [m] [kg/m] [kg] LDRA Test 6: Like the LDRA Test 2, this test was conducted in Expansion 47 West in waste rock. Standard control practice called for holes of the last row to be spaced half the distance and to be charged half the load of regular production holes, while maintaining the 350 mm diameter used for production. Twelve blastholes of this simulated buffer row, usually loaded with 2 3 m of Blendex or ANFO, were charged with about 9 m of LDRA at 0.20 g/cm 3. Two LDRA blastholes were monitored for VOD. In addition, one vibration sensor and one pressure sensor were used along with conventional digital 187

206 video and photography. Figure 6.14 illustrates LDRA Test 6 location while Table 6.4 summarizes blast loading data for the LDRA blastholes. Explosive Conventional (Blendex) LDRA at 0.20 g/cm 3 Geophone Pressure Sensor (PS) PS Geo Figure 6.14: Mine coordinates corresponding to LDRA Test 6 Table 6.4: Blast design data for LDRA Test 6 Hole Coordinates Diameter Hole Length Stem Volume Factor Column Length Unit Load Total LDRA Nº East North [mm] [m] [m] [m2] [m] [kg/m] [kg]

207 Baseline Experiments The three baseline experiments conducted at Chuquicamata were considered representative of typical blast control practices at the mine and were used to establish a baseline for damage comparison with the LDRA product. These tests were conducted by a third party (ASP Blastronics) prior to and after the ones corresponding to the LDRA product. Appendix 6.2 depicts the approximate mine location of the baseline experiments while available blast design data and damage prediction models have been included in Appendix 6.4. A few comments on each baseline experiment follow: Baseline Test B1: This experiment was conducted in Expansion 40 East and consisted of monitoring peak particle velocities behind the 270 mm diameter buffer row typically used in this area of the mine as a damage control practice. Two geophones were used for monitoring vibration levels. Baseline Test B2: This experiment was conducted in Expansion 41 East and like Baseline Test B1, its objective was to establish damage levels behind the control section of the blast, which consisted of two 270 mm buffer rows and a 165 mm diameter presplit row. The test was instrumented to monitor VOD in two blastholes, in addition to two vibration sensors and two gas pressure sensors. One vibration and one gas pressure sensor were positioned in front of the existing presplit, in an attempt to evaluate the effectiveness of the presplit cut in reducing the vibration from the blast. Baseline Test B3: This test was conducted in Expansion 47 West, the same as for LDRA Test 2 and Test 6. It was instrumented with one vibration sensor to monitor peak particle 189

208 velocity levels behind the buffer row in order to asssess damage produced by the standard control blasting practice. This consisted of reducing the explosive load and blasthole spacing for the last row, utilizing the same 350 mm diameter blasthole as for production Velocity of Detonation Experiments The detonation performance of most commercial explosives is greatly affected by external or field factors, such as borehole diameter, degree of confinement, priming conditions, presence of water, etc. These factors, more often than not, differ from the controlled conditions that characterize most laboratory scale experiments; so for a given explosive product, different behaviors are to be expected. For this reason, evaluating the performance of the LDRA explosive under the conditions in which it is applied at the mine is of critical importance, and the velocity of detonation, in turn, is the most practical way of evaluating such performance Explosive Loading and Measurement of VOD A continuous VOD monitor with a low resistance per unit length coaxial cable (nominal 10.8 Ω/m) was used to monitor selected blastholes. The shunted end of the coaxial cable was taped to the explosive primer and lowered into the first hole whose VOD was to be measured. The cable was then extended to chain link the other hole(s) by running it through a wire loop taped to the primer so it slid down until it reached the bottom of the hole. At this point, the primer was raised about one meter, to ensure its placement within the explosive column rather than at the very bottom of the charge. As illustrated 190

209 in Figure 6.15, the LDRA product was manually discharged into the blastholes directly from the 10 kg bags delivered to the site. Figure 6.15: Manual loading of LDRA into test holes Density readings of the LDRA were taken in the field from randomly selected bags as a quality control measure. Density fell well within the expected range and more often than not, right on the mark (0.20 g/cm 3 ) as shown in Figure Figure 6.16: Field control of LDRA density The LDRA/ANFO mixtures prepared at Enaex s plant in Mejillones were also discharged from the bags, however the loading procedure for the initial test was 191

210 improvised on site by filling two 20 litre pails with the required volume of each component and simultaneously emptying them, a pair at a time, into the blast hole. This is shown in Figure 6.17 for a 50/50 volume ratio into a 165 mm diameter hole. Figure 6.17: Bucket loading of LDRA (blue) and ANFO (grey) for a 50/50 volume mix The mixing of both components would occur within the hole, as the two flows meet at the collar and fall down the column. Although crude, the method represents a worstcase loading scenario from which useful input was gathered regarding the effect of segregation on propagation stability and the viability of using mechanized loading equipment. Figure 6.18 shows the VOD traces generated by a 10 m column length of LDRA at 0.20 g/cm 3 in two 165 mm diameter presplit holes having a 10 ms delay accurately provided by electronic detonators. In spite of the observed noise, the slope on the distance time record is perfectly discernible and the VOD easily obtained. Records showed stable propagation along the entire column length. 192

211 g/cc in 165 mm Presplit holes Distance (m) Hole # 1 = 2023 m/s Hole #2 = 2109 m/s Time (ms) Figure 6.18: VOD records of two presplit holes loaded with LDRA at 0.20 g/cm 3 The electrical noise, seen as zero distance spikes on the VOD record, is attributed to a poor shorting of the coaxial cable as a consequence of the low crushing pressures produced by the explosive. This noise tended to decrease for larger diameters or higher density mixtures, as indicated in Figure 6.19 for a 350 mm hole loaded with a 70/30 volume ratio of a LDRA/ANFO mix with a density of 0.37 g/cm g/cc in 350 mm holes Distance (m) VOD = 3311 m/s Time (ms) Figure 6.19: VOD trace of a LDRA/ANFO at 0.37 g/cm 3 in 350 mm diameter hole 193

212 Analysis of VOD Experimental Results Table 6.5 summarizes the VOD data obtained throughout the field trials for both, the LDRA and the LDRA/ANFO mixtures. Table 6.5: Summary of VOD measurements conducted during field trials at Chuquicamata Explosive Volumetric ratio (%) Density (g/cm 3 ) VOD (m/sec) Diameter 165 (mm) 311 (mm) 350 (mm) LDRA LDRA/ANFO 70/ LDRA/ANFO 50/ LDRA/ANFO 70/ LDRA LDRA LDRA LDRA LDRA The expected VOD density and VOD diameter trends are clearly seen in Figure 6.20 and Figure 6.21 respectively. The odd value observed in the VOD diameter graph for the LDRA at 0.20 g/cm 3, whereby the velocity of detonation recorded at 350 mm is lower than the one at the smaller 311 mm diameter hole, can be attributed to either the effect of higher confinement at the smaller diameter tests overcoming the diameter effect or by the fact that the larger diameter experiment was conducted in heavily fractured, less competent rock, offering lower confinement conditions that will, in turn, affect the degree of reaction of the explosive and therefore its VOD. 194

213 VOD-Density VOD (m/s) Density (g/cm (g/cc) ) 165 mm 311 mm 350 mm Figure 6.20: VOD Density data for the various diameters tested at Chuquicamata 3800 VOD-Diameter 3400 VOD (m/s) Diameter (mm) 0.20 g/cm g/cc g/cm g/cc g/cm g/cc 3 Figure 6.21: VOD Diameter relationship for the different LDRA mixtures It is worth noting the small difference in VOD recorded for the LDRA at 0.20 g/cm 3 as the charge diameter is increased from 165 mm to 311 mm and beyond. This is a strong indication of the capability of the low density explosive to deliver much of its maximum available energy at relatively small charge diameters. According to the thermo chemical 195

214 code Cheetah, the ideal VOD (i.e. infinite diameter) for the LDRA at 0.20 g/cm 3 is in the order of 2600 m/s. This ideal velocity is not far off the 2429 m/s recorded at 311 mm, an indication that the LDRA is behaving much closer to ideal than anticipated for such a heterogeneous mixture. A similar behavior can be said for the LDRA/ANFO mixtures where the ideal velocities (3206 m/s and 3575 m/s for 0.37 g/cm 3 and 0.50 g/cm 3 respectively) compare favorably with the experimental values recorded during the field trials, an indication also of the synergy provided by the LDRA within the mix when it is compared to other inert bulking agents. From analyzing the distance time records obtained during these field trials, it can be concluded that the LDRA at 0.20 g/cm 3 reaches stable velocity regimes fairly quickly; in other words, the distance of run to a stable velocity regime is relatively short, regardless of the charge diameter and priming weight. This behavior has practical implications for bottom priming: the explosive is capable of liberating its maximum energy in a short run, ensuring proper fragmentation where needed the most. Similar behavior was observed with the LDRA/ANFO mixtures at diameters of 165 mm or larger as seen in Figure 6.22 for a 50/50 LDRA/ANFO mixture. This behavior is characteristic of over priming. 196

215 LDRA/ANFO (50/ g/cc- 165 mm Distance (m) VOD = 3405 m/s Time (ms) Figure 6.22: VOD trace for a LDRA/ANFO mixture at 0.49 g/cm Gas Pressure Experiments Measurement of pressures generated by the expanding gases behind the last row is a technique that has been recently used in an attempt to identify gas penetration and its eventual contribution to damage. The technique has proven useful mostly as a comparative method of evaluating different explosives and blast designs, rather than as a quantitative approach towards predicting slope failures. For example, measuring gas pressure at both sides of a presplit row will give a qualitative indication of the presplit efficiency reducing gas penetration and protecting bench slopes. The experimental results reported by most researchers (LeJuge et al, 1994; Ouchterlony et al, 1996; Brent et al, 1999) revealed the presence of a negative pressure pulse, followed by a positive pulse when confinement and other blast design conditions are met. Negative pressure pulses have been attributed to vibration induced strain. This strain is 197

216 responsible for generating new fractures or extending existing ones, resulting in dilation or swelling (i.e. elastic rebound) around the hole. As a result of this process, the measuring hole will increase its effective volume, creating a vacuum or suction that will then register as a negative pulse. Given that the swelling was seen to occur before the flow reaches the sensor, it is more likely that gas penetration is the result rather than the cause of this dilation. Figure 6.23 shows an example of an all negative gas pressure record produced by a sensor located 5 m behind a contour line. Figure 6.23: Negative air pressure trace, indicative of damage, at 5 m from contour blast (Ouchterlony, Nie et al, 1996) Positive pressures on the other hand, are attributed to the gas penetrating and pressurizing the measuring holes. There is consensus within the various researchers that these overpressures tend to occur in monitoring holes close to blastholes and in particular under conditions such as presplit and crater blasting, where a higher degree of confinement exists. Figure 6.24 shows a typical example of a positive pressure trace produced behind a presplit row. Worth noting is the short live initial negative pulse, indicative that a certain degree of damage occurs prior to gas pressurization. 198

217 Figure 6.24: Positive pressure trace, indicative of gas pressurization, at 2.8 m behind a presplit (Ouchterlony, Nie et al, 1996) Positive and negative pressure amplitudes vary from author to author and are very much field dependent, with reported values ranging from +280 kpa at a distance of 1 m for a fully confined crater shot to 78 kpa at 8 m behind the last row and with durations in the order of 2 5 seconds. Moreover, positive and negative pressure pulses were reported at distances of up to 20 m behind the blastholes (Brent et al, 1996) implying that both fracture dilation and gas penetration, were present over these distances. It is the generally accepted within the research community (LeJuge et al., 1994; Ouchterlony et al., 1996; Brent et al., 1999) that fracture dilation due to vibration is more significant from a damage point of view than the positive gas pressures associated with gas penetration; nevertheless, the latter can have influence on the stability of slope blocks Gas Sensor Assembly and Field Implementation Each of the six field experiments conducted in Chuquicamata was instrumented with gas pressure sensors. The sensors, supplied by Advanced Custom Sensors Inc. (ACSI), 199

218 consisted of a small piezo resistive silicon elements designed for low pressure applications in harsh environments. The sensing element was embedded in silicone oil, which acted as a pressure transfer medium and was isolated from the external environment by a stainless steel diaphragm, all housed within a ¼ NPT stainless steel port fitting. The selected sensor required a clean and uniform 5 to 10 DC excitation voltage to operate properly, which was supplied via a 9 DC battery through a 5 DC voltage regulator. All electrical components were mounted within a protective plastic box that included an ON OFF switch to preserve battery charge as well as a BNC connector for the coaxial cable running to the recording unit. The whole assembly, as illustrated in Figure 6.25, was designed and assembled at Queen s University. 5 V regulator Sensor Switch BNC 9 V battery Figure 6.25: Sensor assembly used to monitor gas pressure Technical specifications of the sensor include a full scale voltage of 200 ± 40 mv maximum operating pressure of 689 kp, the capability to record vacuum (negative 200

219 pressures) up to 100 kpa, a linear output throughout its measuring range and a relatively fast response (< 1 ms) to the application of pressure. Calibration of the pressure sensor was required in order to convert its voltage output into pressure units. This task was readily accomplished at Queen s University laboratories by subjecting the sensor to a semi static gas pressure provided by a nitrogen cylinder and a quick release valve. Voltage output was recorded with a voltmeter in pressure increments of 35 kpa and results were linearly fitted to obtain the calibration equation relating voltage and pressure. Figure 6.26 depicts a graph with the calibration curves corresponding to all ten sensors assembled for these field experiments, where the linearity of the output and the similarity of their responses are readily observed Pressure Sensor Calibration Curves Pressure (KPa) Voltage (mv) Figure 6.26: Calibration curves of the gas pressure sensors used during field experiments Figure 6.27 illustrates the experimental setup used during these field trials. 201

220 Figure 6.27: Experimental setup used for measuring gas pressure behind the last row At the most, two gas pressure sensors were placed within monitoring holes behind the last row. A 3 m long, 76 mm diameter PVC tube was cemented in the hole to seal off the section of weathered rock usually present at the collar. The plastic tube was capped at the top end and left opened at the bottom end to allow gases in. Once sealed, the measuring borehole acts as a pressure chamber, collecting gases through existing or created fractures and pressurizing the sensor. The sensor assembly itself remained outside, connected to the monitoring hole via a high pressure hose clamped to a metal fitting, which in turn was threaded to the tube s lid. This way, as it happened in most cases, the sensor could be recovered and used for upcoming experiments. Figure 6.28 is a photograph of an actual experiment illustrating the high pressure hose connecting the sensor assembly to the PVC tube, the latter cemented in a 311 mm diameter measuring hole. 202

221 Figure 6.28: Gas pressure sensor assembly in a 311 mm hole Gas Pressure Experimental Results The small number of readings and the variety of loading scenarios of these field trials prevented establishing a correlation between gas pressure and distance or assessing on a relative basis the damage caused by the detonation gases. Therefore, only descriptive comments on the experimental approach will be presented in this section. The effort, however, proved the functionality of the measuring system and provided useful input in lieu of future tests. The location of the gas sensors within each experiment was illustrated in Section 6.5 of this chapter. The pressure traces resembled, in amplitude and shape, those reported by previous investigators; although, none of them consisted of negative pulses alone. Traces typically showed an initial negative pulse of relatively short duration, indicating that a certain degree of damage was occurring at the sensor s location, followed by a positive phase lasting several seconds, the latter indicative of detonation products 203

222 reaching and pressurizing the measuring hole through existing or blast induced fractures. Figure 6.29 illustrates a typical example of a gas pressure trace obtained during LDRA Test 1, where the short initial negative pulse followed by a long lasting overpressure is clearly seen. This particular record corresponds to the gas sensor placed behind three 165 mm diameter blastholes, each loaded with different low density mixes. The closest of the three blastholes was 2.4 m from the sensor and loaded with LDRA at 0.20 g/cm 3 as specified in Table 6.1. Initiation times for each of the blastholes, as obtained from the VOD record in Figure 6.30, have been included in the graph. Hole 3 (144 ms) Hole #1 ( 125 ms) Hole 2 (0.2 ms) Figure 6.29: Gas pressure trace from 165 mm blastholes. LDRA Test 1 (MicroTrap record) The MicroTrap monitor used during these field trials has a dedicated channel for the VOD and 4 channels to record voltage from a variety of sensors (i.e. gas pressure), all events being recorded on the same time scale. The system was set to trigger by the VOD 204

223 event, providing therefore the reference time of initiation ( 125 ms). Using the gas pressure history and the initiation times from the VOD record, it is possible to evaluate the blasthole(s) responsible for the pressure event. As an example, from Figure 6.30 the initiation of the first blasthole occurs at 125 ms while the pressure sensor from Figure 6.29 registers the arrival of a negative gas pulse a few milliseconds later, which lasts about 40 ms. Given the delay established for the remaining two holes (130 ms each), only the very first and closest blasthole can be responsible for the negative pulse. On the other hand, according to Figure 6.29 the much longer positive pressurization pulse peaks around 650 ms, long after all three holes had detonated. Therefore, it is inconclusive whether the closest blasthole or all three of them are contributing to the pressurization phase; although the trace seems to indicate overpressure steps coinciding with the detonation of the last two blastholes. Figure 6.30: Timing of VOD traces corresponding to LDRA Test 1 (MicroTrap record) 205

224 Figure 6.31 next shows the pressure trace corresponding to LDRA Test 6, where the last row consisted of 350 mm diameter blastholes loaded with LDRA at 0.20 g/cm 3. The sensor was positioned 6.5 m from the closest blasthole. A direct comparison of the overpressure (positive) amplitude with that of Figure 6.29 cannot be made due to differences in load, explosive type and distance. However, the longer duration of the negative pressure phase (~300 ms) is an indication that a higher level of damage from the action of gases is to be expected for the same explosive type at larger diameters. Figure 6.31: Gas pressure trace 6.5 m behind a 350 mm hole loaded with LDRA at 0.20 g/cm 3 corresponding to LDRA Test 6 Figure 6.32 and Figure 6.33 show the pressure traces corresponding to LDRA Test 5, which were recorded behind a presplit row consisting of 24 holes loaded with LDRA at 0.20 g/cm 3 and 45 holes loaded with Enaline, the standard emulsion cartridge used for 206

225 pre splitting at Chuquicamata. The two gas pressure sensors were positioned 3.1 m behind the corresponding sections of LDRA and Enaline. Figure 6.32: Pressure at 3.1 m behind presplit section loaded with LDRA at 0.20 g/cm 3 Figure 6.33: Pressure at 3.1 m behind presplit section loaded with Enaline emulsion cartridges 207

226 Interpretation of data is somehow inconclusive given that only two measurements are available for comparison, however the output indicates that both LDRA and Enaline, generate an initial negative pulse, an indication that a certain degree of damage is occurring by the sensor s location. Also worth noting are the amplitudes and durations of the pressure pulses produced by each explosive. While the LDRA shows a longer duration for both the positive and negative phases, the emulsion product registers much higher amplitudes for both phases. Unfortunately, since there are no damage criteria associated with pressure amplitude or duration, it is not possible to evaluate, either quantitatively or qualitatively, which explosive is causing more damage. The fact remains however, that both products generate damage some 3.1 m into the rock, a value that should be taken as a design parameter for control Conclusions of Gas Pressure Results The positive pressures observed during these field trials are a reflection of the blast designs applied in Chuquicamata, which in most cases resembled more a choke blast than a true free face blast. As a consequence, confinement conditions similar to those of a presplit and or crater blast are created and the expected overpressures thus generated. In spite of the relatively small number of measurements conducted in Chuquicamata, some conclusions can be derived from the experimental results. The damage radius expected for a LDRA at 0.20 g/cm 3, loaded into 165 mm blastholes, could reach at least 2.4 m (Figure 6.29), while the corresponding to the 350 mm blastholes at least 6.5 m (Figure 6.31). Standard presplit practices using Enaline explosive cartridges resulted in 208

227 damage radii larger than 3.1 m (Figure 6.33). There are strong indications that the use of LDRA at a density of 0.20 g/cm 3 will generate much lower amplitude pressures but longer duration pulses than Enaline cartridges (Figure 6.32 and Figure 6.33). In order to minimize the possibility of damage arising from dilation (negative pressure due to vibration) and gases (positive pressure due to gas penetration), several blast design issues need to be addressed. Timing and initiation sequence is of particular importance to prevent synergies between detonating blastholes and to provide the necessary relief, allowing the energy remaining after fragmentation to be spent doing useful work (throw) rather than gas penetration, which will tend to generate instabilities or damage to the wall. Proper burden design is fundamental to reduce gas confinement, which is considered the primary factor controlling the extent of gas penetration and the peak levels of positive pressure. Whenever possible, the use of two free faces is justified in order to enhance burden relief, therefore, reduce gas penetration and backbreak. Reduction of vibration levels by implementing proper initiation timing and controlling weight, diameter, density and/or degree of decoupling of the explosive charge, should be addressed in every blast design program. The implementation of a presplit row to allow an effective venting of gases, ensuring no stemming is used and the blast is properly designed to prevent the presplit from being the cause of the problem rather than its solution. Consideration should also be given to avoid excessive delays between 209

228 the presplit and production blasts that may result in loose material plugging the split and reducing its ability to control damage Near Field Vibration Experiments The most widely applied engineering model predicting near field vibrations is the one proposed by Holmberg and Persson (1978), described in earlier chapters and whose general form is presented next for convenience. PPV α H α γ h = k β 2 2 α R + R Tanφ h 0 2 [ 0 ( 0 ) ] Equation 6.2 where k, α and β are site specific constants, γ is the charge per unit length and the remaining factors related to the geometry of the charge with respect to the measuring point as sketched in Figure Xs Ro Xo Explosive Sensor H dh h φ Figure 6.34: Geometric parameters used in Holmberg and Persson s equation 210

229 Assuming β = 2α and integrating the previous equation, the following practical expression is obtained: PPV γ H + X = k a tan R0 R S 0 X 0 X + a tan 0 X R 0 S α Equation 6.3 Where PPV is the peak particle velocity (mm/s), k the rock factor also called velocity factor, α the attenuation factor, γ the linear charge (kg/m), Ro the distance from measuring point to center of incremental charge (m), H the column length (m), Xs the stemming length (m) and Xo the depth of measuring point (m). As the previous formulas imply, near field vibration levels estimated by the Holmberg and Persson s (H&P) model are greatly influenced by the linear charge concentration rather than the charge weight per delay, as is the case for far field applications where the explosive acts as a point source rather than a finite column length. Particle velocity is associated with strain, itself an indication of damage, according to the following relationship: PPV = ε V p Equation 6.4 where PPV is the peak particle velocity, ε the induced strain and Vp the compressional wave velocity. 211

230 Assuming elastic behavior (Hooke s law) and a brittle failure mode for the rock, it is possible to determine the critical particle velocity (PPVcritical) that can be withstood by the rock before tensile failure occurs. This is given by the following equation: PPV critical T V p = σ Equation 6.5 E where σt is the dynamic tensile strength of the rock and E the Young Modulus. Geotechnical data made available by mine personnel related to the areas where the field experiments took place are presented in Table 6.6 along with the corresponding calculation of PPVcritical according to Equation 6.5 above. Table 6.6: PPVcritical for the mine expansions where LDRA field experiments were conducted Geotechnical Unit σt Vp E PPVcritical Mine Expansion [MPa] [m/s] [GPa] [mm/s] Granodiorite Elena E, 42E Granodiorite Este Granodiorite Fortuna W Moderate Shear Zone Intense Shear Zone West Porphyry East Porphyry, Potasic Alteration E, East Porphyry, Chloritic Alteration E, 41E East Porphyry, Serecitic Alteration E, 47W Serecitic Quartz Rock E, 47W Metasedimentary Limestone Dry Metasedimentary Sandstone Saturated Metasedimentary Sandstone Saturated Metasedimentary Limestone E, 42E 212

231 Summarizing, the H&P model allows estimation of the PPV generated in close proximity to a blasthole, where fracturing occurs as a consequence of the high strain levels induced by a detonating charge. Using the PPV strain relationship, a damage criterion can be established by calculating the PPVcritical based on well known geotechnical rock parameters. Fracture or damage at various distances around the blastholes can then be estimated, which in turn provides a good first approximation of the expected level of back break behind the blast and the minimum required standoff between the back row and the toe of the final wall. As mentioned in earlier chapters, the H&P model assumes an infinite VOD, implying that the vibration contributed by each elemental charge the column has been divided into, arrives at the measuring sensor simultaneously. Consequently, the model will tend to overestimate the actual field results. Although different in geometry, the extent of damage generated by the H&P model proved very similar to alternatives such as the Seed Wave model (Starfield et al, 1968). If one adds its simplicity of use and conservative output, it is easy to realize why it has become a popular engineering tool to assess and predict damage. To facilitate the analysis and presentation of data, the Holmberg and Perssons expression presented in Equation 6.2 can be conveniently simplified as follows: PPV α = k (HPF) Equation

232 where HPF refers to the Holmberg and Persson Factor comprising all the geometric aspects related to the explosive s loading configuration and distance to the measuring hole. The factors k and α representing the attenuation characteristics corresponding to a certain sector of the mine are determined from a least square power regression analysis of experimental data. When the data are plotted on a logarithmic scale, the regression becomes linear Field Implementation of Vibration Experiments Vibration sensors arranged in a triaxial configuration and mounted on an aluminum angle were embedded in epoxy poured into a 51 mm diameter, 15 cm long plastic coupling used as a mold. This coupling was then joined to a series of 3 meter long plastic tubes, with the sensors cables running through. The assembly was then lowered into the measuring hole to about 14 meters depth and cemented in place with concrete delivered to the blast site by truck. The photographs included in Figure 6.35 illustrate both the sensor assembly and the placement of the tubes in the measuring holes. Figure 6.35: Vibration sensor embedded in epoxy and setup of PVC tube in the blasthole 214

233 Sensor cables were extended and connected to a Blastronics BMX digital recorder, the resulting orthogonal vibration signals analyzed and the parameters of interest (maximum peak particle velocity and arrival times) determined. The data acquisition rate for recording these particular events was set at Hz (10 KHz). At the most two measuring holes were drilled about one burden behind the last row on each field experiment. Figure 6.36 illustrates the experimental setup. Figure 6.36: Diagram showing field implementation of vibration measurements The distance between the sensor and the explosive column was measured on the surface and later double checked with the survey coordinates provided by the mine. For the presplit holes, drilled at a 75º angle to the horizontal, the sensor to explosive distance was calculated, based on sensor depth, drill angle and surface distance. From a total of nine experimental blasts conducted at Chuquicamata, three were monitored behind buffer rows loaded with the standard explosives used during their 215

234 blasting practices while the remaining six trials were monitored behind holes loaded with LDRA or LDRA/ANFO mixtures. The main objectives of the initial three trials were twofold, to establish a baseline to which to compare the vibration levels produced by the LDRA, and to generate a prediction model applicable to the sector of the mine where the blast took place. Of the six blast experiments monitored using LDRA or its mixes with ANFO, Test 2 and Test 3 failed to record vibrations due to instrumentation malfunction. Moreover, no predictive model was considered for the presplit experiments involving Test 4 and Test 5, since the nature of the blast design itself calls for simultaneous initiation (attained with electronic detonators) of closely spaced holes, preventing therefore the unequivocal identification of individual blastholes from the vibration record. However, the PPV signal generated by the first presplit hole from Test 5, detonating 10 ms prior to the reminder of the presplit blast was considered useful data. Appendix 6.3 and Appendix 6.4 contain vibration data relevant to LDRA tests and the three baseline experiments respectively, including data used for the analysis and generation of the predictive models where possible Vibration Analysis of LDRA Trials As implied in the previous section, from the experiments conducted with LDRA and its mixtures with ANFO, only Tests 1 and 6 provided sufficient vibration data to contribute to a predictive damage model. 216

235 Table 6.7 summarizes the basic information associating the maximum PPV recorded during these LDRA experiments with the characteristics of the blasthole responsible for generating the corresponding vibratory event. Table 6.7: Maximum PPV recorded during field trials using the LDRA and mixes with ANFO Diameter Wt Dist Density PPV (max) Test Geo Explosive mm kg m g/cm 3 mm/s 1 LDRA/ANFO LDRA/ANFO LDRA LDRA As an example of the steps followed to evaluate experimental data and generate the prediction models, information corresponding to LDRA Test 1 is discussed next. For convenience, the locations of the blastholes and measuring holes housing the vibration sensors are duplicated in Figure The experiment consisted of two sets of three holes each, the first set drilled in 165 mm (6 ½ inch) and the second in 311 mm (12 ¼ inch) diameter. In turn, each set was loaded with 100% LDRA (0.20 g/cm 3 ), 70/30 LDRA/ANFO (0.37 g/cm 3 ) and 50/50 LDRA/ANFO (0.49 g/cm 3 ) respectively. 217

236 COORDINATES TEST # Hole 1 Hole 2 Geo Hole 3 NORTH (m) Hole 4 Geo mm 311 mm Hole Hole EAST(m) Figure 6.37: Mine coordinates showing the relative location of the six blastholes and two geophones used during the first experiment conducted with LDRA The vibration records corresponding to Geo 1, the sensor located behind the set of 165 mm blastholes, are shown in Figure The figure includes the records corresponding to the three directions of measurement (radial, transverse and vertical) as well as the peak vector sum they generate. Blast design data corresponding to this particular test are shown in Table 6.8 while Table 6.9 tabulates the data used in the calculation of the H&P prediction model. Arrival times corresponding to each of the three blastholes are clearly observed and they correlate well with the surface delays used in the experiment. The lower vibration level recorded from the third hole, equidistant to the sensor as the first one but loaded with a 218

237 higher density mixture, is attributed to the signal traveling through already fragmented material. A similar behavior was recorded for the larger diameter holes Radial PV (mm/s) Transverse PV (mm/s) Vertical PV (mm/s) Peak Vector Sum PPV (mm/s) Time (s) Figure 6.38: Vibration records showing the three particle velocity components of the PPV corresponding to LDRA Test 1. Sensor located behind 165 mm diameter blastholes 219

238 Table 6.8: Blast design data for LDRA Test 1 Coordinates Initiation Time Diameter Explosive Hole # Comments Density LDE ANFO East [m] North [m] Delay Arrival [ms] [ms] [mm] [g/cm [g/cc] ] [kg] [kg] Geophone (56Kg+0Kg))' (39Kg+65Kg)' (28Kg+108Kg)' Geophone (167Kg+0Kg)' (117Kg+196Kg)' (84Kg+362Kg)' Geophone (56Kg+0Kg))' (39Kg+65Kg)' (28Kg+108Kg)' Geophone (167Kg+0Kg)' (117Kg+196Kg)' (84Kg+362Kg)' Table 6.9: Data used to generate Holmberg and Persson predictive model for LDRA Test 1 Hole # Hole Length Charge Length Linear Charge Geo Distance Geo Depth Diameter Weight Stem H&P PPV [mm] [m] W [kg] H [m] [kg/m] Ro [m] Xo [m] Xs [m] Factor [mm/s] Figure 6.39 depicts the results of the analysis conducted for LDRA Test 1, where the HPF factor as given in Equation 6.6 for each individual blasthole, is plotted against the PPV in a logarithmic scale. A power regression will result in a line whose slope and ordinate represent the values k and α of the function defining the prediction model, thus, the attenuation characteristics of the rock for that particular area. 220

239 10000 Holmberg & Persson Predictive Model LDRA Test 1 PPV [mm/s] PPV = 439.2(HPF) R 2 = H & P Factor Figure 6.39: H&P predictive model for LDRA Test 1 Figure 6.40 shows the H&P model corresponding to LDRA Test 6, which permits a direct comparison of damage between the LDRA and the standard explosive practice as for Baseline Test B 3 discussed later. This particular test was conducted in the same area and under the same drilling and loading scenario as for Baseline Test B 3, allowing a direct comparison of performance between the LDRA and the standard explosive practice of the operation. 221

240 Holmberg & Persson Predictive Model LDRA Test 6 PPV (mm/s) PPV = (HPF) R 2 = H & P Factor Figure 6.40: H&P predictive model for LDRA Test 6. Buffer row, 350 mm. LDRA at 0.20 g/cm Vibration Analysis of Baseline Experiments Three experiments were conducted to establish a baseline for comparison between the LDRA and the standard explosives commonly used in the buffer row, that is, heavy ANFO (Blendex 945 and 930) and micro bubble sensitized emulsions (Emultex). Appendix 6.4 summarizes available blast and vibration data used for construction of the prediction models of all baseline tests. Figure 6.41 illustrates the results for Baseline Test 3, monitored behind a simulated buffer row using 350 mm diameter blastholes loaded with a heavy ANFO. 222

241 Holmberg & Persson Prediction Model Baseline Test B3 PPV (mm/s) PPV = (HPF) R 2 = H & P Factor Figure 6.41: H & P predictive model for Baseline Test B3. Simulated Buffer row, 350 mm loaded with heavy ANFO Table 6.10 summarizes the maximum peak particle velocity recorded on each of the three baseline experiments along with pertinent blasthole and explosive data. Table 6.10: Relevant data from baseline trials for typical blasting practices used at the mine Diameter Weight Distance Density PPV (max) Test Explosive mm kg m g/cm 3 mm/s B1 Emultex S B2 Blendex B3 Blendex Comparison of Vibration Experiments Baseline Test B3 versus LDRA Test 6 The most significant conclusion obtained from the field experiments conducted at Chuquicamata Mine results from the comparison between Baseline Test B3 and LDRA Test 6, both of which were conducted in the same area of the mine (thus, the same rock 223

242 conditions prevail) and under similar drilling and loading scenarios. Moreover, the vibration sensors were located at 5.5 m and 5.9 m respectively from the nearest blasthole, allowing therefore a direct comparison of results between the standard explosive and the low density product. The maximum peak particle velocity of 2730 mm/s recorded during Baseline Test B3 was produced by 200 kg of the Heavy ANFO mixture (Blendex 930) loaded in 350 mm diameter blastholes. At a density of 1.0 g/cm 3 and for the mentioned blasthole diameter, the 200 kg of emulsion will result in a ~2 m long explosive column. On the other hand, LDRA Test 6, which was loaded with 192 kg of LDRA at 0.20 g/cm 3 resulting in a 10 m long explosive column, produced a peak particle velocity of 930 mm/s; that is three times lower than the explosive used in standard blasting practices. Given the 66% decrease in vibration levels experienced when Blendex 930 is replaced by the LDRA, a considerable reduction of damage is therefore to be expected when the low density product is applied. Moreover, the five fold increase in column length as a result of using the LDRA will better distribute the explosive s energy along the blasthole length, so a more evenly distributed fragmentation and a reduction of collar fragments should also be realized. Baseline Test B2 versus LDRA Test 6 Another comparison worth noting can be established between the LDRA Test 6 above and Baseline Test B2. The latter consisted of 200 kg of heavy ANFO at a density of 1.20 g/cm 3 loaded into 165 mm diameter buffer holes. A maximum peak particle velocity 224

243 of 1225 mm/sec at a distance of 4.5 m was recorded for this particular test. A direct comparison between the preceding two experiments is precluded since the sensor for LDRA Test 6 was placed at 5.9 meters from the nearest blasthole, although, by means of the prediction equation defining the Holmberg and Persson model for LDRA Test 6 (Figure 6.40), it is possible to calculate the distance at which a similar level of vibration as attained by Baseline Test B2 is reached. Trial and error calculations show that at a distance of 4.6 m an H&P factor of 4.85 is obtained, which in turn and according to the predictive model corresponding to LDRA Test 6, produces a PPV of 1235 mm/s, similar to the 1225 mm/s generated by Baseline Test B2. As a result, loading 200 kg of LDRA at 0.20 g/cm 3 in 350 mm diameter blastholes should generate the same vibration level at 4.6 m as 200 kg of heavy ANFO at 1.20 g/cm 3 loaded in 165 mm diameter holes at a similar distance of 4.5 m. From an operational point of view, the previous considerations should have important cost savings and productivity implications by reducing the need for small diameter buffer holes, allowing the use of the more efficient, larger diameter production drills while releasing the smaller equipment for drilling of presplit rows solely. Baseline Test B2 versus LDRA Test 5 A further exercise can compare the maximum PPV generated by Baseline Test 2 as detailed above versus the PPV signal from LDRA Test 5, the latter consisting of a presplit experiment using 165 mm diameter holes. The nature of the presplit blast, which calls for simultaneous initiation and reduced spacing between holes, precluded the 225

244 unequivocal identification of other blastholes from the vibration record as well as the construction of a predictive model. As a consequence, only the initial arriving signal originating from the first presplit hole initiated 10 ms before the remaining ones was considered. The initial signal from Test 5 generated a maximum PPV of about 530 mm/s at a distance of 6.8 m from a 51 kg column of LDRA at 0.20 g/cm 3. Baseline Test B2 on the other hand, conducted in the same diameter but charged with 200 kg of heavy ANFO instead, produced a maximum PPV of 1226 mm/s at 4.5 m from the sensor. Since distance and weight are different for each test, a number of distance weight combinations can be used in the prediction equation of Baseline Test 2 (corresponding graph in Appendix 6.4) to reach the 530 mm/s generated by LDRA Test 5. For comparison purposes however, by assuming the same distance to the sensor for both tests, it is possible to predict the explosive weight that produces a similar level of vibration. Therefore, fixing the distance to the sensor at 4.5 m for both experiments, about 92 kg of heavy ANFO would be required to reproduce the 530 mm/s PPV conditions of LDRA Test 5. This weight translates into 3.3 m of heavy ANFO column in contrast to the 12 m column obtained if the LDRA explosive was used instead. Once more, the benefits of having an extended column to avoid large collar fragments and produce a more evenly distributed fragmentation comes to light. 226

245 Post Blast Inspection of LDRA Field Tests A visual inspection of the blast area was conducted after each field test, in order to assess damage and ensure proper detonation of the LDRA blastholes took place. Of particular interest however, were LDRA Tests 4, 5 and 6, since they provided the opportunity to qualitatively evaluate, on the same blast, eventual differences in performance between the LDRA and the standard products used at the mine. Moreover, using survey data for the toe and crest of the new berm generated by LDRA Test 4, it was possible to quantify its width along the presplit line, and thus make a comparative assessment of damage between the LDRA and the standard presplit product. In addition to the previous evaluations, Chuquicamata Mine developed its own wall control assessment methodology, known as Full Control Program (Salazar et al, 2005). The method takes into account two factors, named Design Factors (Fd) and Condition Factors (Fc). The former refers to design parameters, such as the program line, slope angle, toe toe distance and bench height. The latter refers to the condition of the slope face and accounts for the number of half barrels, the number of induced fractures every 10 m of bench face, the presence of unstable blocks, conditions of minor discontinuities, slope profile and crest conditions. Using the Fd and Fc factors, an index indicative of compliance of a predefined standard of damage is established. According to their established criteria, a certain slope would be deemed acceptable if Fd > 0.6 and Fc > 0.7. Since this index is developed by the mine operators themselves, it is considered an 227

246 unbiased evaluation of the results. The only full control data made available was for the presplit corresponding to LDRA Test 4 and the large diameter blast of LDRA Test 6. Pertinent information and assessment results for the mentioned tests follow. Presplit Test 4 and Test 5: These presplit experiments allowed comparison of the LDRA versus the standard 36 mm diameter Enaline emulsion explosive product used in presplit and illustrated in Figure Figure 6.42: Enaline emulsion cartridge used for pre splitting The presplit for LDRA Test 4, illustrated in Figure 6.12 previously, contained a total of 27 holes spaced 2 m apart. A section of 10 blastholes was loaded with LDRA at 0.20 g/cm 3 and the remaining 17 holes with Enaline. As Figure 6.43 depicts, blast results showed the LDRA section producing a coarser fragmentation and higher lift at the collar than the Enaline product, despite the latter having a shorter collar (2 m compared to 5.5 m of the LDRA holes). This result can be attributed to the higher load per unit length 228

247 provided by the LDRA, which at 4.28 kg/m more than doubles the one corresponding to the 1 ½ inch Enaline cartridges (1.27 kg/m). Enaline LDRA Figure 6.43: Presplit of Test 4 indicating Enaline (back) and LDRA (forefront) sections Post blast inspection of Test 4, to assess and compare overbreak at the LDRA and Enaline areas by quantifying the berm width, are presented in Table 6.11: Table 6.11: Berm widths as an indication of damage produced by the LDRA and Enaline sections corresponding to presplit LDRA Test 4 Enaline Sector LDRA Sector Enaline Sector Hole # Berm Width [m] Hole # Berm Width [m] Hole # Berm Width [m] Average

248 As gathered from the previous table, the average berm width behind the section loaded with LDRA falls between the values corresponding to the two sections loaded with Enaline. Although based only on a single experiment, results indicate nevertheless that LDRA and Enaline generate similar levels of overbreak. The bulk loaded LDRA should therefore represent a valid alternative to Enaline, since it eases the loading process, requires less manpower and reduces charging time while generating the same level of damage. It is worth noting however, that these widths are not necessarily the effect of the presplit alone since the two buffer rows and the production blast detonated at a later date could have well induced damage beyond the limits of the presplit row. The quality index from the Full Control program developed by the mining operation concluded that the wall behind the LDRA section of the presplit blast corresponding to LDRA Test 4 complied with the established benchmark, while the one behind the Enaline did not. Results of the evaluation as provided by mine personnel are presented in Table Table 6.12: Full Control program results for LDRA Test 4 Expansion Blast ID Test 40 East Presplit Bench 2289 Nº 4 Blastholes # Explosive Full Control Benchmark From To Index Evaluation LDRA 165 mm 0.73 Comply Enaline 32 x 305 mm 0.68 No comply The berm slope behind the sector loaded with LDRA looked stable, with fair geometry and competent rock despite the low percentage of half barrels observed. The berm crest showed slight signs of overbreak. On the other hand, the evaluation conducted for the 230

249 sector loaded with Enaline was described in basically the same terms as the LDRA, although, the crest presented enough damage to generate a quality index below the established benchmark. Again, these results can only be taken as indicative of damage; nevertheless, they do position the LDRA as a valid alternative to the Enaline cartridges. The presplit blast corresponding to LDRA Test 5 had a total of 71 holes spaced 1.5 m apart, as presented in Figure 6.13 before. The LDRA section consisted of 24 blastholes, while the remaining 47 holes were loaded with Enaline. Results were similar to Test 4 despite the reduction of spacing from 2 m to 1.5 m. No berm survey data were available for damage assessment nor was a Full Control index for the presplit wall developed by the geotechnical department. However, visual inspection conducted right after the blast revealed similar results as for LDRA Test 4, with those holes loaded with the LDRA developing larger fragments and higher lifts at the collar. LDRA Test 6 The standard procedure for production blasting in this area of the mine called for the last row to be drilled at half the spacing and loaded with half the charge of the previous production row. The objective behind this design approach is to simulate a buffer line using the same 350 mm diameter blasthole as used for production. Blast layout corresponding to LDRA Test 6 is depicted in Figure 6.14 elsewhere in this chapter. As shown in Table 6.13 provided by the mining operation, the last row consisted of three sections, the first one loaded with LDRA at 0.20 g/cm 3, the second one with Blendex

250 (1.0 g/cm 3 ) and a last one with Blendex 945 (1.3 g/cm 3 ), the latter two being heavy ANFO mixtures. Collars were about 10 m for all three explosive products. None of the wall sections behind the three sections attained the expected quality index established by the Full Control program, with all of them having been assigned the exact same index value. The wall condition was described as showing very poor and fragmented slope geometry with no half barrels and a heavily damaged crest. Results are a clear indication that in order to reduce damage and produce higher quality bench faces, design changes need to be implemented, regardless of the explosive selected. Table 6.13: Full Control program results for LDRA Test 6 Expansion Blast ID Test Blastholes # From To Explosive Index Full Control Benchmark Evaluation 47 West Bench 2411 E Nº LDRA 350 mm Buffer Blendex mm Buffer Blendex mm Buffer 0.37 No comply 0.37 No comply Post blast visual inspection clearly showed a transition between the LDRA and the Blendex sectors behind the last row. The LDRA sector, illustrated in Figure 6.44, presented surface scabbing produced by reflected tensile stresses, which are indicative of a short collar design. 232

251 Figure 6.44: Surface damage behind LDRA blastholes The sector behind Blendex 930 showed indications of surface cratering, as seen in Figure 6.45, which illustrates the transition zone between the LDRA and the lighter of the Blendex products. Figure 6.45: Transition area from LDRA to Blendex 930 As observed in Figure 6.46, the sector behind the higher density heavy ANFO (Blendex 945) shows the formation of a deeper trench, a consequence of the higher density explosive load used. Therefore, an increase in the level of damage should be expected behind this particular section of the blast. 233

252 Figure 6.46: Trench behind Blendex 945 blastholes Suggested Design Alternatives The impact that explosive loading characteristics, in particular charge density, diameter and length, have on damage control cannot be overstated. A theoretical approach to visualizing such an impact can be clearly seen by using the ABAQUS finite element code, where the influence that explosive density and diameter have on rock damage can be readily identified. Figures 6.47 to 6.49 (Katsabanis, 2005) represent longitudinal cross sections depicting PPV isolines as a function of radial distance from an explosive column. For clarity, the PPV values at about 10 m and 20 m from the axis of the charge have been captioned. 234

253 Emulsion Explosive φ = 311 mm Figure 6.47: PPV isolines generated by an emulsion at 1.20 g/cm 3 in a 311 mm blasthole LDRA Explosive φ = 311 mm Depth (m) Depth (m) 3065 mm/s at 10 m 1532 mm/s at 20 m Radial distance (m) 818 mm/s at 10 m 409 mm/s at 20 m Radial distance (m) Figure 6.48: PPV isolines generated by LDRA at 0.20 g/cm 3 in a 311 mm blasthole 235

254 LDRA Explosive φ = 165 mm Depth (m) 336 mm/s at 10 m 168 mm/s at 20 m Radial distance (m) Figure 6.49: PPV isolines generated by LDRA at 0.20 g/cm 3 in a 165 mm blasthole It can be observed that as explosive density decreases from 1.20 g/cm 3 (emulsion) to 0.20 g/cm 3 (LDRA), so do the vibration levels at radial distances from the charge. In addition, for a given LDRA explosive, vibration levels are further decreased as charge diameter is reduced from 311 mm to 165 mm. However, a feasible solution to the slope stability problem facing Chuquicamata cannot be addressed by a change in explosive loading alone, since they stand little chance of success unless they are part of a more comprehensive approach for addressing damage. Factors such as rock characteristics, priming positioning, blast boundaries, timing and others need to be considered. Present day contour blast designs used in Chuquicamata in Expansions 40 East and 41 East (where LDRA Tests 4 and 5 and Baseline Tests B1 and B2 took place), consist 236

255 basically of a presplit line detonated in advance, followed by two small diameter buffer rows (165 mm) and two larger diameter production rows (311 mm), the former drilled without subgrade and the latter with a 2 m subgrade. Typical burdens are 6 m, 6 m and 8 m for the two buffers and the production rows respectively. Approximate column lengths are 6.5 m for the first buffer row, 13 m for the second buffer row and about 11 m for both production rows, with the first buffer row coinciding with the toe of the bench. Depending on the sector of the blast, the explosive most commonly used are Heavy ANFO mixtures of different densities (Blendex 930 at 1.10 g/cm 3 and Blendex 945 at 1.30 g/cm 3 ). The first buffer row coincides with the toe of the bench. The first buffer row, designed almost exclusively to remove the toe, consists of about 6.5 m of explosive column, 12 m of stemming and 6 m of burden. This exposes a large block of rock to poor fragmentation, due to the low explosive energy reaching the area, while on the other hand it subjects the toe section, the most sensitive area of the bench, to excessive strain and probable damage. The alternative blast design proposed for these areas consists of removing the two 165 mm buffer rows and replacing them with a single 311 mm diameter production row drilled only 2 m from the toe and loaded with a 9 m column of LDRA. Analysis conducted by ASP Blastronics (2005) reproduced in Figure 6.50 clearly shows a reduction of vibration levels at the toe area (smaller blue and yellow contours) with increasing vibrations at the top part of the bench (larger yellow contour) due to the extension of the explosive column length. 237

256 PPV (mm/s) PPV (mm/s) Reduced damage contour at toe area when using LDRA Figure 6.50: PPV isolines for the present and alternative design for Expansions 40 E and 41 E As seen from both cross sections, damage contours indicate that the alternative design using 311 mm diameter holes as a buffer row drilled 2 m from the bench s toe and loaded with 9 m of LDRA at 0.20 g/cm 3 will eliminate the need for the smaller diameter buffer rows, while significantly reducing damage levels around the toe and improving the energy distribution along the bench height. For the shear zone of Expansion 47 West (where LDRA Tests 2 and 6 and Baseline Test B3 took place), the control technique utilized consisted of a single buffer row using the 238

257 same diameter as the production rows (350 mm) but loaded with an explosive charge about half the one used for production. This buffer row was drilled about 3 m from the toe of the bench. The suggested alternative for this particular area of the mine maintains the present buffer row (diameter, burden, etc.) at 3 m distance from the bench s toe but loads it with an 8 m long column of LDRA at 0.20 g/cm 3 instead of the 2 m long Blendex normally used, with the remaining production rows being unchanged. Figure 6.51 compares damage contours for both the actual and the alternate design. PPV (mm/s) PPV (mm/s) Reduced damage contour at toe area when using LDRA Figure 6.51: PPV isolines for the present (top) and alternate (bottom) design for Expansion 47 West 239

258 As observed from the reduced red area towards the back of the last row (PPV>3200 mm/s), the proposed design will generate less damage than the one being currently implemented. Besides an adequate design of the explosive load, particularly for the two buffer rows and the first production row, responsible for affecting the wall and new face of the bench, an important issue to consider for damage control relates to timing. It is necessary to provide enough delay between successive holes of the same row in order to avoid synergies between two or more consecutive blastholes that will tend to create objectionable vibration levels. This tends to occur when short delays (<10 ms) are used. Figure 6.52 shows the timing corresponding to LDRA Test 6, where four different delay combinations having an initiation time less than 10 ms are pointed out. This short delay between holes will most likely result in the coupling of vibratory energies in the near field, which in turn could result in increased damage. 240

259 Figure 6.52: Timing sequence of LDRA Test 6 showing potential areas of vibratory synergies In order to reduce the probability of delay superposition, it is recommended to use shorter in the hole delays than are currently being used in the operation (600 ms), since they will most likely produce superposition with almost any typical surface delay. Another possibility will be to maintain the 600 ms in the hole delay but increase the surface delays between holes of the same row to at least 35 ms. The preferred avenue, however, would be to introduce electronic detonators in all the control blasts, in order to provide total control of timing and greatly reduce the probability of superposition. 241

260 6.6 Conceptual Design of a Manufacturing Plant for the LDRA In preparation for eventual large scale production of LDRA, the opportunity was sought to evaluate, at a conceptual level, the most important stages of a manufacturing process for the low density product. Critical to the mentioned opportunity was the involvement of the Technology, Engineering and Management (TEAM, 2006)) group, a multidisciplinary project course offered by the Department of Chemical Engineering at Queen s University that links 4 th year students with industries seeking additional consulting resources. The multidisciplinary group was composed of four 4 th year students, staff form the Department of Chemical Engineering of Queen s University, the author and a professional chemical engineer in the capacity of external advisor to the group. The proposal for the Mining TEAM group called for the conceptual design of the low density explosive manufacturing plant to be located in Chile, within the existing ammonium nitrate manufacturing factory of Enaex. Also considered in the design were local conditions, such as dry climate and labor rates, the latter having a large influence on the plant s design. The final design includes a process flow diagram, equipment selection for each process stage, equipment sizing quotes from suppliers, piping and instrumentation diagram for the designed manufacturing plant, logistics; plant operational procedures (batch times, precautions and methods), process hazard analysis and a cost analysis of the manufacturing operation. 242

261 From them all, the process flow diagram, the rationale behind the selection and sizing of the equipment and the cost analysis were considered relevant input to include in this research thesis. A summary of said items are included in Appendix Pending Research After the experience gained in Chile during pilot scale manufacturing and field testing of the LDRA, three operational issues needed to be addressed for the product to have a reasonable probability of becoming part of the standard blasting practices at Chuquicamata mine: 1. preliminary design of a LDRA processing plant (above section and Appendix 6.5) 2. mechanization of blasthole loading (Appendix 4.3) and 3. generation of post detonation flammable gases, which will be addressed next The Problem of Flammable Gases Up scaling from laboratory to field experiments brought up the unexpected issue of flammable gases, which were not observed during the developmental stages of the LDRA product. It was known that the fuel rich composition of the LDRA (OB = 330 g O/kg LDRA) would preclude its use from underground operations unless proper ventilation provisions existed or were provided to expel toxic fumes. Unforeseen however, was the generation of post detonation flames lasting a couple of seconds after the ejection of the stemming material. These flames were caused by an excess of methane and carbonaceous residues present in the products of detonation due to the large fuel content provided by the poly beads. Upon venting and mixing with 243

262 atmospheric oxygen, these detonation products ignited, producing short lived flames that dissipated in a matter of a few seconds. However, post blast inspection revealed the presence of entrapped hot gases within the fragmented muck blowing sporadically into the atmosphere. Although site inspection and particularly VOD records gave a clear indication of a full and stable detonation of the LDRA blastholes, the presence of these entrapped hot gas pockets presented a safety concern that needed to be addressed. In spite of the advanced stage of this research effort, different concepts were investigated and preliminary experiments conducted in an attempt to identify avenues that may solve the issue of combustion/flames. Velocity of detonation tests along with color high speed digital video running at 500 frames/s were used to assess detonation performance and evaluate the magnitude and duration of the flames relative to those produced by the baseline explosive, that is the LDRA at 01.5 g/cm 3. Figure 6.53 shows a sequence of three frames corresponding to the baseline LDRA explosive loaded in a 51 mm steel tube. The strength of the combustion flames is clearly depicted in the initial two frames, while the last frame, corresponding to 28 ms after initiation, is shown as an indication of the duration of the event. Figure 6.53: Frame sequence at 2, 10 and 28 ms after initiation. LDRA at 0.15 g/cm 3 244

263 As a result of these late investigations, various alternatives were deemed feasible to solve the problem of flames, including the addition of chemically inert bulking agents and/or flame retardant powders or by modifying the stemming practices. Addition of bulking agents such as expanded perlite, a volcanic rock having a bulk density of 0.10 g/cm 3, is a feasible approach towards improving the oxygen balance of the mixture and reducing flame occurrence. Perlite being an inert ingredient, the oxygen balance of the resulting mixture will improve in direct proportion to the weight of perlite being added. Perlite granules having a bulk density of 0.10 g/cm 3 and standard LDRA at 0.15 g/cm 3 were mixed in a 50/50 and 40/60 volume ratio for testing. The oxygen balance of these two mixes was calculated at 200 g O/kg and 160 g O/kg respectively, with corresponding densities measured at 0.13 g/cm 3 and 0.11 g/cm 3. VOD of these mixtures was measured at 1580 m/s and 1215 m/s respectively, values that are remarkably different despite having a similar density. Worth noting is that the VOD trace of the 40/60 LDRA/Perlite mixture showed indications of discontinuous propagation. Figure 6.54 shows selected high speed video frames of a 50/50 LDRA/perlite mix at a density of 0.20 g/cm 3. Flames took longer to fully develop, although they were still quite large and lasted as long as the baseline experiment. 245

264 Figure 6.54: Frame sequence at 2, 16 and 26 ms after initiation. LDRA and perlite mixed at 50/50 volume. Mixture density of 0.20 g/cm 3 The resulting flames from a further alternative product are shown in Figure This product consists of a mix of ammonium nitrate and perlite, the latter replacing in its entirety the expanded polystyrene beads used in the LDRA. The product is currently known by the name of PANFO. Figure 6.55: Frame sequence at 2, 4 and 6 ms after initiation. Perlite ammonium nitrate based mixture (PANFO) at a density of 0.28 g/cm 3 For all practical purposes, PANFO is a completely different explosive. As tested, the mixture resulted in a density of 0.28 g/cm 3 and detonated with a VOD of 1710 m/s. The magnitude and duration of the flames, as seen in the high speed frames, was in the order of 4 ms maximum, negligible duration compared to the 28 ms of the baseline LDRA. 246

265 The advantage of this particular product resides in the fact that the source of flammable gases is eliminated and the mixture can be balanced in oxygen, making it eligible for underground applications. In addition to the elimination of flames, the manufacturing process and operational logistics (handling, transporting, storing) of this new low density concept are greatly improved in comparison to the difficulties experienced during pilot scale production and field testing of the LDRA in Chuquicamata. For this reason the product is currently being considered for further development in Chile, where a patent application has recently been presented. The addition of flame retardant powders to the LDRA, such as a proprietary product supplied by Mining Resource Engineering Limited (MREL, 2005) under the brand name H 10, produced remarkable results, suppressing the flames to practically negligible values. This product is supplied as an extremely fine powder (90% < 100 μm), therefore, segregation is greatly reduced in spite of its high bulk density (1.8 g/cm 3 ). However, the issue of cost could play a definite role, since about 25% by weight of this fine powder is required to suppress the flames. Three VOD tests were conducted using 10%, 20% and 30% by weight of the H 10 powder. A decrease in VOD performance was noticeable, with recorded values of 1950 m/s and 1700 m/s for the 10% and 30% additions respectively. Corresponding mix densities were 0.18 g/cm 3 and 0.21 g/cm 3. These experiments clearly show that the addition of denser inert powders will decrease detonation pressure in spite of an 247

266 increase in mixture density. Figures 6.56 and 6.57 show the results of LDRA mixtures containing 20% and 30% of H 10 powder respectively. Figure 6.56: Frame sequence at 2, 4 and 6 ms after initiation. LDRA at 0.15 g/cm % H 10 flame suppressor powder Figure 6.57: Frame sequence at 2, 4 and 6 ms after initiation. LDRA at 0.15 g/cm % H 10 flame suppressor powder There are other ingredients that could have good flame suppressing capabilities and may represent a feasible alternative to H 10 powder. These include rice hulls and bentonite rock, also in the form of fine ground powders. Ground rice hulls are not an inert product, although their fuel content is remarkably low and contains silica, which suppresses flames. Moreover, ground rice hull bulk density and cost are much lower than the H 10 powder, representing an alternative worth investigating. A similar situation occurs with bentonite, a volcanic clay that finds its main application in the petroleum industry as drilling mud. The chemical composition of this clay is particularly 248

267 interesting, since it contains hydrated silica oxides, so both the silica and the water will contribute to suppress the flames. Although the bulk density of bentonite is higher than ground rice hulls, it is not as high as H 10. The third avenue investigated during this preliminary evaluation was to use water as a stemming material. The experiment consisted of placing a 50 cm long plastic sleeve filled with water on top of the 1 meter long column of LDRA at 0.15 g/cm 3 loaded in a 50 mm diameter steel tube. This approach produced the most remarkable results, eradicating the flames to a point in no flame was observable by the high speed video but the one produced by the primer. Further alternatives as stemming material include perlite or similar absorbent rocks soaked in water, as well as limestone rock, salt rock, etc., all of which tend to cool the detonation products and consequently reduce the risk of flames. Regardless of the method finally chosen to deal with the issue of post detonation flames, the above clearly shows that there are several avenues to solve this issue, some requiring minor adjustments to the composition or simple modifications to the loading procedure. The latter, although a straightforward task to implement in the field, requires additional labor time to place the stemming, since water bags will require careful placement to avoid tearing and consequently leakage, which in turn could lead to desensitization and misfires. Moreover, provisions need to be taken for water to be available at the blast site for loading the bags. The use of wetted perlite or similar cooling stemming material not only involves the cost of consumables, but more important, the requirement for additional equipment to transport them to the blast site. 249

268 The small scale experiments conducted at Queen s laboratory gave a clear indication of the different avenues that are worth pursuing to solve the issue of post detonation flames. Extending these observations to predict the behavior of the mixtures at larger operational scenarios should be given careful consideration, since the conditions in which they will eventually be used at the mine will dictate, to a great extent, the performance of the product. 6.8 Chapter Summary This chapter describes the field trials conducted at the Chuquicamata Mine, Chile, and the experimental observations of the LDRA under operational conditions. Important input was gathered during manufacturing at a semi industrial scale of the LDRA, which no doubt will prove useful in future projects. The LDRA offered flexibility during its manufacturing, resulting in a stable product in spite of unexpected changes in the ingredients and the larger production scales involved. Velocities of detonation were measured in several large blastholes, confirming the full and stable detonation of the explosive column. Gas pressure measurements were taken behind the last rows in some of the experiments, in an attempt to validate the approach and discriminate damage associated with gas from that arising from shock vibration. Close range vibration measurements were also taken behind the last row to record peak particle velocities (PPV) and establish a predictive damage model for the blast area. The most important result arising from these latter measurements refers to the PPV generated by the LDRA at 0.20 g/cm 3, where the recorded amplitude proved to 250

269 be almost three times lower than the one generated by the standard explosive used at the mine, all under very similar conditions of distance, load and diameter. The implementation of comprehensive and sound blasting practices to minimize damage should include the ability to modify explosive density approaching the final wall. The use of low density explosives as a damage control tool opens the opportunity of avoiding the two phase approach (production phase followed by a control phase) currently implemented at the Chuquicamata Mine, an operational issue with large economic impact on the productivity of the mine. Modifying the buffer holes, by replacing the less productive small diameter equipment presently used to accommodate the larger diameter production drills, will certainly reflect on productivity and cost. Moreover, the increased slope stability and overall safety of the operation are both good examples of areas where improvements will reflect positively in productivity and costs of a mining operation. 251

270 Chapter 7. Research Contributions 7.1 Summary of Main Thesis Contributions The main effort involved during this research thesis was to develop novel low density explosives, determine their detonation characteristics at a laboratory scale and evaluate their behaviour in large scale applications; all with the final expectation of introducing an improved and feasible alternative to the existing explosive technologies currently used in wall control blasting. For the latter expectation to have a reasonable chance of success, the ingredients selected for the low density composition needed to be readily available in the industry, the explosive mixture amenable for mechanized bulk loading operations, its manufacturing process simple and its detonation characteristics predictable. The investigation concluded with the development of the so called low density reactive agent (LDRA); in other words, a bulking agent which upon surface treatment acquires a reactive nature while maintaining its bulking characteristics. The LDRA represents a flexible system to control density, and as such, a feasible and practical way of controlling explosive energy when used as a diluting agent of ANFO and/or emulsion explosive products. In other words, the LDRA mix is foreseen as a viable approach for matching explosive characteristics to a variety of application requirements, including wall control in open pit mines and construction quarries, as an alternative to detonating cord in dimensional stone operations, as a tool to improve 252

271 collar fragmentation by increasing column length without increasing powder factor, as a secondary blasting explosive and for any condition requiring a flexible system to control the energy delivered to the rock. The characteristics of the LDRA product provided a distinctive approach towards minimizing the effects of segregation that occur when diluting standard ANFO prills with any type of bulking agent. The substantial density difference between ANFO and typical bulking agents will make segregation an unavoidable fact for this kind of mixture. Segregation will lead to uneven propagation, but more importantly, to initiation and/or detonation failures, the latter representing an unacceptable condition. The reactive nature of the LDRA eradicates the risk of initiation and/or propagation failure due to segregation, regardless of the degree of dilution initially sought or the level of segregation finally attained during mixing and loading. This characteristic is considered an important research contribution towards alternate solutions for damage control. The detonation behavior exhibited by the LDRA product, as tested in the laboratory, proved stable and reliable. Product density during manufacturing was controllable within a 10% 15% range, the latter considered reasonable since it has no practical consequences at the low density levels involved. The scatter observed in the VOD values was mostly the result of experimental ambiguities and of testing different batches throughout the developmental stages of the product. 253

272 During the characterization stage of the LDRA, novel experimental approaches were investigated to evaluate performance parameters of the mixture with various degrees of success. A modified aquarium technique used to infer detonation pressure without resorting to photographic methods proved unreliable for the low pressure events generated by the LDRA, although indications were encouraging for products generating higher detonation pressures. For the latter, further experimentation is warranted to corroborate results. Improvements of measuring techniques using thin aluminum probes that allowed a better interpretation of VOD records generated by low order or marginal detonations were evaluated. In addition, methodologies to measure dynamic pressures inside a detonating explosive, which involved the evaluation of different gauge technologies were developed and implemented. A non destructive experimental approach was designed to evaluate the borehole and/or explosion pressure produced by decoupled products, such as detonating cord, or by bulk loaded mixtures, such as the LDRA. The system was able to withstand the detonation pressure of the low density explosive without rupturing, thus allowing the retention of the gases and the recording of the pressure profile of the event until venting. Important lessons were learned during field application of the LDRA explosive in Chuquicamata. Large scale manufacturing proved to be flexible enough to accommodate considerable variations from the mixing pail method used during the course of the investigations at Queen s University. Contrary to the batches prepared at Queen s, 254

273 where the ammonium nitrate fines were taken from a reject stream of the prilling process of a local manufacturing plant, the fines used in Chuquicamata were generated by crushing and sieving ammonium nitrate prills. The previous issue brought up two areas of concern. The first one was related to the manufacturing process, that is, the need for a crushing and sieving phase that would produce ammonium nitrate fines of the desired size range. The second concern was related to the eventual effect on storage life and/or sensitivity as a consequence of the higher hygroscopic behavior exhibited by the fines that are exposed upon grinding, given the lack of surfactant on their fresh surfaces. In spite of these apparent difficulties, the quality control VOD tests conducted throughout the manufacturing process, as well as the actual field experiments, indicated reliable performance of the LDRA explosive. In addition to the above issue, the expanded polystyrene beads supplied at the operation differed from the ones used at the Queen s University laboratory in both size and density. This characteristic does not play a critical role, as long as it is within a reasonable size range, although it weighs heavily on the proportions of ammonium nitrate fines required to obtain a certain LDRA density, and thus, the desired explosive performance. To account for this unexpected difference in ingredient characteristic, the composition of the LDRA was readily modified until the desired performance was reached. Only a small number of VOD quality control tests were necessary to reach a VOD of 2000 m/s sought for the field experiments. The corresponding target density for the velocity was 0.20 g/cm

274 7.2 Patent Information on the LDRA The novelty of the LDRA product is best supported by the patent rights granted to Mr. Guillermo Silva by the United States Patent and Trademark Office (USPTO), US patent number 6,425,965 B1 dated July 30, 2002 and registered under the name Ultra Low Density Explosive Composition. Similar patent rights have also been claimed and granted in the Republic of South Africa, where a Certificate of Grant number 2002/3684 dated November 26 th, 2003 was issued. In addition, a new patent application (# ) dated November 24 th, 2006, has been filed in the name of Mr. Guillermo Silva at the Republic of Chile. The application relates to the alternate low density reactive agent (PANFO) and has been filed under the title Agente de Tronadura Granular de Baja Densidad con Aplicación en Minería. 256

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281 Meng, X; Hustrulid, W; McCarter, M. (2005). Some New Insights on Borehole Wall Pressure When Using Decoupled Charges. Proceedings of the 31 st Conference Explosives and Blasting Techniques (ISEE). Volume 1 Mirzaev, E. S; Matochkin, V. A. (1988). Methods to increase the efficiency of Blasting Operations. Soviet Mining Journal. Volume 2 Number 1. January March edition; pp# 97 Mohanty, B. (1987). Strength of Rock under High Strain Rate Loading Conditions Applicable to Blasting. Proceedings of the 2 nd Int. Symp. on Rock Fragmentation by Blasting. pp# Moxon, N; Armstrong, L. (1990). Low Shock Energy Emulsion Based Wet Hole Explosives. Third International Symposium on Rock Fragmentation by Blasting. Brisbane, Australia. August 26 31; pp# 45 Moxon, N.; Mead, D.; Richardson, S. (1991). Reducing Blasting Costs Using Air Decks: The Do s and Don ts. Third High Tech Seminar. Blasting Technology, Instrumentation and Explosives Applications. Blasting Analysis International (BAI) MREL (2005). Mining Resource Engineering Limited.1555 Sydenham Rd. Kingston, Ontario, Canada Nicholls, H; Hooker, V. (1962). Comparative Studies of Explosives in Salt. Report of Investigation # 6041, US Bureau of Mines. Page 46 Nie, S. (1999). Borehole Pressure in Blast Holes. SveBeFo Report # 42. Swedish Rock Engineering Research Nifad ev, V; Kalinina, N. (1992/a). Gasification of foamed polystyrene during shock loading. Combustion, Explosion and Shock Waves; pp# 630 Nifad ev, V; Kalinina, N. (1992/b). Detonation mechanism in explosive mixtures containing polystyrene foam. Combustion, Explosion and Shock Waves; pp# 650 Obert, L; Duvall, W. (1950). Generation and Propagation of Strain Waves in Rock. Part I, Report of Investigation # US Bureau of Mines Ouchterlony, F; Nie, S; Nyberg, U; Deng, J. (1996). Monitoring of large open cut rounds by VOD, PPV and gas pressure measurements. Rock Fragmentation by Blasting; pp#

282 Ouchterlony, F. (1997). Prediction of Crack Lengths in Rock After Cautious Blasting with Zero Inter Hole Delay. International Journal of Blasting and Fragmentation I pp# Persson, P; Ladegaard Pedersen, A; Kihlstrom, B. (1969). The Influence of Borehole Diameter on Rock Blasting Capacity of an Extended Explosive Charge. Int. Journal of Rock Mechanics and Mining Science, Vol 6, pp# Porter, D; Fairhurst, C. (1970). A Study of Crack Propagation Produced by the Sustained Borehole Pressure in Blasting. Proceedings 12 th Symposium on Rock Mechanics, Rolla, MO. Page 497 Rinehart, J. S. (1965). Dynamic Fracture strength of Rocks. Proceedings AIME, VII Symposium on Rock Mechanics. The Pennsylvania State University Rock, J. (2004). Improving Blasting Outcomes Using SoftLoad Low Density Explosives. Proceedings EXPLO 2004, Perth, Western Australia; pp# 153 Rowe, Goodridge, Stow and Molloy. (2001). Variable Energy Explosives for Soft Ground Blasting. Explo AusIMM Proceedings. Hunter Valley, NSW. Oct. pp# Rozovski, A. S. (1976). Initiation and development of detonation in porous materials impregnated with liquid oxygen. Combustion, Explosions and Shock Waves Rustan, A. (1983). Linear Shaped Charges for Contour Blasting or Stone Cutting. National Swedish Board for Technical Development, Report # Salazar, G.; Quinzacara, N. (2005). Proyecto Full Control. Mina Chuquicamata. Jornadas de Tronadura ASIEX, Mayo Sanchidrian, J. A. (1996). Improvement of Productivity in Quarrying Dimension Stone Using New Drilling and Blasting Techniques. Paper presented at Brite Euram Workshop on Ornamental Stones, Montpellier. Sanchidrian, J. A., and A. Patino. (2002). Numerical Modeling of Detonating Cords in Uncoupled Holes. Proceeding of The Seventh International Symposium On Fragmentation By Blasting, Beijing. pp# Sheahan, R. M; Beattie, T. A. (1998). Effect of Explosive Type on Fines Generation in Blasting. Proceedings of Explo 98, Australia. pp # 41. Shvedov, K. K. (1985). Explosion processes pulsating along the charge in porous explosives. Combustion, Explosions and Shock Waves; pp#

283 Schvedov, K; Koldnunov, V; Gruzdeva, T. (1984). Detonation of low density granules of explosives based on foamed polystyrene. Detonatsiya: Materialy II Vses. Sov. Po Detonatsii. IKhF, AN SSSR, Chernogolovka; pp# (In Russian) Singh, S; Lamond, R. (1996). Applications of Tracer Blasting during Stoping Operations. Rock Fragmentation by Blasting. Fragblast 5. pp#425 Starfield, A. M; Pugliese, J. M. (1968). Compression waves generated in rock by cylindrical explosive charges: a comparison between ca computer model and field measurements. Int. Journal of Rock Mechanics and Mining Science 5 (1); pp# Sudweeks, W. (2000). Dyno Nobel, West Jordan Office, Utah. Confidential Report Tapia Cubillos, M. (1992). Análisis técnico para la obtención de explosivos secos diluídos. V Jornadas de Especialistas en Tronadura. Expomin 92. Depto. de Ingeniería Civil de Minas. Universidad de Chile. pp# 5 TEAM Report (2006): Technology, Engineering and Management: Manufacturing Process of a Low Density Explosive. Department of Chemical Engineering, Queen s University Torrance A; Richardson, S; Moxon, N. (1987). Tailoring Explosive Shock Waves to Improve Fragmentation in Open Cut Mines. Second International Symposium of Rock Fragmentation by Blasting Wackerbarth, D. E., Anderson, M. U., Graham, R. A. (1992). PVDFSTRESS: A PC based computer program to reduce Bauer PVDF stress rate gauge data. Structural Physics and Shock Chemistry Division. Sandia National Laboratories. Report SAND Wang, X. (1994). Emulsion Explosives. Metallurgical Industry Press. Beijing; pp# 118 Wei, Y; Wang, S. (1987). A new Method of Fracture Plane Control and its Application. Second International Symposium of Rock Fragmentation by Blasting Wieland, M. (1987). Cross Borehole Wave Measurements in Underground Coal. 13 th Conference of Explosives and Blasting Techniques (ISEE) Wieland, M. (1993). Instrumenting Delay Blast Malfunctions in Underground Coal. 19 th Conference of Explosives and Blasting Techniques (ISEE) 265

284 Winzer, S; Anderson, D; Ritter, A. (1983). Rock Fragmentation by Explosives. First International Symposium of Rock Fragmentation by Blasting. Lulea, Sweden, pages Winzer, S; Ritter, A. (1979). The Role of Stress Waves and Discontinuities in Rock Fragmentation. Proceedings 21 st US Symposium on Rock Mechanics, Rolla, MO. pp# Zerill, F. J. (ed). (1981). Notes from Lectures in Detonation Physics. NSWC MP Naval Surface Weapons Center, Silver Spring, Maryland. Zhou, X; Yu, Y. Z. (1992). Estimation of Detonation Velocity and Detonation Pressure from CHNO Explosive Mixtures. Journal of the Industrial Explosives Society, Japan. Vol. 53, No. 1, pp# Jan Feb

285 Appendix 2.1: Survey of Explosive Technologies Applied to Wall Control Considerable effort has been placed at developing explosive technologies to mitigate the undesirable effect of over break and dilution. In an attempt to emphasize this effort, a review of selective low density explosive products and techniques used in wall control applications throughout the last years was conducted. In many of the cases surveyed, the explosive products were the result of diluting ANFO by addition of lighter ingredients. In other cases, denser inert ingredients were added and or the chemical composition altered, in order to reduce the VOD and the explosion pressure. A well known low density explosive developed in Sweden under the name of Isanol consisted of a straight mixture of ANFO prills and expanded polystyrene beads; the latter ranging between 3 mm to 5 mm in diameter (Heltzen, 1980). Claims of 95/5 polystyrene/anfo volume ratios (density < 0.10 g/cm 3 ) are documented to detonate at 1600 m/sec under confined conditions in diameters as small as 35 mm and initiated by a blasting cap. Oxygen balance cannot be obtained with this composition at densities below 0.30 g/cm 3, however, due to the extremely light weight of the polystyrene beads, densities as low as 0.10 g/cm 3 will produce fairly acceptable oxygen balance values. It is possible to modify the composition of ANFO by reducing the amount of fuel oil and improve the oxygen balance of the mixture, however, the quality of mix between fuel and oxidizer will not be the same and as a consequence, the risk of generating nitrogen oxide fumes (several times more toxic than carbon monoxide) becomes an issue of 267

286 consideration. Segregation and static build up upon loading are also issues that need to be addressed. Excessive segregation could lead to propagation failures and static buildup is certainly a safety consideration to be aware of when dealing with explosives. Bulk low strength water gel products developed in 1992 by Ireco (now Dyno Nobel) consisted of an aqueous solution of oxidizing salts as the continuous phase, fuel, thickeners and a density reducing agent (bulking agent) to which ammonium nitrate prills or ANFO prills were added to control energy output. Sensitivity was imparted via the bulking agents in the form of gassing, micro balloons, foamed polystyrene beads and combinations thereof. Density could be adjusted between 0.40 g/cm 3 and 0.70 g/cm 3. Experimental VOD for a mixture containing wt. 25% ANFO ranged between 2400 m/s and 3000 m/s for semi confined tests in 100 mm to 200 mm diameter PVC pipes. The minimum primer was in the order of 90 g of high explosive (Pentolite) for a 100 mm diameter pipe while the critical diameter was in the order of 100 mm for the semiconfined shots (Jackson, 1993) Bulk low shock energy emulsion mixtures were also investigated with the objective of providing water resistance to the explosive product for use under wet conditions (Moxon et al, 1991). They consisted of a mixture of three basic components: an emulsion matrix containing an aqueous solution of ammonium nitrate, a sensitizer (glass microballoons) and a bulking agent (polystyrene, perlite, vermiculite or bagasse). Experimental results indicated an increase in water resistance of the explosive mixture was obtained when bulking agents like perlite and vermiculite were added to the 268

287 emulsion matrix. On the other hand, bulking agents like bagasse and sawdust failed to pass the test due to their lipophilic (affinity to oil) behavior, which resulted in high oil absorption leading to a breakdown of the emulsion matrix. The emulsion mixtures developed had critical densities of 0.78 g/cm 3 and 0.62 g/cm 3 when mixed with perlite and polystyrene respectively. VODs for micro balloon sensitized emulsions containing polystyrene and perlite as bulking agents were measured at 4700 m/s to 3400 m/s for corresponding densities of 1.05 g/cm 3 and 0.80 g/cm 3 when tested in 100 mm diameters. Critical diameters were dependent on the degree of sensitization imparted to the emulsion matrix. For example, the critical diameter for an emulsion polystyrene mixture decreased from 75 mm to 25 mm when micro balloon addition was increased from 1.5% to 3.5% by weight. The emulsion perlite mixture showed a smaller critical diameter than the polystyrene mixture when sensitized with wt. 1.5% micro balloons but at the higher range (3.5% weight.) the critical diameter proved the same as with polystyrene (75 mm). Sensitization with 5% weight of micro balloons did not produce noticeable increases in VOD for either mixture. Other researchers have exhaustively tested low density ANFO mixtures obtained by addition of bagasse, sawdust, rubber, polystyrene, peanut skin, coal refuse, wheat, perlite, vermiculite and similar (Moxon et al, 1990). The bulking agents exhibit different levels of mix quality, with perlite and vermiculite showing a comparatively poor mixing capability. According to Moxon et al, from all the bulking agents tested, the best results were obtained with ANFO/bagasse at volume ratios ranging between 30/70 and 70/

288 Within the previous group of explosives we can include mixtures of ANFO prills with ingredients that will increase rather than decrease the density of the resulting mix; however, they will reduce the VOD and consequently the explosion pressure. Examples of these ingredients are sodium chloride (Crossland, 1982) and Stonegrit (Hardwick, 2002), the latter a mineral product commonly used for sandblasting. The actual manufacturing process of the ammonium nitrate prills themselves has been yet another approach taken to reduce explosive bulk density. Such examples include ANFEX, ammonium nitrate prills from African Explosives Limited (AEL) whose nominal bulk density of around 0.71 g/cm 3 is achieved by chemically inducing air voids during the manufacturing process; and EXPAN, ammonium nitrate prills manufactured by Sasol SMX also from South Africa, where densities as low as 0.61 g/cm 3 are obtained by the controlled addition of plastic microspheres. Detonating cord and ANFO tracing are two techniques regularly used for wall control. The former achieves its objective not through its low density but through decoupling effects and has been widely used in dimensional stone quarrying. It consists of a PETN explosive core enveloped by several layers of protecting materials. Cords come with various loads, ranging from 1.60 g/m to 85 g/m. For dimensional stone quarrying, core loads of 3.8 to 5.3 g/m have been successfully used to split the rock. Water is usually added to the holes to enhance shock transfer to the rock (Barkley et al, 2001). 270

289 Tracing consists of placing ANFO and a detonating cord in the hole and initiating them with a non electric detonator positioned at the bottom. The mechanism of tracer blasting will reduce damage due to a combination of effects including partial burning of ANFO, side initiation of ANFO, desensitization of ANFO and a partition of explosive energy towards gas rather than shock. Booster cords, which differ from normal detonating cords in that they have a low strength core (3 g/m) with bumps of a heavier core load distributed every 3 m, represent an interesting alternative since they simplify the loading operation (Singh, 1996). Low density water gels such as Dynolite, a Dyno Nobel product consisting of an aqueous solution of inorganic salts (continuous phase) sensitized with glass micro spheres or gas bubbles (chemical gassing or entrapped air bubbles), reported densities ranging between 0.40 g/cm 3 and 0.70 g/cm 3. The minimum primer recommended for Dynolite is 150 g (400 g for holes > 150 mm) while critical density is around 1.30 g/cm 3. The VOD attained for a mixture density of 0.50 g/cm 3 was about 1000 m/s lower than ANFO under the same diameter and confinement conditions. Dynolite can be loaded using standard truck delivery systems or pumpable and heavy ANFO delivery systems. It is very susceptible to water and has a 3 week sleeping time when dry conditions prevail. It has found particular applications in controlled blasting programs, ore dilution/grade control, strip coal mines and places where strict vibration levels are imposed (Hunsaker, 1995). 271

290 ANRUB consists of a mixture of standard ANFO prills and rubber particles of various sizes (Harries, 1993). Although rubber is not a particularly good diluting agent, its mixture with ANFO will not produce low density mixtures; however, due to the poor quality of the resulting fuel oxidizer mix, a reduction of VOD (thus shock energy) is obtained, while the total energy remains at the same level. In other words, a slow reacting explosive producing a larger level of heave energy will be obtained. Depending on the particle size of the rubber, ANRUB mixtures showed VODs between 2600 m/s and 3900 m/s. The critical diameter under unconfined conditions ranged between 100 mm and 150 mm. Moreover, initial heave velocities of ANRUB compared favorably with those of ANFO (about 35% higher). In recent years, an emulsion based product developed by Orica (Rowe et al, 2001) was introduced to the market under the name of Novalite, with densities ranging between 1.10 g/cm 3 to 0.30 g/cm 3 and adjusted by the addition of expanded polystyrene beads. The lower density mixtures were successfully used for removing soft overburden (0.60 g/cm 3 ) and as a presplit product (0.30 g/cm 3 ) in coal mines. Softload LDE consists basically of a bulk loaded heavy ANFO whose density can be controlled in the range of 1.25 g/cm 3 to 0.45 g/cm 3 by addition of rice hulls as bulking agent. It is being used for cast blasting in open pit coal operations (Rock, 2004) and due to the excess fuel available in its composition; the serious problem of nitrogen oxide clouds has been eliminated. In some coal operations, Softload (0.45 g/cm 3 ) is being used in the front row as a substitute of a high density product (1.20 g/cm 3 ), in the last row for 272

291 wall control purposes and in the collars of holes to eliminate stab holes and achieve an effective breakup of the cap rock. Blastlite, a recent development of Dyno Nobel Asia Pacific in Australia (Beach et al, 2004) is a mixture of ANFO prills with an undisclosed bulking agent. The manufacturer claims that the characteristics of the bulking agent are such that a homogeneous mixture with little segregation can be obtained, without using emulsion as a binding agent. However, the spread in VOD data does not correlate well with such a claim. The VOD reported for Blastlite at 0.56 g/cm 3 under confinement ranged between 2700 m/s and 3200 m/s for 200 mm and between 2800 m/s and 3400 m/s for 311 mm diameter blastholes. Although no information is available on the bulking agent used in Blastlite, close up photographs reveal the presence of rice or similar grain hulls. If this proves right, the bulking agent would then be compressible, a definite advantage when it comes to transportation costs. Considerable effort was placed in the design of a delivery system that minimizes and controls dust generation, an indication that the bulking agent tends to produce it. A recent product by Orica called Flexigel, is an emulsion mixed at various ratios with expanded polystyrene beads to achieve mixture densities between 0.30 and 1.10 g/cm 3 depending on the amount of beads added. Flexigel has been designed primarily for dry and de watered blastholes; however, it is suitable for wet holes only at higher densities, becoming water resistant at or above 0.80 g/cm 3. The VOD data available at the lower density (0.30 g/cm 3 ) ranged between 2200 and 2800 m/s in a 250 mm diameter hole. 273

292 Appendix 4.1: Information on Bulking Agents Nanogel, consists of translucent, irregularly shaped particles of an aerogel product prepared by treating amorphous silica gel with an organo silicon compound. The particles range between 0.5 mm to 4 mm in diameter with a bulk density ranging between 0.08 g/cm 3 and 0.25 g/cm 3 depending on particle distribution. Nanogel particles are hydrophobic in nature; thus, they will readily absorb oil; have a highly porous structure consisting of 95% air and 5% solids (volume basis) and an extremely small pore size of ~20 nanometers in diameter. Nanogel tends to collapse upon the application of low shear forces; however, it does exhibit moderate compressive strength. From a bulking point of view, these particles present no particular advantage to expanded polystyrene beads except the irregular shape which will help to reduce segregation. Moreover, Nanogel s application is not directed to the massive volumes that would be required for blasting and as such, prices in the order of US $40 to $55 per kg would practically exclude it as a commercially viable bulking agent ingredient for the LDRA. Expancel, consists of small spherical plastic particles having of a polymer shell encapsulating a gas. Upon the application of heat, the gas expands and the shell softens, resulting in an increase of volume of up to 40 times its original value, reaching bulk densities in the order of 0.03 g/cm 3. However, the maximum particle size of the expanded product is around 90 μm in diameter, size considered too small 274

293 for the purposes of this research, where the bulking agent is sought to provide a support structure to the energetic ingredient. The interesting aspect about this product is the simplicity of the expansion process, where unlike polystyrene, no vapor heat treatment and drying techniques are necessary. This opens the opportunity for high density transport and on site expansion, an important cost reduction consideration. Cellulose granules (K Sorb), form a particulate product manufactured from recycled cellulose, capable of retaining more than five times their weight in liquid. It is a lightweight and highly absorbent product typically used in oil spill cleanups, sludge stabilization and liquid waste solidification prior to transportation and disposal. It represents a valid alternative to clay based absorbents and peat moss, having the capability of absorbing 500 % more than clay products such as cat litter. If required, these cellulose granules can be supplied with an oleophilic treatment which enhances its absorption characteristics to about 13 times its own weight in hydrocarbon liquids while still repelling water. They are an economical alternative to polypropylene, sphagnum peat moss, diatomaceous earth, clay, and treated wood fiber. It is also manufactured as packing media to cushion containers and absorb leaking liquids during transportation. UltraFoam: is a state of the art reticulated vitreous carbon rigid foam obtained by pyrolysis of a polymeric material. It is extremely light and strong and can be manufactured having distinct mechanical properties. These foams are thermally 275

294 stable, chemically pure and resistant to shock and stress. They are used in aeronautics for insulation purposes. The high porosity that characterizes this material makes it suitable to retain ammonium nitrate particles within its structure and form the fuel oxidizer combination necessary for detonation. However, the product was immediately discarded due to its high cost (unit price is quoted by the cubic inch) and physical characteristics that preclude it from being bulk loaded into holes. Perlite: refers to a generic term for a naturally occurring siliceous rock. The distinguishing feature that sets perlite apart from other volcanic glasses is that when heated to a suitable point in its softening range, it expands from four to twenty times its original volume. This expansion is due to the presence of 2% to 6% combined water in the crude perlite rock. When quickly heated to above 870 C, the crude rock bursts open as the combined water vaporizes and creates countless tiny bubbles that account for the lightweight and other physical properties of the expanded perlite product. Expanded perlite is a very absorbent product that can be manufactured at densities ranging between 0.03 g/cm 3 to 0.40 g/cm 3, making it adaptable for numerous applications. Since perlite is a form of natural glass, it is classified as chemically inert, an important contrasting feature when compared to expanded polystyrene, given that it will have no influence on the oxygen balance of the final mixture. 276

295 Industrial applications of perlite are the most diverse, ranging from high performance fillers for plastics to cement for petroleum, water and geothermal wells. Other applications include its use as a filter medium for pharmaceuticals, food products, chemicals and water for municipal systems and swimming pools. Additional applications include its use as an abrasive in soaps, cleaners, and polishes; and a variety of foundry applications utilizing perlite insulating properties and high heat resistance. This heat resistant property is taken advantage of when perlite is used in the manufacture of refractory bricks, mortars, and pipe insulation. Perlite is also used as an absorbent to control and clean up liquid spills, as well as a carrier for pesticides, feed concentrates, herbicides and other similar applications. As mentioned elsewhere in this thesis, perlite has been used in dry explosive mixtures for diluting ANFO. Mixture densities of 0.45 g/cm 3 have detonated successfully (Tapia, 1992). From all these properties and potential uses, the ones foreseen as being the most important for the purposes of this research are its low density, its absorption ability and its chemically inert characteristics. 277

296 Appendix 4.2: Recommendations on LDRA Final Mix 1. Recommendations Regarding LDRA Ingredients The experience gained during development and mixing of the LDRA proved valuable in suggesting the following recommendations regarding the explosive ingredients: Expanded polystyrene beads (EPS) having a bulk density around 17 kg/m 3 proved most suitable to achieve lower LDRA densities while maintaining stable detonation conditions. In addition, when preparing mixtures of LDRA and standard ANFO prills, a 2 mm diameter EPS bead is recommended in order to reduce segregation between the two components of the mix Use of mineral oil or any other paraffin or vegetable based oil is a key factor to providing an adherence layer to the EPS beads. Mineral oils have a strong surface affinity with polystyrene and will readily disperse onto the EPS surface, eliminating the static clinging always present during handling of the beads Care should be exercised to prevent exposure of the polystyrene beads to naphthenebased oils such as gas oil, fuel oil and the likes. These products are incompatible with polystyrene, as their vapours will chemically attack the polystyrene structure leading to the collapse and dissolution of the expanded beads within a short period of time The addition of tackifying agents to lighter, less viscous oil has preference over the use of high viscosity oils since the latter are more difficult to disperse onto the beads 278

297 and mix with the ammonium nitrate fines. The tackifyer agent increases adhesive strength of the resulting oily solution without increasing oil viscosity The ammonium nitrate fines should be balanced in oxygen with the mineral oil solution so the resulting ANMO fines would have improved hygroscopic characteristics and higher detonation sensitivity. Being ANMO the main energetic ingredient of the LDRA mixture, increasing its sensitivity will ensure not only a more stable propagation but also a longer storage life than if ammonium nitrate fines were used instead. Preferably, the same mineral oil solution applied on the polystyrene beads is recommended for mixing with the ammonium nitrate fines Using an oil soluble dye to colour the mineral oil solution for quality control purposes of the wetted EPS, ANFO fines and ultimately the LDRA mixture is considered good practise The preferred source of ammonium nitrate fines are rejects from the manufacturing stream of prilling plants. These rejects cannot be recycled back into the process given the relatively high content of surfactant (already added to the process) that remains entrapped within the fines. The presence of said surfactants provides improved hygroscopic characteristics to the fines, making them free flowing and extremely easy to work with Other sources of ammonium nitrate fines include those obtained from grinding explosive or fertilizer grade prills as well as those from grinding pure ammonium nitrate crystals; however, extra consideration should be given to the effect of 279

298 humidity absorption on caking and its influence on mixture sensitivity and performance Particle size of the ammonium nitrate fines should be such that they are easy to mix with the mineral oil solution and small enough to adhere to the wetted EPS surface to reduce segregation. Fines ranging from 600 μm down to +100 μm provide both ease of mixing and good surface retention, as long as they are evenly distributed throughout this range Recommended tackifying agent dose ranges between 0.5% and 5% depending on brand; however, a 3% by weight proved appropriate regardless of the supplier Wetting the 2 mm EPS beads with 15% by weight of the mineral oil solution proved sufficient for the 2 mm bead size 2. Recommendations Regarding LDRA Manufacturing The process of learning from preparing low density mixtures throughout this research provided a few practical tips as to the manufacturing stages of the LDRA, which will be presented next. These tips were to prove useful later, when manufacturing of much larger volumes was required for a semi industrial scale field evaluation as will be discussed in upcoming chapters. Basic Preparation Steps The following five simple steps are recommended when manufacturing an LDRA mixture based on the 2 mm diameter, 17 kg/m 3 density EPS bead: 280

299 A. Prepare a mineral oil solution containing ~3% by weight of the tackifyer agent, adding a coloring dye in order to provide a strong contrast to the oil solution B. Classify ammonium nitrate fines to the recommended range of 600 μm to +100 μm C. Wet the 2 mm EPS beads with 15% by weight mineral oil solution D. Mix the ammonium nitrate fines with the same mineral oil solution in a 94:6 weight ratio (oxygen balance proportions) to obtain ANMO fines E. Mix the wetted EPS beads with the ANMO fines at the weight ratio required to obtain the desired LDRA density Mixing Considerations Two basic components result from the preparation steps described above: wetted EPS beads (step 3) and ANMO fines (step 4). All that is required to produce an LDRA mixture (step 5) at a specified density is to mix these two components at the appropriate proportion. The following issues are considered to be of practical significance when mixing these two components: A. The density of the final LDRA mixture will depend on two factors: the bulk density of the polystyrene beads and the proportion of ANMO fines added B. To obtain an LDRA of a given density, a stepwise addition of ANMO fines to the wetted EPS is recommended, with density readings taken in between steps to assess conditions. Once the required weight percentage of fines has been determined, successive batches can be prepared by adding all the fines in a single step, since any gentle mixing action will disperse them evenly throughout the wetted beads 281

300 C. The mixing action can be supplied by a cement mixer or similar tumbling equipment. Hand mixers, such as egg beaters or similar slow moving tools do not damage the beads and can be safely used when dealing with small quantities D. For the 2 mm diameter, 17 kg/m 3 density EPS beads, the weight percentages of ANMO fines (relative to the EPS) shown in Table A should be considered as a first approximation to obtain the listed densities: Table A: Weight % of ANMO fines required to achieve a given LDRA density Addition of ANMO Fines Expected LDRA Density (as % weight of EPS) (g/cm 3 ) E. Detonation experiments with the LDRA at densities lower than 0.10 g/cm 3 showed the characteristic signs of unstable propagations, thus densities lower than 0.12 g/cm 3 are not recommended for reliable initiation and stable propagation Density Considerations The following experimental observations regarding the density of an LDRA mixture prepared with the suggested ingredients and procedures are worth consideration: A. The maximum LDRA density obtained just by adhesion of ANMO fines onto the surface of the expanded polystyrene beads is in the order of 0.08 g/cm 3. Mixture densities above this value will contain ANMO fines distributed within the interstitial spaces of the beads. Since unstable regimes were observed at this density, higher values should be used. Therefore, ANMO fines would be entrapped within these 282

301 interstitial spaces in a quantity that will depend on the final density sought for the LDRA B. A target density of 0.15 g/cm 3 is recommended as the density of choice for an LDRA mixture to have a reliable initiation and stable propagation and where a reasonably small amount of fines will remain within the interstices of the beads without segregating when subjected to handling or shaking C. A density of 0.20 g/cm 3 is seen as the upper limit above which an accumulation of ANMO fines within the interstitial spaces of the beads becomes excessive. Although some of these fines will tend to settle upon transportation and handling, they will however remix upon loading of the blastholes D. When explosive densities above 0.20 g/cm 3 are required, it is recommended to dilute standard ANFO (0.82 g/cm 3 ) with LDRA at 0.15 g/cm 3 rather than add more ANMO fines to the LDRA 283

302 Appendix 4.3: Potential Field Improvements 1. Technologies for on Site Expansion of Polystyrene As mentioned in Chapter 4, the price of low density explosives, formulated around any bulking agent, will be strongly influenced by the high transportation costs of these light consumables. From the various bulking agents available, polystyrene is one that can be supplied unexpanded, in the form of small beads (< 0.5 mm diameter) having densities in the order of 0.8 to 1.1 g/cm 3. Expansion will increase diameter by as much as times while reducing its density by more than an order of magnitude. If a low density explosive based on expanded polystyrene beads becomes part of the standard blasting practice of an operation, the volume requirements may well justify the acquisition or development of technology that allows for on site expansion of the beads. Having onsite capabilities to address this issue will greatly reduce transportation costs and make the low density mixture more competitive with other existing technologies. Standard expansion technology for polystyrene beads is based on two stages, the expansion stage itself, using vapour as the source of heat, followed by a settling period where the vacuum created due to the expansion process dissipates and stabilizes to atmospheric conditions. Compared to other expanded bulking agents such as perlite, where a much higher expansion temperature is required, the technology available for 284

303 polystyrene represents an alternative worth evaluating for on site expansion due to its simplicity and cost saving potential. On a different angle, expandable polystyrene beads are manufactured using two types of blowing agents: pentane and carbon dioxide. Pentane gas has no effect on the upper ozone layer; although, if not recovered, it can contribute to low level smog formation. Therefore, manufacturers use state of the art technology to capture pentane emissions. With ever evolving technologies, some producers are using carbon dioxide (CO2) as the expansion agent of choice. CO2 is non toxic, non flammable, does not contribute to smog and has no ozone depletion potential. In addition, the CO2 used for this expansion process is recovered from existing commercial and natural sources, making the technology more attractive from an environmental standpoint since it does not increase the levels of CO2 in the atmosphere. In a similar fashion, recent technological developments resulted in water expandable polystyrene beads, in other words, water instead of pentane is used as the blowing agent. The process creates a molecular bond encapsulating starch in a shell of polystyrene. The chemically bonded starch absorbs micro drops of water, which becomes a safer, more environmentally friendly blowing agent inside the beads and does not generate the environmental issues presented when pentane gas is used as blowing agent. 285

304 2. Mechanized Loading Considerations Two approaches are foreseen for mechanized delivery of the LDRA into the blastholes: pre mixing at a mixing plant, transporting to the blast site and loading into the blastholes, or alternatively, using a mixing/delivery unit at the blast site. The former has the advantage that it requires simpler mechanized loading equipment, its main disadvantage being the logistics involved in handling, storing and transporting a material classified as an explosive. On the other hand, mixing at the collar will greatly simplify the logistics since no handling of explosive material occurs until loading, however, special mixing/loading trucks need to be designed to handle the different ingredients. Mechanized loading of pre mixed LDRA should be easily accomplished using readily available equipment such as the ANFO delivery truck illustrated in Figure A 1, consisting basically of a truck mounted storage bin with a belly auger discharging into a vertical auger which in turn feeds a horizontal rotating discharge auger. The prill fuelling capabilities of these trucks would not be of any use for the LDRA. In addition to these delivery systems, standard cement mixing trucks having a rotating drum mixer can also be adapted for mechanized loading of LDRA. 286

305 FigureA I: Bulk delivery truck for standard ANFO A question mark is raised when it comes to deciding which alternative will generate less segregation when mixes of LDRA and ANFO are desired, in other words, the convenience or not of premixing the two ingredients to the desired volume ratio at the plant, or doing it at the blast site prior to loading. The answer to this question requires a more detailed study; however, from laboratory experiments conducted at Queen s, as well as field trials by pre mixing LDRA and ANFO, it was observed that upon mixing, ANFO prills tend to settle as a bottom layer regardless of mixing time or speed. Moreover, pre mixing these two ingredients will be subjected to segregation during handling and transportation. During auger or drum discharge, further separation into an upper and a lower layer will occur. In addition, discharging the pre mix at the collar will result in further segregation due to the different fall velocities of the two ingredients. 287

306 From the previous observations there seems to be no special advantage to pre mix these two ingredients just to expose them later to a list of sources that will induce their segregation. Since segregation due to free falling is inevitable and to a great extent dictates the level of segregation that will finally occur in the blasthole, pre mixing appears to be a needless exercise as long as the mix is being discharged at the collar. Indications are that a simple mechanized delivery system should address this problem in a satisfactory manner. Slightly modified ANFO trucks, with an independent storage bin and augers for each ingredient, should be adequate to provide the desired volume ratios by controlling feeding rates on each discharge stream, with the two streams merging at the collar and mixing during free fall. Moreover, the explosive alternative mentioned in Chapter 6 (PANFO), where all the expanded polystyrene beads are substituted by perlite, could represent an interesting alternative worth investigating in the future. To the mentioned capability of being oxygen balanced, this mixture has the additional advantage that the fuel can be incorporated to the mixture prior to discharge, in a similar fashion as for standard ANFO prills; thus, existing mixing truck technology should prove feasible without the need to implement major changes. Moreover, given that the mixture can be treated as an oxidizer material rather than a high explosive as the LDRA, handling and transportation problems are greatly eased. 288

307 Appendix 5.1: Velocity of Detonation Records: VOD Diameter Relationship 0.10 g/cc- 75 mm steel 0.12 g/cc- 50 mm steel Distance (m) VOD = 1871 m/s Distance (m) VOD = 1862 m/s Time (ms) Time (ms) Distance (m) g/cc- 75 mm steel VOD = 1858 m/s Time (ms) Distance (m) g/cc- 50 mm steel VOD = 1712 m/s Time (ms) 289

308 g/cc- 75 mm steel 0.16 g/cc- 50 mm steel 1 1 Distance (m) VOD = m/s Distance (m) VOD = 1930 m/s Time (ms) Time (ms) g/cc- 75 mm steel g/cc- 75 mm steel (VOD sensor~1.8 m) Distance (m) VOD = 1893 m/s Distance (m) VOD = 1971 m/s Time (ms) Time (ms) 290

309 g/cc- 75 mm steel (VOD sensor~1.2 m) g/cc- 50 mm steel (VOD sensor~1.8 m) Distance (m) VOD = 1969 m/s Distance (m) VOD = 1925 m/s Time (ms) Time (ms) g/cc- 50 mm steel (VOD sensor~1.6 m) g/cc- 75 mm steel Distance (m) VOD = 2216 m/s Distance (m) VOD= Time (ms) Time (ms) 291

310 g/cc- 36 mm steel 0.15 g/cc- 32 mm steel Distance (m) VOD = 1758 m/s Distance (m) VOD = 1786 m/s Time (ms) Time (ms) g/cc- 25 mm steel g/cc- 25 mm steel (blasting cap) Distance (m) VOD = 1680 m/s Distance (m) VOD = 1632 m/s Time (ms) Time (ms) 292

311 g/cc- 75 mm steel (cap + A-3) 0.15 g/cc- 50 mm steel (cap alone) Distance (m) VOD = 1742 m/s Distance (m) VOD = 1760 m/s Time (ms) Time (ms) g/cc- 25 mm steel (cap alone) 1.4 ANFO Initiation Test Distance (m) VOD = 1570 m/s Distance (m) LDRA = 1905 m/s ANFO = 3192 m/s Time (ms) Time (msec) 293

312 g/cc- 50 mm Aluminum 0.15 g/cc- 50 mm steel Distance (m) VOD = 1703 m/s Distance (m) VOD = 1744 m/s Time (ms) Time (ms) g/cc- 150 mm steel g/cc- 150 mm steel Distance (m) VOD = 1965 m/s Distance (m) VOD = 1903 m/s Time (ms) Time (ms) 294

313 LDRA 0.15 g/cc- 25 mm steel (cap alone) 0.15 g/cc- 32 mm steel (5x75 gr/ft detcord ends) Distance (m) VOD = 1515 m/s Distance (m) VOD = 1826 m/s Time (ms) Time (ms) g/cc- 38 mm steel (2x75 gr/ft detcord ends) g/cc- 38 mm steel (1x75 gr/ft detcord end) Distance (m) VOD = 1831 m/s Distance (m) VOD = 1676 m/s Time (ms) Time (ms) 295

314 g/cc- 50 mm steel (aquarium) 0.15 g/cc- 38 mm steel Distance (m) VOD= 1872 m/s Distance (m) VOD = 1891m/s Time (ms) Time (ms) g/cc- 50 mm aluminum g/cc- 50 mm PVC plastic Distance (m) VOD = 1862 m/s Distance (m) VOD = 1712 m/s Time (ms) Time (ms) 296

315 g/cc- 50 mm Cardboard 0.15 g/cc- 50 mm Cardboard Distance (m) VOD = 1200 m/s Distance (m) VOD = 1229 m/s Time (ms) Time (ms) 0.05 g/cc- 50 mm steel 0.11 g/cc- 50 mm steel Distance (m) VOD = 1550 m/s Distance (m) VOD = 1740 m/s Time (ms) Time (ms) 297

316 g/cc- 75 mm Papertube (160 g Primer) g/cc- 75 mm Papertube (80 g Primer) Distance (m) VOD = 1253 m/s Distance (m) VOD = 1352 m/s Time (ms) Time (ms) g/cc- 75 mm Papertube (40 g Primer) g/cc- 75 mm PVC Plastic Distance (m) VOD = 1395 m/s Distance (m) VOD = 1638 m/s Time (ms) Time (ms) 298

317 g/cc- 75 mm steel (aquarium) 0.15 g/cc- 75 mm steel (aquarium) Distance (m) VOD = 1866 m/s Distance (m) VOD = 1890 m/s Time (ms) Time (ms) 0.15 g/cc- 75 mm steel (aquarium) 0.15 g/cc- 50 mm steel (aquarium) Distance (m) VOD = 1860 m/s Distance (m) VOD = 1832 m/s Time (ms) Time (ms) 299

318 g/cc- 50 mm steel (aquarium) 0.10 g/cc- 38 mm steel Distance (m) VOD = 1806 m/s Distance (m) VOD = 1631 m/s Time (ms) Time (ms) 0.15 g/cc- 45 mm Plexiglas 0.15 g/cc- 32 mm steel Distance (m) VOD = 1117 m/s Distance (m) VOD = 1700 m/s Time (ms) Time (ms) 300

319 g/cc- 50 mm steel 0.15 g/cc- 32 mm steel Distance (m) VOD = 1806 m/s Distance (m) VOD = 1595 m/s Time (ms) Time (ms) 1.6 LDRA 0.16 g/cc- 75 mm steel g/cc- 50 mm steel Distance (m) VOD = 1981 m/s Distance (m) VOD = 1925 m/s Time (ms) Time (ms) 301

320 g/cc- 75 mm steel 0.15 g/cc- 125 mm steel Distance (m) VOD = 1954 m/s Distance (m) VOD = 2083 m/s Time (ms) Time (ms) g/cc- 200 mm steel g/cc- 100 mm Cardboard Distance (m) VOD = 2177 m/s Distance (m) VOD = 1592 m/s Time (ms) Time (ms) 302

321 g/cc- 89mm Cardboard 0.15 g/cc- 75 mm Cardboard Distance (m) VOD = 1545 m/s Distance (m) VOD = 1525 m/s Time (ms) Time (ms) VOD Trace using Brass Sensor Diameter Effect (confined steel, ρ = 0.15 g/cc) Distance (m) VOD = 1929 m/sec VOD = 1708 m/sec VOD = 1430 m/sec VOD (m/s) Time (ms) Diameter (mm) 303

322 VOD at Infinite Diameter (0.15 g/cc, steel) Experimental / Ideal VOD Ratio (0.15g/cc, steel) VOD (m/s) VOD intercept ~ 2100 m/s Relative VOD Inverse Diameter (1/mm) Inverse Diameter (1/mm) Confinement Effect (0.15 g/cc) 0.27 g/cc- (SCH 80 cardboard tubes) VOD (m/s) VOD (m/s) Diameter (mm) Diameter (mm) 304

323 Density Effect (steel confinement) 0.15 g/cc- 76-mm, 3.2 m long steel tube VOD (m/sec) mm 50 mm Distance (m) Probe 1 = 1971 m/s Probe 2 = 1969 m/s Density (g/cc) Time (ms) 1.4 ANFO Initiation Test 3600 ANFO mixtures Distance (m) LDRA = 1905 m/s ANFO = 3192 m/s VOD (m/s) Ideal ANFO/LDRA ANFO/EPS Time (ms) Density (g/cc) 305

324 1.4 Modified Aquarium ANFO 0.82 g/cc in 75-mm steel Modified Aquarium 0.15 g/cc (2 Al probes) 600 Distance (m) VOD anfo = 2834 m/s Shock in w ater = 2705 m/s Distance (mm) LDRA = 1806 m/s Shock w ater~1350 m/s Time (ms) Time (microsec) 306

325 Appendix 5.2: Detonation Pressure Techniques Flash Radiography and Densitometry Techniques Much of the observations conducted with flash X rays concern geometrical relationships between shock fronts and contact surfaces from which pressures are inferred. Examples include the experimental work conducted at the Institute de Saint Louis in France (Engineering Design Handbook, 1972) where pressures were estimated by measuring the curvature of a brass plate in the vicinity of a shock front when accelerated by the detonation of the explosive. Flash X ray densitometry is a method that measures the reacted explosive density by comparing the optical density in an X ray film immediately behind and in front of the detonation wave, then calculating the pressure via the conservation equations (mass and momentum). Flash X ray densitometry tends to underestimate pressure values, perhaps as a result of the limitations proper of the actual technology for this kind of measurement. Another approach using flash radiography consists of embedding a micro thin metal foil within the explosive mass in a plane perpendicular to the direction of flow. Triggering two or more flash X ray tubes at pre selected times allows travel distances of this foil to be determined from the film and the corresponding particle velocities readily calculated. The conservation of momentum equation is then applied to determine the detonation pressure. 307

326 The use of the flash X ray technique is not widely spread since the instrumentation is costly, not readily available and its application is restricted to laboratory environments only; nevertheless, it is an established and valid alternative to determine detonation pressure of explosives. Standard Aquarium Technique This section briefly describes the fundamental concepts behind the standard aquarium method and applies these principles towards developing a simpler experimental approach for determining the detonation pressure. The standard aquarium method requires the use of a streak camera and back lighting photographic techniques. Figure A 2 sketches the experimental setup and the film traces required to infer detonation pressure. The light emitted by the detonating explosive will enter the optical system through the camera slit and produce a record in the film whose slope is indicative of the explosive s velocity of detonation (VOD). The non luminous shock transferred into the water is recorded in the film by cutting off the light emitted from an argon bomb. The slope of the corresponding record at a particular point is indicative of the shock velocity in water (Usw) at that point. 308

327 Distance axis Camera Slit Explosive charge Argon light bomb Velocity of Detonation (VOD) Water Aquarium Argon light Shock in water (Usw) Time axis Figure A 2: Sketch of the standard aquarium technique used to measure detonation pressure (left) and the corresponding photographic traces (right) The shock Hugoniot of water is a known empirical relationship between its shock velocity (Usw) and particle velocity (uw) that can be conveniently expressed in the following form: Us w = Co w + S w u w where Cow and Sw are constants which depend on the shock behavior characteristics of the medium and are readily available in the literature (Cooper, 1997). For the case of water, these values are Cow = km/s and Sw = Once the values for the two constants Cow and Sw for water are known and the shock velocity in water at the interface Usw is assessed from the film record, the particle 309

328 velocity (uw) in water is determined from the above equation and thus, the pressure in the water (Pw) is calculated from the conservation of momentum equation given by: P w = ρ Us w w u w where ρw is the density of water and ρw*usw, known as the impedance, is a factor that remains constant for a given material. Knowing the pressure in water at the interface, the detonation pressure of the explosive is inferred by using the impedance mismatch technique, which is a linear approximation of the reflected Hugoniot technique commonly used for evaluating shock interactions. The equation relating the incident detonation pressure (Pd) and the transmitted shock pressure in water (Pw) is as follows: P d = P w ρ w Usw + ρe VOD 2 ρ Us w w where ρe (g/cm 3 ) is the initial density of the explosive. 310

329 Appendix 5.3: PVDF Pressure Records PVDF #1 (Charge Mode, Hardware Integrator at scope) PVDF #2 (Charge Mode, Hardw are Integrator at scope) 8 8 Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) PVDF #3 (Charge Mode, Hardware Integrator at scope) PVDF # 4 (Charge Mode, Hardware Integrator by gage) 8 10 Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) 311

330 PVDF# 5 (Charge Mode, Hardware Integrator by gage) PVDF # 6 (Charge mode, RC circuit soldered at leads) Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) PVDF # 8 (Current mode, 1 Ohm CVR) PVDF # 10 (Current mode, 1 Ohm CVR) 6 6 Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) 312

331 PVDF#11 (Current mode, 1 Ohm CVR) PVDF # 13 (Current mode, 1 Ohm CVR) 6 4 Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) PVDF # 14 (Current mode, 1 Ohm CVR) PVDF # 15 (Current mode, 1 Ohm CVR) 6 8 Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) 313

332 PVDF # 16 (Current mode, 1 Ohm CVR) PVDF # 17 (Current mode, 1 Ohm CVR) Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) PVDF # 19 (Current mode, 1 Ohm CVR) PVDF # 21 (Current mode, 1 Ohm CVR) 5 8 Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) 314

333 PVDF #22 (Current mode, 1 Ohm CVR) PVDF #24 (Current mode, 1 Ohm VCR) 8 5 Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) PLOTDATA ANALYSIS PVDF # 8 (Current mode, 1 Ohm CVR) PVDF # 10 (Current mode, 1 Ohm CVR) Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) 315

334 PVDF # 11 (Current mode. 1 Ohm CVR) PVDF # 14 (Current mode, 1 Ohm CVR) Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) PVDF # 15 (Current mode, 1 Ohm CVR) PVDF # 16 (Current mode, 1 Ohm CVR) 8 8 Pressure (Kbar) Pressure (Kbar) Time (u sec) Time (u sec) 316

335 PVDF #22 (Current mode, 1 Ohm CVR) 8 Pressure (Kbar) Time (u sec) 317

336 Appendix 5.4: CCR Pressure Records: Detonation Pressure in gel and water filled bottles LDRA Detonation Pressure Test 1 LDRA Detonation Pressure Test Pressure (Kbar) Water LDRA Pressure (Kbar) Water LDRA Time (μsec) Time (μsec) LDRA Detonation Pressure Test 3 LDRA Detonation Pressure Test Pressure (Kbar) Water LDRA Pressure (Kbar) Water LDRA Time (μsec) Time (μsec) 318

337 LDRA Detonation Pressure Test 5 LDRA Detonation Pressure Test Pressure (Kbar) Water LDRA Pressure (Kbar) Water LDRA Time (μsec) Time (μsec) 10 LDRA Detonation Pressure Test 7 10 LDRA Detonation Pressure Test Pressure (kbar) Gel LDRA Pressure (kbar) Gel LDRA Time (μsec) Time (μsec) 319

338 10 LDRA Detonation Pressure Test 9 10 LDRA Detonation Pressure Test Pressure (Kbar) Gel LDRA Pressure (Kbar) Gel LDRA Time (μsec) Time (μsec) LDRA Detonation Pressure Test 11 LDRA Detonation Pressure Test Pressure (Kbar) Water LDRA Pressure (Kbar) Water LDRA Time (μsec) Time (μsec) 320

339 LDRA Detonation Pressure Test 13 LDRA Detonation Pressure Test Pressure (Kbar) Water LDRA Pressure (Kbar) Water LDRA Time (μsec) Time (μsec) LDRA Detonation Pressure Test 15 LDRA Detonation Pressure Test Pressure (Kbar) Water LDRA Pressure (Kbar) Water LDRA Time (μsec) Time (μsec) 321

340 LDRA Detonation Pressure Test 17 LDRA Detonation Pressure Test Pressure (Kbar) Water LDRA Pressure (Kbar) Water LDRA Time (μsec) Time (μsec) 8 LDRA Detonation Pressure Test 19 6 Detonation Pressure with CCR 0.15 g/cc in 51-mm steel Pressure (Kbar) Time (μsec) Water LDRA Pressure (Kbar) Test # in Wax in Water in Gel 322

341 Appendix 5.5: CCR Pressure Graphs: Detonation Pressure in Cardboard tubes 0.15 g/cc (50-mm Cardboard tube) 0.15 g/cc (50-mm Cardboard tube) 3 3 Pressure (Kbar) Pressure (Kbar) Time (μsec) Time (μsec) 0.15 g/cc (75-mm Cardboard tube) 0.15 g/cc (75-mm Cardboard tube) 6 6 Pressure (Kbar) Pressure (Kbar) Time (μsec) Time (μsec) 323

342 g/cc (87-mm Cardboard tube) 0.15 g/cc (87-mm Cardboard tube) 6 6 Pressure (Kbar) Pressure (Kbar) Time (μsec) Time (μsec) 0.15 g/cc (100-mm Cardboard tube) 0.15 g/cc (100-mm Cardboard tube) 6 6 Pressure (kbar) Pressure (kbar) Time (μsec) Time (μsec) 324

343 g/cc (25-mm steel pipe) 0.15 g/cc (100-mm steel pipe) 6 6 Pressure (Kbar) Pressure (kbar) Time (μsec) Time (μsec) 0.15 g/cc (Confinement effect) Average Pressure Values 6 Pressure (Kbar) Cardboard Steel Diameter (mm) 325

344 Appendix 5.6: Reflected Hugoniot Technique The fundamental physical laws governing the behavior of elastic waves when they encounter an interface separating two different media can also be applied to shock waves, including detonation shocks. Therefore, when a detonation shock traveling through a given incident medium encounters an interface, it will be partly reflected back to the incident medium and partly transferred into the transmitting medium in a manner that depends on the shock characteristics of the two media. Five basic parameters are needed to fully describe a detonation shock wave: pressure (P), shock velocity (D), particle velocity (u), density (ρ) or specific volume (v = 1/ρ) and specific internal energy (e). The conservation equations of mass, momentum and energy along with an Equation of State (EOS) for the detonation products provide four out of the five equations required to solve the unknowns. The EOS is a function relating energy, pressure and specific volume [e= f(p,v)] and represents all the equilibrium states that are possible for a particular material. The fifth equation, the so called Shock Hugoniot or simply Hugoniot equation, is a relationship between two of the four unknowns present in the conservation of mass and momentum equations (P, D, u and v). The Hugoniot can be defined as the locus of all possible states behind the shock front originating from a certain initial state and is determined experimentally by subjecting the material to varying shock strengths. From the six possible pair combinations (P D, P v, P u, D v, D u and u v), the D u and P u pairs are of particular usefulness in shock interaction. The reflected Hugoniot method is an analytical/graphical technique that makes use of the Hugoniot characteristics of the material and the conservation of momentum equation of the explosive to evaluate shock interaction at the explosive material interface. The technique allows the inference of the detonation pressure of the explosive 326

345 from the shock pressure values recorded by a sensor positioned in the transmitting medium. The Hugoniot relationship in the D u plane has been experimentally determined for many different materials and the parameters defining it are readily available in the literature. It has been proven that a linear relationship exists between both variables, this commonly expressed by an equation of the form: D = C 0 + S u Equation A where D (km/s) and u (km/s) are respectively the shock and particle velocity in the transmitting medium, Co (km/s) a constant representing the Y axis intercept and S a dimensionless constant representing the slope of a linear fit to the experimental data in the D u plane. The different terms of Equation A are best shown in the next diagram: D (km/s) S Co u1 (km/s) The conservation of momentum equation for a detonating explosive is given by: P ρ D u Equation B = 0 327

346 where P (GPa) is the detonation pressure of the explosive, ρ0 (g/cm 3 ) the original density of the explosive, D (km/s) the detonation velocity and u (km/s) the particle velocity behind the detonation front. Given the fact that both ρ0 and D remain fairly constant throughout the detonation process, so does its product, which is normally known as shock impedance (I). Therefore, the conservation of momentum equation (Equation B) for a detonating explosive can be approximated by: P = I u Equation C When plotted in the P u plane, Equation C will show as a liner relationship where the impedance (I) is represented by the slope of the linear function as shown in the following diagram: P (GPa) I = ρ 0.D u (km/s) Combining Hugoniot Equation A and conservation of momentum Equation B results in: P = ρ C ρ Equation D 2 0 u + S u 328

347 Equation D above assumes that the initial state of the material is in repose (i.e. Lagrangian terms: u0 = 0). If the material was in motion (i.e. Eulerian terms: u0 0) the equation becomes: P = ρ C ρ Equation E 2 0 ( u u0 ) + S ( u u0 ) Moreover, since particle velocity is a vector quantity characterized by direction, the mirror images of the previous two equations, representing the reflected shock traveling in the opposite direction, becomes: P = ρ C ( u u + ρ S u u Equation F ) ( 0 ) Equations E and F are also known as right going and left going Hugoniots respectively. The corresponding curves for Equations D, E and F plotted in the P u plane are shown in the following diagram: Equation D (Lagrangian, u o = 0) P Equation E (Eulerian, u o 0) Equation F (Mirror images) u o = 0 u o = 1 u o = 2 u 329

348 The process to infer the detonation pressure of an explosive in contact with a given material whose shock characteristics are known can be determined from solving the previous equations; however, a graphical approach to solve the shock interaction problem is considered a more explanatory approach and will be presented next in the form of an example. Let s assume a detonation front of unknown amplitude generated by the LDRA travels through the explosive, it reaches an interface and is partly reflected back to the explosive and partly transmitted into the Wax material embedding the recording gage. Conditions of continuity require that both pressure and particle velocity remain the same at both sides of the explosive Wax interface. The conditions in the P u plane for the incident right going detonation shock traveling in the LDRA are given by Equation B, while those for the reflected shock going back into the LDRA are given by the mirror image of Equation B, that is the left going shock Hugoniot. In both cases, ρ0 and D are known values. The conditions for the shock transmitted into the Wax material on the other hand should lay in the right going Hugoniot of Wax given by Equation D, where the pressure recorded by the gage (Pgage) embedded in the Wax material is a known value. Given the conditions of continuity of pressure and particle velocity at the explosive Wax interface, the solution to the interface problem (i.e. detonation pressure of the LDRA) has to necessarily lay in the intersection of the right going Hugoniot of Wax and the leftgoing Hugoniot of the LDRA passing through Pgage from where the detonation pressure of the explosive can be determined. An example of the process followed to infer the detonation pressure of the LDRA is depicted in the following graph. 330

349 PLDRA =?? P gage = P epoxy Pressure (GPa) Per( u) Pel( u) Pwr( u) u Particle Velocity (mm/usec) Right Hugoniot LDE Left Hugoniot LDE Right Hugoniot of Wax

350 Appendix 6.1: Quality Control of LDRA (Enaex) PLANILLA DE CONTROL DE ANFO LDE rango aceptación densidad: 0,18-0,22 gr/cc rango aceptación V.O.D.:mayor que 1520 m/s Vel, Densidad cartón 6 OBSV, Fecha de Identificación de la muestra Pour APD 450 análisis Fecha de Producción Hora Batch Nº gr/cc m/seg Muestra llevada a Lab, 15/04/ /04/ por operador promedio /04/ /04/ Muestra llevada a Lab, por operador promedio /04/ /04/ /04/ /04/ : Muestra llevada a Lab, por operador Muestra llevada a Lab, por operador promedio máximo mínimo /04/ /04/ : /04/ /04/ : /04/ /04/ : /04/ /04/ : Muestra llevada a Lab, por operador Muestra llevada a Lab, por operador Muestra llevada a Lab, por operador Muestra llevada a Lab, por operador V.O.D D. POUR promedio 0.18 máximo 0.20 mínimo

351 PLANILLA DE CONTROL DE ANFO LDE rango aceptación densidad: 0,18-0,22 gr/cc rango aceptación V.O.D.:mayor que 1520 m/s Vel, Densidad cartón 6 OBSV, Fecha de Identificación de la muestra Pour APD 450 análisis Fecha de Producción Hora Batch Nº gr/cc m/seg Muestra llevada a Lab, 25/04/ /04/2005 9: por operador Muestra llevada a Lab, 25/04/ /04/ : por operador Muestra llevada a Lab, 25/04/ /04/ : por operador Muestra llevada a Lab, 25/04/ /04/ : por operador V.O.D D. POUR promedio 0.19 máximo 0.19 mínimo /04/ /04/2005 9: /04/ /04/ : /04/ /04/ Muestra llevada a Lab, por operador Muestra llevada a Lab, por operador Muestra llevada a Lab, por operador V.O.D D. POUR promedio 0.20 máximo 0.22 mínimo /04/ /04/ : ***** 28/04/ /04/ /04/ /04/ /04/ /04/ /04/ /04/ Muestra llevada a Lab, por operador Muestra tomada en polvorines por LCC Muestra tomada en polvorines por LCC Muestra tomada en polvorines por LCC Muestra tomada en polvorines por LCC V.O.D D. POUR promedio 0.21 máximo 0.21 mínimo

352 PLANILLA DE CONTROL DE ANFO LDE rango aceptación densidad: 0,18-0,22 gr/cc rango aceptación V.O.D.:mayor que 1520 m/s Vel, Densidad cartón 6 OBSV, Fecha de Identificación de la muestra Pour APD 450 análisis Fecha de Producción Hora Batch Nº gr/cc m/seg Muestra llevada a Lab, 28/04/ /04/ : por operador Muestra tomada en 04/05/ /04/ polvorines por LCC Muestra tomada en 04/05/ /04/ polvorines por LCC Muestra tomada en 04/05/ /04/ polvorines por LCC Muestra tomada en 04/05/ /04/ polvorines por LCC Muestra tomada en 04/05/ /04/ polvorines por LCC Muestra tomada en 04/05/ /04/ polvorines por LCC V.O.D D. POUR promedio 0.21 máximo 0.23 mínimo /04/ /04/ : Muestra llevada a Lab, por operador 04/05/ /04/ V.O.D D. POUR promedio 0.20 máximo 0.20 mínimo /05/ /05/ : /05/ /05/ Muestra llevada a Lab, por operador Muestra tomada en polvorines por LCC V.O.D D. POUR promedio 0.22 máximo 0.22 mínimo

353 PLANILLA DE CONTROL DE ANFO LDE rango aceptación densidad: 0,18-0,22 gr/cc rango aceptación V.O.D.:mayor que 1520 m/s Vel, Densidad cartón 6 OBSV, Fecha de Identificación de la muestra Pour APD 450 análisis Fecha de Producción Hora Batch Nº gr/cc m/seg Muestra llevada a Lab, 03/05/ /05/2005 9: por operador Muestra tomada en 04/05/ /05/ polvorines por LCC Muestra tomada en 04/05/ /05/ polvorines por LCC Muestra tomada en 04/05/ /05/ polvorines por LCC Muestra tomada en 04/05/ /05/ polvorines por LCC Muestra tomada en 04/05/ /05/ polvorines por LCC V.O.D D. POUR promedio 0.22 máximo 0.25 mínimo /05/ /05/ : /05/ /05/ /05/ /05/ /05/ /05/ /05/ /05/ Muestra llevada a Lab, por operador Muestra tomada en polvorines por LCC Muestra tomada en polvorines por LCC Muestra tomada en polvorines por LCC Muestra tomada en polvorines por LCC V.O.D D. POUR promedio 0.21 máximo 0.23 mínimo

354 E XPANSIO N 38 SUR EX PANSION 38 SUR Appendix 6.2: Geotechnical Data of Testing Areas Expansion 42 East (LDRA Test 1 and Test 3) Geotechnical Unit Granodiorite Geotechnical Zone # 1 UCS > n/a FF/m = n/a Wi = n/a Rock Density = 2.52 ton/m3 Conatact OBL Waste Expansion 47 West (LDRA Test 2 and Test 6; Baseline Test B 3) Geotechnical Unit Granodiorite Fortuna Geotechnical Zones 2, 3 and 4 UCS = MPa FF/m = 1 to 8 Rock Density = 2.52 ton/m3 336

355 EXPA NSIO N 38 SUR EXPAN SION 38 SUR Expansion 40 East (LDRA Test 4; Baseline Test B 1) Geotechnical Unit East Potasic Porphyry Geotechnical Zone # 3, 4 and 5 UCS > 180 MPa FF/m = 3 to 7 Wi = Rock Density = 2.62 ton/m3 Transition Zone Porphyry UCS > MPa FF/m = 3 to 6 Wi = Expansion 41 East (LDRA Test 5; Baseline Test B 2) Geotechnical Unit Granodiorite Elena Geotechnical Zones 4, 5 and 6 UCS = MPa FF/m = 5 to 8 Rock Density = 2.52 ton/m3 337

356 Appendix 6.3: Blast and Vibration Data of LDRA Experiments LDRA Test # 1 (Expansion 42 East) COORDINATES TEST # Hole 1 Hole 2 Geo Hole mm NORTH (m) Hole mm 5016 Geo Hole Hole EAST(m) mm blastholes (Geophone 1) 1000 Hole 2 PPV (mm/sec) Hole 1 Hole Time (sec) 338

357 311 mm blastholes (Geophone 2) Hole 5 PPV (mm/sec) Hole 4 Hole Time (sec) mm blasthole (Geophone 2) PPV (mm/sec) Hole 1 Hole 2 Hole Time (sec) Hole # Comments Coordinates Initiation Time Diameter Explosive LDE ANFO Density Delay Arrival East [m] North [m] [mm] [g/cc] [kg] [kg] [ms] [ms] Geophone (56Kg+0Kg))' (39Kg+65Kg)' (28Kg+108Kg)' Geophone (167Kg+0Kg)' (117Kg+196Kg)' (84Kg+362Kg)' Geophone (56Kg+0Kg))' (39Kg+65Kg)' (28Kg+108Kg)' Geophone (167Kg+0Kg)' (117Kg+196Kg)' (84Kg+362Kg)'

358 Hole # Hole Length Charge Length Linear Charge Geo Distance Geo Depth Weight Stem H&P PPV [m] W [kg] H [m] [kg/m] Ro [m] Xo [m] Xs [m] Factor [mm/s] Selected Holes Hole Geo H&PF PPV Holmberg & Persson Predictive Model LDRA Test 1 PPV [mm/s] PPV = 439.2(HPF) R 2 = H & P Factor 340

359 LDRA Test # 5 (Presplit Expansion 41 East) Explosive Conventional (Enaline) LDRA at 0.20 g/cm 3 Geophone PS 2 Pressure Sensors (PS) Geo PS 1 TEST # 5 LDRA IN PRESPLIT ROW 1000 PPV (mm/s) First LDRA hole Time (msec) Hole Coordinates Expl Density Diameter Hole Length Column Length Stem Weight LDRA Nº East North [g/cm 3 ] [mm] [m] [m] [m] [kg] LDRA LDRA LDRA LDRA LDRA LDRA

360 LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA LDRA Hole Coordinates Expl Density Diameter Hole Length Column Length Stem Weight ENALINE Nº East North [g/cm 3 ] [mm] [m] [m] [m] [kg] Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Short hole Enaline Enaline Enaline

361 Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline / Enaline /

362 LDRA Test # 6 (Expansion 42 West) 4420 BLAST COORDINATES TEST # NORTH (m) LDRA HOLES GEO EAST (m) Test # 6 LDRA in 350 mm blastholes Hole # 970 PPV (mm/sec) Time (msec) 344

363 BLASTHOLES TEST # NORTH (m) EAST (m) 350 mm (13 3/4 inch) Geophone 0.20 g/cc Hole Coordinates East North Hole Length Weight Column Length Linear Charge Distance Geo Depth Geo Stem # [m] [m] [m] [kg] H [m] [kg/m] Ro [m] Xo [m] Xs [m]

364 Hole HandP PPV # [mm/s] Holmberg & Persson Predictive Model LDRA Test 6 PPV (mm/sec) PPV = (HPF) R 2 = H & P Factor 346

365 Appendix 6.4: Blast and Vibration Data of Baseline Experiments Baseline Test B 1 (Expansion 40 East) 347