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Influence of bolt-grout bonding on MCB conebolt performance
International Journal of Rock Mechanics & Mining Sciences 49 (2012) 165–175
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International Journal of Rock Mechanics & Mining Sciences journal homepage: www.elsevier.com/locate/ijrmms
Technical Note
Influence of bolt-grout bonding on MCB conebolt performance M. Cai a,n, D. Champaigne b a b
School of Engineering, Laurentian University 933 Ramsey Lake RoadSudbury, Sudbury, Ont, Canada P3E 2C6 Mansour Mining Technologies Inc., Sudbury, Ont, Canada
a r t i c l e i n f o Article history: Received 27 November 2010 Received in revised form 25 June 2011 Accepted 11 November 2011 Available online 29 November 2011
1. Introduction 1.1. Rockburst and rockburst damage As the depth of mining and underground construction increases, stress-induced rock fracturing is inevitable and in some cases the rocks can fail violently, leading to seismic events and rockbursts. A rockburst is defined as damage to an excavation that occurs in a sudden or violent manner and is associated with a seismic event [1]. Many hard rock mines in Canada, China, Chile, South Africa, Australia, Sweden, and other countries, and some ¨ deep civil tunnels in Switzerland (e.g., Loetschberg, Gotthard, Furka Base Tunnels), China (e.g., Erlang Mountain Highway Tunnel, Jingping-II intake tunnels), and Peru (e.g., Olmos TransAndean Tunnel) have experienced rockbursts to various degrees. Rockbursts can be classified into three types [2]: (a) fault slip burst, (b) pillar burst, and (c) strain burst. When the mininginduced stresses exceed the strength capacity of the unsupported or supported rock, even temporarily, failure is initiated or triggered. Typical rockburst damage to underground excavations include stress-induced rock fracturing, bulking of roof and sidewalls, floor heave, shearing of rock, rock falls and ejections, etc. Rockbursts can cause fatalities and injuries to workers, damage to mine infrastructure and equipment, and substantially increase investment risk. For example, a magnitude mN ¼ 3.8 (Nuttli magnitude) seismic event, which happened on January 6, 2009, caused extensive damage to four levels at the Kidd Mine in Timmins, Ontario, Canada. One drift supported by standard rock support with mesh and rebar collapsed totally and a few pieces of mining equipment were damaged. Luckily, the event occurred at a time as the nightshift was going off duty and heading to surface and thus no injuries were reported. For obvious reasons, we must
n
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rely on proper rockburst support to safeguard workers and investments. It is imperative to design and utilise proper rockburst-resistant support when mining or tunnelling at depth.
1.2. Rockburst support Rock support in burst-prone ground differs from conventional rock support in shallow ground where controlling gravityinduced rock falls is the main concern. Rock support in burstprone ground needs to resist dynamic loading and large rock dilation due to rock failure. Much research has been conducted (e.g., [1–3] and many others) to address the rockburst problem. These studies recommend and provide design of rockburst support as a means to mitigate rockburst damage. A rock support system helps the rock support itself [4,5]. This view of rock support function was derived from the observation of rock support under static loading. In general, the role of rock support in underground construction can be classified into three major functions—reinforce, retain, and hold [1]. The goal of reinforcing the rockmass using rockbolts and/or rebars is not only to strengthen it, thus enabling the rockmass to support itself but also to control the bulking process as rockbolts/rebars prevent fractures from propagating and opening up. Under high stress conditions, fractured rock between the reinforcing/holding elements may unravel if not properly retained. An important function of rock support is realized through retaining the fractured surface rock so as to confine rocks further away from the opening to prevent further fracturing. Widely used retaining elements include wire mesh, shotcrete, or cast-in concrete liners. The holding function is needed to tie retaining elements of the support system and loose rock back to stable ground, to dissipate dynamic energy due to rock ejection and rock movement, and to prevent gravity-driven falls of ground. Under dynamic loading condition, the rock support system must be able to absorb energy to survive the rockburst event. When rockburst damage is
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anticipated, yielding holding elements such as conebolts and friction bolts must be used in the support system. Otherwise, severe damage or complete failure of the excavation could result. 1.3. Need for yielding support in bursting ground 1.3.1. Rock dilation due to rock fracturing When the rock fractures, its volume increases significantly, and the phenomenon is often referred to as dilation or bulking. Fig. 1 shows the results of a numerical simulation using ELFEN [6], an FEM/DEM combined code that can properly simulate rock fracturing processes. Because a displacement loading boundary is applied, the rock failure process is stable and there is no violent rock ejection and disintegration. When the peak strength is reached, the rock starts to dilate and an 8% to 10% volume increase is observed in this case when the vertical strain reaches about 2%. In other words, when a 2 m thick ring of rock near the excavation boundary fails in this fashion with an 8% to 10% volume increase, the maximum wall displacement would be 160 to 200 mm, which is far beyond the deformation capacity of most standard ground support systems. Rock bulking resulting from disintegration of rock due to violent rock fracturing can lead to even greater volume increase. When rock fractures under low confinement conditions its volume increases significantly. Kaiser et al. [1] proposed the concept of bulking factor and used it to estimate maximum wall deformation. One typical example of rock bulking is floor heave in an underground excavation. In South Africa, complete drift closure due to floor heave has been encountered during severe rockbursts. In Canadian mines, bulking in the floor is commonly observed on the order of 20% to 40% of the fractured rock zone (if unconfined), typically leading to floor heave of 0.2 to 0.4 m. 1.3.2. Yielding design principle When mining at depth, the excavation-induced stresses around underground openings are so high that rock fracturing is inevitable. The stresses are further elevated when the rock is loaded dynamically. In many situations, it is no longer possible to increase the load carrying capacity of the reinforced rock system economically, and the support behaviour must be fundamentally changed to allow for yielding. Jager et al. [7] showed that the dilation pressure resulting from brittle failure in a deep tunnel sidewall exceeded 0.4 MPa. When the rock fracturing process is dynamic, it is likely that the pressure will be considerably higher, exceeding the capacity of most practical rock support systems.
Fig. 1. Stress-displacement relation of a rock sample with dilation behaviour.
For reference, the pressure (support pressure) to the excavation surface that can be supplied by a moderately dense pattern of rockbolts is about 0.1 MPa. To provide a support pressure of 0.4 MPa, very thick cast-in concrete would be required, which is clearly not economically practical for most mining applications. In general, rock load or pressure to the support system decreases with increasing rock deformation. If the rock support system is able to yield in a controlled fashion, it reaches a dynamic equilibrium during the deformation process and the system will eventually reach a new static equilibrium. Hence, the key point is that the rock support system must allow the rock mass to deform, which means that the rock support system must be deformable. A yielding support system should tolerate large tunnel convergence without ‘‘self-destruction’’ of the system while providing necessary support to ensure safety and maintain serviceability of the tunnel. A yielding rock support system is a system in harmony with the surrounding rock. Fortunately, rock bulking can be controlled effectively using appropriate rock support elements, but the rock support system used must be yieldable to accommodate large deformation. Field observation in a few mines confirmed that when yielding support is installed, the wall dilation can be effectively controlled (Fig. 2). Cook and Ortlepp [8] first suggested the use of yielding rockburst support in the deep gold mines in South Africa. Smooth bar yielding bolts were developed and tested by Ortlepp in 1969. Over the years, various rock support products have been developed and used in practice. Among them, conebolts stand out as good holding and energy absorbing elements in a rockburst support system. In the following, we present a brief review of the development of various types of conebolts, identify debonding as a most important issue for the effective performance of conebolts, describe a recent development to address the debonding problem, and compare the static and dynamic test results to reveal the influence of bolt-grout bonding on the behaviour of conebolts.
Fig. 2. Large wall bulking contained by yielding support.
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2. Historical development of conebolts 2.1. South African conebolts The concept and idea of conebolt was firstly developed in South Africa. The conebolts were groutable yielding tendons developed by the Chamber of Mines Research Organisation (COMRO) in 1987 [9] for use in cement grouted holes. When subjected to static loading, the cone functions as a wedge-style mechanical anchor similar to standard mechanical rockbolts. However, when subjected to dynamic loading, the cone yields or ploughs through the grout, thus absorbing the dynamic energy through controlled deformation. Ease of installation was one of the important considerations in the initial design of the South Africa conebolt. Because cement grouting of shepherd’s crook tendons was generally practiced in South Africa, this consideration led to the use of cement grout for the conebolts. To facilitate the bolt’s yielding or ploughing through the grout, little or no bonding between the grout and the bolt was permitted. Hence, debonding between the bolt and grout was required. Debonding was achieved by coating the bolt with wax. A special saponified wax was developed and patented. The wax, which adheres well to steel, has an appropriate melting point and viscosity for ease of applying a thin layer of coating to the bar. One issue related to the use of wax was that its stability was affected in the presence of ammonia. If the wax coated bolts were stored for long periods of time underground in areas contaminated by blasting fumes, the wax could potentially lose its effectiveness. Ortlepp [10] conducted a quarry site blasting test and demonstrated the effectiveness of South African conebolts in absorbing kinetic energy. Cement grouted conebolts have been used in a few Australia mines such as Big Bell Mine [11]. However, conebolts were reluctantly accepted in South African mines, even in deeplevel rockburst prone mines. This was largely due to the high initial cost [12]. South African conebolts were first tested in Canada in 1994 [13] using cement grout and polyester resin (injected with quicksetting resin grout). The cement grouted conebolt had a maximum displacement of 240 mm and the load-displacement curve resembles that of a steel stretching curve, indicating that there was little cone movement during pull-out test. In resin grout testing, the conebolt could not shear through the resin and failed after only about 100 mm of displacement. To our knowledge, no
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mine in Canada systematically used the South African conebolts in its operation. The main reason being that resin injection was not practical for use at the time and cement grouting was time consuming and required a third pass to torque the bolts once the grout was cured. 2.2. Modified conebolts (MCB) The increase in violent ejection failures in 1999/2000 at Brunswick Mine in Canada led to the urgent development of a complete yielding support system [14]. The South African conebolts were modified to accommodate resin-grouting applications in Canada. Norand Inc. modified the conebolt head and added a blade for the purpose of resin mixing and the new bolt was called Modified Cone Bolt (MCB). This new tendon can shear through the surrounding column of polyester resin grout. It is a smooth bar threaded at its outer end, with a forged cone and mixing blade at the other end. Grease was used as the debonding agent. Initial dynamic testing was conducted using a one tonne dynamic drop facility designed by MIRARCO (Laurentian University, Canada) for the Noranda Technology Centre (NTC) (a modified version of the NTC drop test facility is now known as the CANMET drop test facility shown in Fig. 7). At Brunswick Mine, MCB38 conebolts (grease-debonded, installed in 38 mm boreholes) at 1 m spacing with 150 150 mm2, 9.5 mm thick domed plates, were used with #0 gauge mesh straps (8.23 mm wire diameter, 10 10 cm2 aperture, typical sheet size: 30.5 365.8 cm2) to overlay the diamond mesh (5 cm aperture, 5.0 mm diameter), installed on a second-pass system [14,15]. A close-up view of the rockburst support system installed at Brunswick Mine can be seen in Fig. 3. There was a distinct boundary in one drift (tunnel) that defined stable and unstable grounds in Fig. 3a. This boundary was called the ‘‘extreme edge’’ at the mine site. Beyond the extreme edge, there were only standard rock support systems installed. A series of rockburst events that occurred between October 13 and 17, 2000 severely damaged the ground where standard supports were installed, but the section that was supported by the rockburst support system suffered no damage. On Oct. 13, 2000, a Nuttli magnitude mN ¼2.5 rockburst event caused the collapse of an intersection involving approximately 1860 t of material. A small portion of the 326 crosscut immediately east of the intersection had MCB conebolts installed (a close-up view of the 326 crosscut can be seen in Fig. 3b). The right-hand side of the tunnel was supported by 2.3 m
#6 Chain-link mesh+ #0 mesh strap + MCB conebolts(17.2 mm)
Chain-link mesh
Thread-bar +
#6 Chain-link mesh
shotcrete + #9
+ #0 mesh strap +
gauge weld mesh
MCB conebolts (17.2mm)
+ end anchored rockbolts
Fig. 3. ‘‘Ground truthing’’ of modified conebolts at Brunswick Mine (photo courtesy: Brunswick Mine and B. Simser).
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long MCB conebolts in a 1 1 m2 pattern with #0 gauge mesh straps and chain-link mesh. The left-hand side of the tunnel was supported only by standard support (rebar þshotcrete þ#9 gauge weld mesh). The rockburst support system installation had not been fully completed at the time of the bursts as the chain-link mesh was temporarily suspended with 0.9 m end anchored rockbolts and the next step of installing MCBs and mesh straps had not been completed. The foreground of the area was supported by a rebar and steel fibre reinforced shotcrete. After the rockburst event, it was revealed that the right-hand side of the tunnel suffered no visible damage at all while the rest of the area suffered moderate to severe damage. The left-hand side showed excessive large bulking. While this was not intended as a ‘‘field experiment,’’ the observed rockburst damage to different sections clearly showed that a well-designed rockburst support system can effectively limit or eliminate rockburst damage to underground openings. The excellent performance of the MCB supported tunnel section held the ground so well that many people speculated that the rock might not be hit hard as the rest of the area. However, close onsite examination revealed that the MCBs had fulfilled their role in dissipating dynamic energy as some conebolts displaced as much as 180 mm [14]. Subsequent use of MCBs based rock support system at the mine site performed exceptionally well. Following the success at Brunswick Mine, MCBs have enjoyed a slow but gradual acceptance by some Canadian mines having rockburst problems, and have been successfully applied in their operations.
3. Influence of bolt-grout bonding on conebolt performance 3.1. The debonding issue The South African (SA) conebolts have a 0.1 mm thick saponified wax coating as the debonding agent. Approximately 10 months after receiving the SA bolts at Big Bell Mine in Australia,
it was noted that some bolts exhibited a substantial deterioration or flaking of the protective wax coating [11]. The original MCB was greased to reduce the bonding effect between the grout and the bolt. However, it was noted that the spinning action of the bolt to mix the resin stripped off almost all the grease coating, leading to unwanted bond between the bar and the resin. If wax or grease is not working as intended as a debonding agent, then bonding between the bolt and the grout can reduce the effectiveness of conebolts. The question is how much bonding exists between the bolt and the grout? This can be found out by static and dynamic tests in the laboratory or in the field. Eight South African conebolts (22 mm diameter), four with intact wax coating and four with the wax coating physically removed, were tested at Big Bell Mine to compare the bonding behaviour of bolt and the cement grout (sc E35 MPa), with their cones cut off [11]. The bolt lengths were between 1.0 and 1.5 m. It was seen that for bolts with wax coating physically removed, the peak loads vary between 42 and 92 kN and the residual loads vary between 35 and 70 kN, at a displacement of 150 to 200 mm. For the bolts with wax coating, the peak loads vary between 40 and 91 kN and the residual loads vary between 17 and 25 kN, at a displacement of 150 to 200 mm (see Fig. 4). The conclusion was that the wax coated smooth bar had on average 25% and 50% less peak and residual bond strength per metre of embedment than bolts with wax removed, respectively. In our opinion, these test results suggest that wax coatings are not effective in debonding the conebolts from the cement grout. Cai and Champaigne [16] conducted a series of similar pull tests using steel pipes. Greased MCBs (17.2 mm diameter, 1.2 m long) were installed in the steel pipes (inner diameter 33 mm, wall thickness 5.2 mm) using resin. The bolts were fully encapsulated. After the resin was cured, the cone section (about 0.1 m) was cut off and the remaining bolt was pull tested until the bolt started to slide in the resin. For comparison purposes, the greased bar test results are presented in Fig. 4 with the results of Player [11]. In our test, the displacement range was limited to about 35 mm, due to stroke limitation of the ram. It was found that for an embedment length of 0.94 m, the maximum load recorded was 127 kN. This load
Fig. 4. Load-displacement curves of smooth bars with different debonding agents.
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is in fact close to the yield point of the steel bar. Another bar tested with a grease coating had a peak load of 81.6 kN. This is also very high, considering the short embedment length of 0.94 m. Similarly to the South African wax coated conebolts, the pull test results showed that the grease coating is not effective in debonding the MCBs from the resin grout. It can then be concluded that the grease coating was stripped off of the MCBs during the installation process of spinning the bolt through the resin at the manufactures recommended 30 to 40 revolutions.
3.2. The next generation conebolt—MCB33 It is obvious that an effective debonding agent is required in order to improve the static and dynamic performances of MCBs. No better wax or different grease debonding agent could be identified or developed. As a consequence, a new debonding agent in the form of a plastic (PVC) sleeve installed over the shaft of the MCB conebolt was developed [16]. The lab test results show that this debonding agent can eliminate the bonding force effectively. As shown in Fig. 4, the load for a 0.94 m section of smooth bar to slide in resin grout is only 4.5 to 6.8 kN if a bar is
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fully covered by a plastic sleeve. In one test, about 20 mm of the bar was not covered by the sleeve and it resulted in a peak load of 22.7 kN. The residual load of this bar reached to 9 kN at a displacement of 35 mm. The effectiveness of the plastic sleeve as a debonding agent was confirmed. The new design of the MCB33 conebolt is shown in Fig. 5. The bar diameter is 17.2 mm. The bolts are manufactured from C1055 modified steel with a typical yield strength of 517 MPa, a typical ultimate tensile strength of 745 MPa, and a typical elongation capacity of 16%. The new debonding agent, heat shrink plastic sleeve (20 mm diameter), is applied to the whole bar shaft except the cone section and thread section. The sleeve is durable and sleeve damage is unlikely if the bar is handled without abuse. The sleeve is tightly around the bar shaft and there is no sleeve damage or movement during bolt installation. There are currently two different cone sizes in use—MCB38 for 38 mm boreholes and MCB33 for 33 mm boreholes. Because regularly used rebars and rockbolts in underground mines all use 33 mm boreholes, it makes much more economical sense to use the MCB33 conebolts, i.e., one bit and one resin. This eliminates extra bits and resin sizes from inventory, and greatly simplifies rock support installation. In addition, it is possible to
Fig. 5. Fully debonded MCB33 conebolt (patented).
180 Rebar
160
MCB33 (grease debonded) S.A conebolt
140
(GRC test result)
Load (kN)
120 100 80 MCB33 (PVC debonded)
60 40 20 0 0
20
40
60
80 100 Displacement (mm)
120
140
160
Fig. 6. Load-displacement relations of MCB33 conebolts with different debonding agents obtained from in-situ pull tests.
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install all standard and rockburst support systems in one pass, which is desired for rapid drift development. The final resin matrix quality is important to ensure that the dynamic response of the MCB consistently performs on a repetitive basis and as designed. A number of factors can affect the mixing performance of resin cartridges: insertion rate, rotational speed, bolt to hole annulus, resin viscosity, etc. In the case of MCB33, the borehole diameter was reduced from 38 mm down to 33 mm, therefore increasing resin mix-ability considerably. In the following discussion, we compare in-situ pull test and dynamic drop test results of MCB33 conebolts with different debonding agents. The goal is to understand how the conebolt performance can be affected by the method of debonding. 3.3. Comparison of static behaviour with different debonding agents Extensive in-situ pull tests were performed at five mines in Canada to evaluate the load-displacement response of MCB33 conebolts with grease and PVC sleeve as the debonding agent. The MCB33 conebolt total length is 1.7 m and the resin embedment length is roughly 1 m. A few test results are presented in Fig. 6. For comparison, load-displacement curves of a 20 mm rebar and a 16 mm South African conebolt are also shown in the same figure. The cement grouted conebolt (tested by GRC (Geomechanics Research Centre, Laurentian University, Canada) [1]) shows similar stiff behaviour as rebar until the load reaches 125 kN. After the peak load, there is an initial drop in load followed by an ever increasing load. The response of grease debonded MCB33 conebolts show a linear trend before the load reaches about 100 kN. When the load is further increased, the bolt debonds from the resin and the displacement rate increases drastically. Compared with the response of rebar and the South African conebolt, grease debonded conebolts are less stiff but the displacement that correspond to bond breakage (load deviation from linearity) is small, generally less than 5 mm. The reason that the grease debonded MCB33 conebolts are less stiff than the South African conebolt is that the bolts are not fully encapsulated with resin in the boreholes. For fully encapsulated bolts, the initial stiffness will be similar to that of rebar. The response of PVC sleeve debonded MCB33 conebolts is characterised with a gradual increase of load with deformation and a plateau is reached when the displacement is in the range of 20 to 40 mm. After the plateau, the load continues to increase. Yielding of the steel was noticed from the reduction of bolt diameter. The tests were stopped because the displacement reached the stroke limit of the ram. Compared with the behaviour of grease debonded MCB33, PVC sleeve debonded conebolts are less stiff before the plateau is reached. The ultimate loads of the bolts are not affected by the type of debonding agents used, but the ultimate deformations of PVC debonded conebolts are slightly larger (most likely due to some cone movements). The behaviour of the MCB33 conebolts with PVC debonding agent is attributed mostly to the cone resistance in resin without the interference of frictional forces between the bolt and resin. This feature may not be very desirable under slow loading rates, however under high loading conditions, as will be the case under seismic load, complete debonding of the shaft of the bar from the grout can enhance the cone plough performance and is demonstrated in the dynamic drop tests described next.
rockbursts. However, this difficulty does not stop researchers from finding practical means to test the dynamic behaviour of rock support elements. Drop test facilities, which can perform repeatable dynamic loading, have been built in Canada, South Africa, Sweden, and Australia to test the dynamic properties of reinforcing and holding elements. The drop test is characterised by its ease to repeat tests at a relatively low cost. The drop test facility employed in this study is shown in Fig. 7. The facility is used for rockbolt testing in steel tubes. The drop weight is lifted with an electromagnet, which in turn is lifted by a pair of cranes mounted in parallel on the top of the machine. By cutting the power to the magnet, the weight free falls onto the sample. The maximum weight capacity is 3 t, and the maximum input energy is 58.9 kJ at a drop height of 2 m. A target for measuring the plate displacement is installed at the end of the bolt, on the remaining threaded section. Displacements are measured using line-scan cameras. A rod is connected to the upper end of the bolt test specimen and a second target is fixed at the top of the rod to measure the cone displacement. To ensure that the bolts are properly installed into the steel tubes, CANMET also developed a special installation rig, which can control the insertion time and the number of bolt revolutions to minimise the difference in resin mixing quality between samples.
3.4.2. Dynamic behaviour of non-greased and grease debonded MCB conebolts One of the key factors, which influence conebolt performance is the type of debonding agent used. Another important factor that affects conebolt performance is the mixing of the resin during bolt installations. The normal installation procedure for MCB conebolt is to spin the bolt to mix the resin as soon as the bolt is inserted into the tube (or borehole in the field). Spinning is continued at a constant insertion rate until the bolt reaches the end of the tube (borehole). Immediately the bolt is spun another 30 to 40 revolutions to further mix the resin and the installation is completed. We also investigated an alternative installation method, i.e., to push the bolt into the tube as much as possible before spinning the resin for 30 to 40 revolutions. Sample autopsy by splitting the tube to examine the mixed resin indicated that
Steel tube
Drop weight
Plate
Test bolt
3.4. Comparison of dynamic behaviour with different debonding agents 3.4.1. Dynamic drop test method and facility It is almost impossible to have a testing system that can simulate the real dynamic loading to a bolt resulting from
Fig. 7. CANMET drop test facility and test setup.
M. Cai, D. Champaigne / International Journal of Rock Mechanics & Mining Sciences 49 (2012) 165–175
both the normal and the alternative installation methods can mix the resin satisfactorily. First, we conducted six drop tests of MCB33s, of which three were greased and three non-greased. The bolts were installed in the steel tubes using the alternative installation method. A mass of 1115 kg and a drop height of 1.5 m were used in this test series, and each impact produced 16 kJ of input energy. The test results are presented in Fig. 8, showing the plate displacement (in mm) and the relative cone displacement and steel stretching (in %) to the plate displacement. In the first impact loading (Drop-1), the cone plough, expressed as a percentage of the cone movement to the total deformation measured at the plate, varied from 11.2% to 97.5% (with an average of 57%) for non-greased bolts, showing a large variability. For greased debonded bolts, the cone plough varied from 58.1% to 95.4% with an average of 74.5%. In the second impact loading (Drop-2), for non-greased bolts, the cone plough varied from 1.1% to 91.1% with an average of 31.1%.
For greased debonded bolts, the cone plough varied from 43.1% to 72.0% with an average of 61.0%. One non-greased bolt and all three greased debonded bolts were dynamically loaded the third time (Drop-3). Again, the test results showed large variability in terms of cone plough behaviour. For one greased bolt, the cone was ‘‘locked’’ in place and 100% of the plate displacement came from steel stretching. As can be seen from the results, with respect to cone movement, grease debonded bolts seem to perform only marginally better on average (65% of total displacement is due to cone movement in all drops) than non greased bolts where cone movement accounts for 48% of total displacement. This test result confirms that grease is not an ideal debonding agent for MCB conebolts. In a separate dynamic drop tests conducted at CANMET, St-Pierre et al. [17] also found that the influence of grease on the bolt behaviour was not obvious. In two of their tests, the cones were cut off and a drop weight of 1016 kg from 0.5 m (5 kJ input energy) was used to test the dynamic capability of the
Non greased Drop-1
171
Greased
100%
16 kJ
80% 60% 40% 20%
0% Plate displ.(mm)
116
101
161
283
126
129
Steel stretch %
88.8%
37.6%
2.5%
4.6%
30.2%
41.9%
Cone plough %
11.2%
62.4%
97.5%
95.4%
69.8%
58.1%
Non greased Drop-2
Greased
100%
16 kJ
80% 60%
32 kJ (accumulative)
40% 20%
0% Plate displ.(mm)
79
89
93
130
122
150
Steel stretch %
8.9%
98.9%
98.9%
56.9%
32.0%
28.0%
Cone plough %
91.1%
1.1%
1.1%
43.1%
68.0%
72.0%
Non greased Drop-3 16 kJ
Greased
100% 80% 60%
48 kJ (accumulative) 40% 20% 0% Plate displ.(mm)
102
137
92
120
Steel stretch %
28.4%
46.7%
100.0%
37.5%
Cone plough %
71.6%
53.3%
0.0%
62.5%
Fig. 8. Comparison of dynamic behaviours of non-greased and grease debonded MCB33 conebolts, showing the percentage of cone plough and steel stretching at 16 kJ impact energy.
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energies of 16 kJ and the results are compared with the results of previous tests using grease as the debonding agent. All the tested bolts were the same type (MCB33), installed in the same steel tube using the normal installation procedure with the same type of resin. Dynamic drop test results showed that on average 99.6% of the plate displacement of PVC sleeve debonded bolts was from cone plough on the first drop as compared to 82.9% in the old design using grease as the debonding agent (Fig. 9). In the second drop, the cone plough percentages for the new and old designs are 100.0% and 29.2%, respectively. In the third drop, the cone plough percentages for the new and old designs are 99.6% and 4.7%,
smooth bars without cones. The resin embedment length was estimated at 1.2 m. The bar, which was greased sustained six drops with a total input energy of 35 kJ and the non-greased bolt accepted five drops with a total input energy of 30 kJ. Note that 5 kJ impact energy is large enough to break a fully bonded rebar at the threads.
3.4.3. Dynamic behaviour of PVC sleeve debonded MCB33 conebolts Dynamic drop tests were conducted to validate the new design of debonding agent using PVC sleeve. Firstly, plastic sleeve debonded conebolts (17.2 mm diameter) were tested at impact
PVC sleeve Drop-1
Greased
100%
16 kJ
80% 60% 40% 20% 0%
Plate displ.(mm)
334
385
484
201
127
153
150
161
101
Steel stretch %
0.3%
0.0%
0.0%
4.5%
22.0%
9.8%
9.3%
2.5%
54.5%
Cone plough %
99.7%
100.0%
100.0%
95.5%
78.0%
90.2%
90.7%
97.5%
45.5%
PVC sleeve Drop-2
Greased
100%
16 kJ
80% 60%
32 kJ (accumulative)
40% 20% 0%
Plate displ.(mm)
297
264
386
172
98
85
90
93
Steel stretch %
0.0%
0.0%
0.0%
0.0%
58.2%
100.0%
96.7%
98.9%
Cone plough %
100.0%
100.0%
100.0%
100.0%
41.8%
0.0%
3.3%
1.1%
PVC sleeve Drop-3
Greased
100%
16 kJ
80% 60%
48 kJ (accumulative)
40% 20% 0%
Plate displ.(mm)
350
359
85
Steel stretch %
0.6%
0.3%
95.3%
Cone plough %
99.4%
99.7%
4.7%
Fig. 9. Comparison of dynamic behaviours of PVC sleeve debonded and grease debonded MCB33 conebolts, showing the percentage of cone plough and steel stretching at 16 kJ impact energy.
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repeated shocks is significantly reduced. For example, a greased bolt resulted in 54.5% of steel stretching in the first drop. With this degree of steel stretching, the bolt did not survive the second drop. The new design of the conebolts eliminates this problem completely. In fact, two new bolts received a third energy input of 16 kJ and the cone ploughs were 99.4% and 99.7%, respectively. The test results of grease-debonded conebolts shown in Figs. 8 and 9 indicate that the installation method may have a minor influence on the effectiveness of grease as a debonding agent. Compared with the results using the alternative installation method shown in Fig. 8, the normal installation method leads to more steel stretching in the second drop as shown in Fig. 9. This is because the normal installation procedure causes more spinning of the bolt, thus has a larger potential of stripping more grease off the bolt during installation. When the debonding is not effective, consecutive loading can stiffen the bolt and the bolt relies on more steel stretching to absorb the impact energy. Three new MCB33 conebolts with plastic sleeves were also tested at 26 kJ impact energy (1784 kg weight, 1.5 m drop height). These tests resulted in 100% and 99.9% cone ploughs in the first and second drops, respectively (Fig. 10). This means that almost all of the impact energy was absorbed by cone plough. In our previous drop tests of grease debonded MCB conebolts, the highest impact energy tested was 22 kJ, which resulted in significant steel stretch. Close to perfect cone plough at 26 kJ was very encouraging and it was decided to use the last available bolt to test at an even higher energy level. The bolt was subjected to a 33 kJ impact load from a single drop (2229 kg weight, 1.5 m drop height) and the bolt survived. Cone plough constituted 76.3% of the total
respectively. It is seen that the new design allows the cone to plough through the resin very effectively while the old design relied increasingly on steel stretch to absorb the impact energy of repeated dynamic loading. It is of interest to explore why greased MCBs performed worse on the second drop. On the second drop one can clearly see more steel stretch and less cone plough in Fig. 9b. To explain this phenomenon, we need to refer to Fig. 4. As explained before, grease could be stripped off during the bolt installation, allowing the shaft of a MCB to come directly in contact with the resin. The smooth bar behaves more or less like the curves of ‘‘No wax’’ shown in Fig. 4. Although the curve was obtained in cement grout, we speculate that the response of a clean smooth bar in resin grout will be similar. Load resistance increases as the bars slide in the grout. The first drop test of greased MCB conebolts resulted in 149 mm plate displacement (average value). On the second drop, the frictional resistance along the bolt shaft is higher than the resistance in the first drop. With this effect, coupled with a change in cone resistance, the cone plough action became less effective in the second and subsequent drops. As mentioned above, greased MCBs could withstand 16 kJ of impact energy 3 to 4 times before the bolt broke. This means that the bolts become stiffer after repeated dynamic loading and the cause of this phenomenon can be attributed to the in-effectiveness of grease as a debonding agent. The implication of this observed phenomenon is that greased MCBs can become ‘‘stiff’’ after an initial moderate dynamic loading (e.g., 16 kJ). The same conclusion can be reached by examining the non-greased bolt test data (Fig. 8). With less cone plough and more steel stretching, their ability to withstand
26 kJ
33 kJ
100%
Drop-1
80% 60%
40% 20% 0% Plate displ. (mm)
602
597
729
531
Steel stretch %
0.0%
0.0%
0.0%
23.7%
Cone plough %
100.0%
100.0%
100.0%
76.3%
26 kJ Drop-2
100% 80% 60%
40% 20% 0% Plate displ. (mm)
173
393
636
Steel stretch %
0.3%
0.0%
Cone plough %
99.7%
100.0%
Fig. 10. Percentage of cone plough and steel stretching of MCB33 conebolts at 26 kJ and 33 kJ impact energy.
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displacement and steel stretch was 23.7%, which represents only 5.6% steel strain. This bolt did not survive the second drop at the same energy input level (33 kJ). It should be noted that the third bolt tested at 26 kJ did not fail in the second drop. The plate hit the floor so that displacement measurement could not be correctly conducted. Fig. 11 presents the relationship between bolt displacement and absorbed impact energy. Data other than MCBs are obtained from Player et al. [18] and Li and Charette [19]. It is seen that toe anchored thread bars fail at the thread and absorb very little energy ( o2 kJ). The fully bonded threadbars are tested using load transfer by separating the bolt in the middle of the installed pipe. The debonded threadbars have about 1.6 m debonded section, which allow the steel to stretch. All data points are from the first drop loading. Four threadbars failed but there were no conebolt failures on the first drop. Very good linear correlations between steel strain and energy absorption can be seen from the threadbar data and the D-bolt data. It is observed that the cement-grouted South Africa conebolts and at least two of the old MCB conebolts follow the trend line of threadbar, indicating that the dominant energy absorption mechanism of those conebolts had been steel stretch with very little cone plough. On the other hand, the new MCB design allows the cone to plough much more efficiently, and their capacity to absorb dynamic energy is thus substantially increased. Using MCB conebolt and threadbar data as an example, we identified in Fig. 11 three zones showing three distinct mechanisms of energy absorption by rockbolts. In Zone A, pure cone plough or friction resistance is responsible for energy dissipation. The energy dissipation capacity of most PVC sleeve debonded
conebolts fall into Zone A. The energy dissipation capacity of friction bolts such as Split Set will fall into Zone A as well, but at a much lower energy dissipation rate (indicated by the slope of the line). In Zone B, the energy is dissipated through a combination of steel stretching and cone plough (or other yielding mechanism). Most grease debonded MCB data points fall into this zone. In Zone C, including the boundary indicated by the blue solid line, energy absorption is achieved by pure steel stretching. The energy dissipation rate is hence dependent on the yielding mechanism. When steel stretching is involved in energy dissipation, high energy dissipation rates can be achieved. The highest energy dissipation rates can only be realized by deploying high strength steel (e.g., D-bolt) and/or using larger diameter bars, at an additional material cost. For other yielding bolts such as Roofex [20], their behaviour can be classified into Zone A or Zone B accordingly (the slope of the zones may be different than the ones shown in Fig. 11 because MCB data are used here). It is critical to differentiate the deformation source of the rockbolts shown in Fig. 11. For MCB conebolts, the deformation comes from the plate displacement, reflecting all the deformation due to cone sliding and steel stretching. For threadbar and D-bolt, the deformation comes from the steel stretching from a free deformation section between two internal anchor points/sections. Dynamic loads were directly applied to the plates for MCBs but not for threadbars and D-bolts. Hence, the former displacement is external and the later is internal. In an integrated support system where the yielding support elements are combined with surface retaining elements such as mesh and shotcrete, a large deformation capacity at the plate would be preferred to accommodate large wall deformation and load transferred from the surface
60 100% Steel stretching (high strength steel)
100% Steel stretching to absorb energy (normal rebar steel) 5.6% steel stretching (94.4% cone plough)
50
Dynamic input energy (kJ)
Increasing steel strength/bar diameter
Increasing steel stretching
0% Steel stretching (100% cone plough) to absorb energy
40
Zone C 30
Zone B 20
Pure cone plough or friction resistence
MCB33 (new, 17.2 mm) MCB33 (old, 17.2 mm) Fully bonded threadbar (20 mm)
Zone A
Fully debonded threadbar (20 mm)
10
Toe anchored threadbar (20 mm) SA Conebolt (22 mm, 40 GPa grout) D-bolt (22 mm, 0.846 m segment)
0 0
100
200
300 400 500 600 Displacement (mm)
700
800
Fig. 11. Energy dissipation of MCB33 and other bolts versus deformation. ‘‘new’’ and ‘‘old’’ stand for plastic sleeve and grease debonded conebolts, respectively. (For interpretation of the references to colour in this figure, the reader is referred to the web version of this article.)
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retaining system. A better understanding of the classification of energy absorption mechanism of rockbolts shown in Fig. 11 is important not only for selecting the right bolts for dynamic rock support, but also for guiding new development of yielding support products in the future.
4. Conclusion The use of yielding support is a key component when designing a rockburst support system. Modified conebolt based rockburst support systems have been proven to be very effective to mitigate/limit rockburst damage. To facilitate the conebolt’s yielding or ploughing through the grout, little or no bonding between the grout and the bolt is required. The effect of various debonding agents on MCB conebolt performance is investigated in this study. It is found that grease debonded conebolts perform only marginally better than nongreased ones, indicating that grease is not an ideal debonding agent for MCB conebolts. A new debonding agent in the form of PVC sleeve was developed for MCB conebolts. The new design has greatly improved the dynamic capability of the conebolt. Compared with the test results of grease debonded conebolts, the new design allows the cone to work in the fashion it is intended for – to crush the resin and absorb the dynamic energy while maintaining the load capacity of the bolt. By eliminating the bond between the bolt and resin, the bolt performance is now more consistent with less variability. Reducing the rock support installation time is important for the underground construction industry. To achieve the best support effect, an integrated rock support system should include reinforcing element such as rebars and holding/yielding elements such as conebolts. The previous design of the MCB conebolt requires 38 mm boreholes for bolt installation. Because most mines use 33 mm diameter boreholes to install rebars, yielding support system installation thus requires a second pass if other borehole sizes are used, adversely impacting the construction productivity. MCB33 conebolts, suitable for 33 mm diameter boreholes, can be used in a one-pass rock support system to facilitate rapid drift development in underground mines.
Acknowledgements The authors wish to acknowledge Xstrata Nickel, Xstrata Copper, Kirkland Lake Gold, Vale, and Goldcorp for their support during the field test of the development of the new MCB33 conebolt. The authors also wish to thank CANMET for executing the dynamic drop test. Comments and suggestions by Peter Kaiser are highly appreciated, and permission to publish the dynamic
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test results from Mansour Mining Technologies Inc. is gratefully acknowledged.
References [1] Kaiser PK, Tannant DD, McCreath DR. Canadian Rockburst Support Handbook.Geomechanics Research Centre. Sudbury, Ont, Canada: Laurentian University; 1996. [2] Hedley DGF. Rockburst handbook for Ontario hardrock mines. CANMET Special report SP92-1E; 1992. [3] Ortlepp WD. The design of support for the containment of rockburst damage in tunnels an engineering approach. In: Kaiser PK, McCreath DR, editors. Rock Support in Mining and Underground Construction. Rotterdam: Balkema; 1992. p. 593–609. [4] Hoek E, Brown ET. Underground excavations in rock. London: Institution of Mining and Metallurgy; 1980. [5] Windsor CR, Thompson AG. Rock reinforcement—technology, testing, design and evaluation. In: Hudson JA, editor. Comprehensive Rock Engineering. Oxford: Pergamon; 1993. p. 451–84. [6] Rockfield Software Ltd. ELFEN. 2003; Version 3.7. [7] Jager A, Wojno LN, Henderson NB. New developments in the design and support of tunnels under high stress. In: Technical Challenges in Deep-level Mining, SAIMM Congress; 1990. vol 1, pp. 1155–71. [8] Cook NGW, Ortlepp WD. A yielding rockbolt. Chamber of Mines of South Africa Research Organization Bulletin; 1968. (14). [9] Jager AJ. Two new support units for the control of rockburst damage. In: Kaiser PK, McCreath DR, editors. Rock Support in Mining and Underground Construction; 1992. p. 621–31. [10] Ortlepp WD. Grouted rock-studs as rockburst support: a simple design approach and an effective test procedure. J South Afr Inst Min Metall 1994;94:47–53. [11] Player JR. Field performance of cone bolts at Big Bell mine. In: Villasescusa E, Potvin Y, editors. Ground Support in Mining and Underground Construction; 2004. p. 289–98. [12] Ortlepp WD. Dynamic capacity in cable anchors and rockbolts. In: Proceedings of the First International Symposium on Block and Sub-Level Cave Mining. Johannesburg: S Afr Inst Min Metall; 2007. pp. 405–20. [13] Tannant DD, Buss BW Yielding rockbolt anchors for high convergence or rockburst conditions. In: Proceedings of the 47th Canadian Geotechnical Conference; 1994. [14] Simser B, Andrieus P, Gaudreau D. Rockburst support at Noranda’s Brunswick Mine, Bathurst, New Brunswick. In: Hammah R, Bawden W, Curran J, Telesnicki M, editors. Proc NARMS 2002, vol. 1. University of Toronto Press; 2002. p. 805–13. [15] Simser B, Joughin WC, Ortlepp WD. The performance of Brunswick Mine’s rockburst support system during a severe seismic episode. J South Afr Inst Min Metall 2002:217–23. [16] Cai M, Champaigne D. The art of rock support in burst-prone ground. In: Tang CA, editor. RaSiM 7: Controlling Seismic Hazard and Sustainable Development of Deep Mines. Rinton Press; 2009. p. 33–46. [17] St-Pierre L, Hassani FP, Radziszewski PH, Ouellet J. Testing and analyzing the dynamic response of rock support elements. In: Proceedings of the Fourth International Symposium on High-Performance Mine Production; 2007. pp. 353–68. [18] Player JR, Thompson AG, Villasescusa E. Dynamic testing of threadbar used for rock reinforcement. In: Proceedings of the Third CANUS Rock Mechanics Symposium. Diederichs M, Grasselli G, editors; 2009. Paper 4030. [19] Li C, Charette F. Dynamic performance of the D-bolt. In: Proceedings of the Fifth International Seminar on Deep and High Stress Mining. Van Sint Jan M, Potvin Y, editors; 2010, pp. 321–8. [20] Plouffe M, Anderson T, Judge K. Dynamic testing of tendons (Roofex). CANMET-MMSL 07-053 (CR); 2008, p. 71.
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Technical Note
Influence of bolt-grout bonding on MCB conebolt performance M. Cai a,n, D. Champaigne b a b
School of Engineering, Laurentian University 933 Ramsey Lake RoadSudbury, Sudbury, Ont, Canada P3E 2C6 Mansour Mining Technologies Inc., Sudbury, Ont, Canada
a r t i c l e i n f o Article history: Received 27 November 2010 Received in revised form 25 June 2011 Accepted 11 November 2011 Available online 29 November 2011
1. Introduction 1.1. Rockburst and rockburst damage As the depth of mining and underground construction increases, stress-induced rock fracturing is inevitable and in some cases the rocks can fail violently, leading to seismic events and rockbursts. A rockburst is defined as damage to an excavation that occurs in a sudden or violent manner and is associated with a seismic event [1]. Many hard rock mines in Canada, China, Chile, South Africa, Australia, Sweden, and other countries, and some ¨ deep civil tunnels in Switzerland (e.g., Loetschberg, Gotthard, Furka Base Tunnels), China (e.g., Erlang Mountain Highway Tunnel, Jingping-II intake tunnels), and Peru (e.g., Olmos TransAndean Tunnel) have experienced rockbursts to various degrees. Rockbursts can be classified into three types [2]: (a) fault slip burst, (b) pillar burst, and (c) strain burst. When the mininginduced stresses exceed the strength capacity of the unsupported or supported rock, even temporarily, failure is initiated or triggered. Typical rockburst damage to underground excavations include stress-induced rock fracturing, bulking of roof and sidewalls, floor heave, shearing of rock, rock falls and ejections, etc. Rockbursts can cause fatalities and injuries to workers, damage to mine infrastructure and equipment, and substantially increase investment risk. For example, a magnitude mN ¼ 3.8 (Nuttli magnitude) seismic event, which happened on January 6, 2009, caused extensive damage to four levels at the Kidd Mine in Timmins, Ontario, Canada. One drift supported by standard rock support with mesh and rebar collapsed totally and a few pieces of mining equipment were damaged. Luckily, the event occurred at a time as the nightshift was going off duty and heading to surface and thus no injuries were reported. For obvious reasons, we must
n
Corresponding author. Tel.: þ1 705 675 1151x5099; fax: þ 1 705 675 4838. E-mail address: [email protected] (M. Cai).
1365-1609/$ - see front matter & 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijrmms.2011.11.006
rely on proper rockburst support to safeguard workers and investments. It is imperative to design and utilise proper rockburst-resistant support when mining or tunnelling at depth.
1.2. Rockburst support Rock support in burst-prone ground differs from conventional rock support in shallow ground where controlling gravityinduced rock falls is the main concern. Rock support in burstprone ground needs to resist dynamic loading and large rock dilation due to rock failure. Much research has been conducted (e.g., [1–3] and many others) to address the rockburst problem. These studies recommend and provide design of rockburst support as a means to mitigate rockburst damage. A rock support system helps the rock support itself [4,5]. This view of rock support function was derived from the observation of rock support under static loading. In general, the role of rock support in underground construction can be classified into three major functions—reinforce, retain, and hold [1]. The goal of reinforcing the rockmass using rockbolts and/or rebars is not only to strengthen it, thus enabling the rockmass to support itself but also to control the bulking process as rockbolts/rebars prevent fractures from propagating and opening up. Under high stress conditions, fractured rock between the reinforcing/holding elements may unravel if not properly retained. An important function of rock support is realized through retaining the fractured surface rock so as to confine rocks further away from the opening to prevent further fracturing. Widely used retaining elements include wire mesh, shotcrete, or cast-in concrete liners. The holding function is needed to tie retaining elements of the support system and loose rock back to stable ground, to dissipate dynamic energy due to rock ejection and rock movement, and to prevent gravity-driven falls of ground. Under dynamic loading condition, the rock support system must be able to absorb energy to survive the rockburst event. When rockburst damage is
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anticipated, yielding holding elements such as conebolts and friction bolts must be used in the support system. Otherwise, severe damage or complete failure of the excavation could result. 1.3. Need for yielding support in bursting ground 1.3.1. Rock dilation due to rock fracturing When the rock fractures, its volume increases significantly, and the phenomenon is often referred to as dilation or bulking. Fig. 1 shows the results of a numerical simulation using ELFEN [6], an FEM/DEM combined code that can properly simulate rock fracturing processes. Because a displacement loading boundary is applied, the rock failure process is stable and there is no violent rock ejection and disintegration. When the peak strength is reached, the rock starts to dilate and an 8% to 10% volume increase is observed in this case when the vertical strain reaches about 2%. In other words, when a 2 m thick ring of rock near the excavation boundary fails in this fashion with an 8% to 10% volume increase, the maximum wall displacement would be 160 to 200 mm, which is far beyond the deformation capacity of most standard ground support systems. Rock bulking resulting from disintegration of rock due to violent rock fracturing can lead to even greater volume increase. When rock fractures under low confinement conditions its volume increases significantly. Kaiser et al. [1] proposed the concept of bulking factor and used it to estimate maximum wall deformation. One typical example of rock bulking is floor heave in an underground excavation. In South Africa, complete drift closure due to floor heave has been encountered during severe rockbursts. In Canadian mines, bulking in the floor is commonly observed on the order of 20% to 40% of the fractured rock zone (if unconfined), typically leading to floor heave of 0.2 to 0.4 m. 1.3.2. Yielding design principle When mining at depth, the excavation-induced stresses around underground openings are so high that rock fracturing is inevitable. The stresses are further elevated when the rock is loaded dynamically. In many situations, it is no longer possible to increase the load carrying capacity of the reinforced rock system economically, and the support behaviour must be fundamentally changed to allow for yielding. Jager et al. [7] showed that the dilation pressure resulting from brittle failure in a deep tunnel sidewall exceeded 0.4 MPa. When the rock fracturing process is dynamic, it is likely that the pressure will be considerably higher, exceeding the capacity of most practical rock support systems.
Fig. 1. Stress-displacement relation of a rock sample with dilation behaviour.
For reference, the pressure (support pressure) to the excavation surface that can be supplied by a moderately dense pattern of rockbolts is about 0.1 MPa. To provide a support pressure of 0.4 MPa, very thick cast-in concrete would be required, which is clearly not economically practical for most mining applications. In general, rock load or pressure to the support system decreases with increasing rock deformation. If the rock support system is able to yield in a controlled fashion, it reaches a dynamic equilibrium during the deformation process and the system will eventually reach a new static equilibrium. Hence, the key point is that the rock support system must allow the rock mass to deform, which means that the rock support system must be deformable. A yielding support system should tolerate large tunnel convergence without ‘‘self-destruction’’ of the system while providing necessary support to ensure safety and maintain serviceability of the tunnel. A yielding rock support system is a system in harmony with the surrounding rock. Fortunately, rock bulking can be controlled effectively using appropriate rock support elements, but the rock support system used must be yieldable to accommodate large deformation. Field observation in a few mines confirmed that when yielding support is installed, the wall dilation can be effectively controlled (Fig. 2). Cook and Ortlepp [8] first suggested the use of yielding rockburst support in the deep gold mines in South Africa. Smooth bar yielding bolts were developed and tested by Ortlepp in 1969. Over the years, various rock support products have been developed and used in practice. Among them, conebolts stand out as good holding and energy absorbing elements in a rockburst support system. In the following, we present a brief review of the development of various types of conebolts, identify debonding as a most important issue for the effective performance of conebolts, describe a recent development to address the debonding problem, and compare the static and dynamic test results to reveal the influence of bolt-grout bonding on the behaviour of conebolts.
Fig. 2. Large wall bulking contained by yielding support.
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2. Historical development of conebolts 2.1. South African conebolts The concept and idea of conebolt was firstly developed in South Africa. The conebolts were groutable yielding tendons developed by the Chamber of Mines Research Organisation (COMRO) in 1987 [9] for use in cement grouted holes. When subjected to static loading, the cone functions as a wedge-style mechanical anchor similar to standard mechanical rockbolts. However, when subjected to dynamic loading, the cone yields or ploughs through the grout, thus absorbing the dynamic energy through controlled deformation. Ease of installation was one of the important considerations in the initial design of the South Africa conebolt. Because cement grouting of shepherd’s crook tendons was generally practiced in South Africa, this consideration led to the use of cement grout for the conebolts. To facilitate the bolt’s yielding or ploughing through the grout, little or no bonding between the grout and the bolt was permitted. Hence, debonding between the bolt and grout was required. Debonding was achieved by coating the bolt with wax. A special saponified wax was developed and patented. The wax, which adheres well to steel, has an appropriate melting point and viscosity for ease of applying a thin layer of coating to the bar. One issue related to the use of wax was that its stability was affected in the presence of ammonia. If the wax coated bolts were stored for long periods of time underground in areas contaminated by blasting fumes, the wax could potentially lose its effectiveness. Ortlepp [10] conducted a quarry site blasting test and demonstrated the effectiveness of South African conebolts in absorbing kinetic energy. Cement grouted conebolts have been used in a few Australia mines such as Big Bell Mine [11]. However, conebolts were reluctantly accepted in South African mines, even in deeplevel rockburst prone mines. This was largely due to the high initial cost [12]. South African conebolts were first tested in Canada in 1994 [13] using cement grout and polyester resin (injected with quicksetting resin grout). The cement grouted conebolt had a maximum displacement of 240 mm and the load-displacement curve resembles that of a steel stretching curve, indicating that there was little cone movement during pull-out test. In resin grout testing, the conebolt could not shear through the resin and failed after only about 100 mm of displacement. To our knowledge, no
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mine in Canada systematically used the South African conebolts in its operation. The main reason being that resin injection was not practical for use at the time and cement grouting was time consuming and required a third pass to torque the bolts once the grout was cured. 2.2. Modified conebolts (MCB) The increase in violent ejection failures in 1999/2000 at Brunswick Mine in Canada led to the urgent development of a complete yielding support system [14]. The South African conebolts were modified to accommodate resin-grouting applications in Canada. Norand Inc. modified the conebolt head and added a blade for the purpose of resin mixing and the new bolt was called Modified Cone Bolt (MCB). This new tendon can shear through the surrounding column of polyester resin grout. It is a smooth bar threaded at its outer end, with a forged cone and mixing blade at the other end. Grease was used as the debonding agent. Initial dynamic testing was conducted using a one tonne dynamic drop facility designed by MIRARCO (Laurentian University, Canada) for the Noranda Technology Centre (NTC) (a modified version of the NTC drop test facility is now known as the CANMET drop test facility shown in Fig. 7). At Brunswick Mine, MCB38 conebolts (grease-debonded, installed in 38 mm boreholes) at 1 m spacing with 150 150 mm2, 9.5 mm thick domed plates, were used with #0 gauge mesh straps (8.23 mm wire diameter, 10 10 cm2 aperture, typical sheet size: 30.5 365.8 cm2) to overlay the diamond mesh (5 cm aperture, 5.0 mm diameter), installed on a second-pass system [14,15]. A close-up view of the rockburst support system installed at Brunswick Mine can be seen in Fig. 3. There was a distinct boundary in one drift (tunnel) that defined stable and unstable grounds in Fig. 3a. This boundary was called the ‘‘extreme edge’’ at the mine site. Beyond the extreme edge, there were only standard rock support systems installed. A series of rockburst events that occurred between October 13 and 17, 2000 severely damaged the ground where standard supports were installed, but the section that was supported by the rockburst support system suffered no damage. On Oct. 13, 2000, a Nuttli magnitude mN ¼2.5 rockburst event caused the collapse of an intersection involving approximately 1860 t of material. A small portion of the 326 crosscut immediately east of the intersection had MCB conebolts installed (a close-up view of the 326 crosscut can be seen in Fig. 3b). The right-hand side of the tunnel was supported by 2.3 m
#6 Chain-link mesh+ #0 mesh strap + MCB conebolts(17.2 mm)
Chain-link mesh
Thread-bar +
#6 Chain-link mesh
shotcrete + #9
+ #0 mesh strap +
gauge weld mesh
MCB conebolts (17.2mm)
+ end anchored rockbolts
Fig. 3. ‘‘Ground truthing’’ of modified conebolts at Brunswick Mine (photo courtesy: Brunswick Mine and B. Simser).
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long MCB conebolts in a 1 1 m2 pattern with #0 gauge mesh straps and chain-link mesh. The left-hand side of the tunnel was supported only by standard support (rebar þshotcrete þ#9 gauge weld mesh). The rockburst support system installation had not been fully completed at the time of the bursts as the chain-link mesh was temporarily suspended with 0.9 m end anchored rockbolts and the next step of installing MCBs and mesh straps had not been completed. The foreground of the area was supported by a rebar and steel fibre reinforced shotcrete. After the rockburst event, it was revealed that the right-hand side of the tunnel suffered no visible damage at all while the rest of the area suffered moderate to severe damage. The left-hand side showed excessive large bulking. While this was not intended as a ‘‘field experiment,’’ the observed rockburst damage to different sections clearly showed that a well-designed rockburst support system can effectively limit or eliminate rockburst damage to underground openings. The excellent performance of the MCB supported tunnel section held the ground so well that many people speculated that the rock might not be hit hard as the rest of the area. However, close onsite examination revealed that the MCBs had fulfilled their role in dissipating dynamic energy as some conebolts displaced as much as 180 mm [14]. Subsequent use of MCBs based rock support system at the mine site performed exceptionally well. Following the success at Brunswick Mine, MCBs have enjoyed a slow but gradual acceptance by some Canadian mines having rockburst problems, and have been successfully applied in their operations.
3. Influence of bolt-grout bonding on conebolt performance 3.1. The debonding issue The South African (SA) conebolts have a 0.1 mm thick saponified wax coating as the debonding agent. Approximately 10 months after receiving the SA bolts at Big Bell Mine in Australia,
it was noted that some bolts exhibited a substantial deterioration or flaking of the protective wax coating [11]. The original MCB was greased to reduce the bonding effect between the grout and the bolt. However, it was noted that the spinning action of the bolt to mix the resin stripped off almost all the grease coating, leading to unwanted bond between the bar and the resin. If wax or grease is not working as intended as a debonding agent, then bonding between the bolt and the grout can reduce the effectiveness of conebolts. The question is how much bonding exists between the bolt and the grout? This can be found out by static and dynamic tests in the laboratory or in the field. Eight South African conebolts (22 mm diameter), four with intact wax coating and four with the wax coating physically removed, were tested at Big Bell Mine to compare the bonding behaviour of bolt and the cement grout (sc E35 MPa), with their cones cut off [11]. The bolt lengths were between 1.0 and 1.5 m. It was seen that for bolts with wax coating physically removed, the peak loads vary between 42 and 92 kN and the residual loads vary between 35 and 70 kN, at a displacement of 150 to 200 mm. For the bolts with wax coating, the peak loads vary between 40 and 91 kN and the residual loads vary between 17 and 25 kN, at a displacement of 150 to 200 mm (see Fig. 4). The conclusion was that the wax coated smooth bar had on average 25% and 50% less peak and residual bond strength per metre of embedment than bolts with wax removed, respectively. In our opinion, these test results suggest that wax coatings are not effective in debonding the conebolts from the cement grout. Cai and Champaigne [16] conducted a series of similar pull tests using steel pipes. Greased MCBs (17.2 mm diameter, 1.2 m long) were installed in the steel pipes (inner diameter 33 mm, wall thickness 5.2 mm) using resin. The bolts were fully encapsulated. After the resin was cured, the cone section (about 0.1 m) was cut off and the remaining bolt was pull tested until the bolt started to slide in the resin. For comparison purposes, the greased bar test results are presented in Fig. 4 with the results of Player [11]. In our test, the displacement range was limited to about 35 mm, due to stroke limitation of the ram. It was found that for an embedment length of 0.94 m, the maximum load recorded was 127 kN. This load
Fig. 4. Load-displacement curves of smooth bars with different debonding agents.
M. Cai, D. Champaigne / International Journal of Rock Mechanics & Mining Sciences 49 (2012) 165–175
is in fact close to the yield point of the steel bar. Another bar tested with a grease coating had a peak load of 81.6 kN. This is also very high, considering the short embedment length of 0.94 m. Similarly to the South African wax coated conebolts, the pull test results showed that the grease coating is not effective in debonding the MCBs from the resin grout. It can then be concluded that the grease coating was stripped off of the MCBs during the installation process of spinning the bolt through the resin at the manufactures recommended 30 to 40 revolutions.
3.2. The next generation conebolt—MCB33 It is obvious that an effective debonding agent is required in order to improve the static and dynamic performances of MCBs. No better wax or different grease debonding agent could be identified or developed. As a consequence, a new debonding agent in the form of a plastic (PVC) sleeve installed over the shaft of the MCB conebolt was developed [16]. The lab test results show that this debonding agent can eliminate the bonding force effectively. As shown in Fig. 4, the load for a 0.94 m section of smooth bar to slide in resin grout is only 4.5 to 6.8 kN if a bar is
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fully covered by a plastic sleeve. In one test, about 20 mm of the bar was not covered by the sleeve and it resulted in a peak load of 22.7 kN. The residual load of this bar reached to 9 kN at a displacement of 35 mm. The effectiveness of the plastic sleeve as a debonding agent was confirmed. The new design of the MCB33 conebolt is shown in Fig. 5. The bar diameter is 17.2 mm. The bolts are manufactured from C1055 modified steel with a typical yield strength of 517 MPa, a typical ultimate tensile strength of 745 MPa, and a typical elongation capacity of 16%. The new debonding agent, heat shrink plastic sleeve (20 mm diameter), is applied to the whole bar shaft except the cone section and thread section. The sleeve is durable and sleeve damage is unlikely if the bar is handled without abuse. The sleeve is tightly around the bar shaft and there is no sleeve damage or movement during bolt installation. There are currently two different cone sizes in use—MCB38 for 38 mm boreholes and MCB33 for 33 mm boreholes. Because regularly used rebars and rockbolts in underground mines all use 33 mm boreholes, it makes much more economical sense to use the MCB33 conebolts, i.e., one bit and one resin. This eliminates extra bits and resin sizes from inventory, and greatly simplifies rock support installation. In addition, it is possible to
Fig. 5. Fully debonded MCB33 conebolt (patented).
180 Rebar
160
MCB33 (grease debonded) S.A conebolt
140
(GRC test result)
Load (kN)
120 100 80 MCB33 (PVC debonded)
60 40 20 0 0
20
40
60
80 100 Displacement (mm)
120
140
160
Fig. 6. Load-displacement relations of MCB33 conebolts with different debonding agents obtained from in-situ pull tests.
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install all standard and rockburst support systems in one pass, which is desired for rapid drift development. The final resin matrix quality is important to ensure that the dynamic response of the MCB consistently performs on a repetitive basis and as designed. A number of factors can affect the mixing performance of resin cartridges: insertion rate, rotational speed, bolt to hole annulus, resin viscosity, etc. In the case of MCB33, the borehole diameter was reduced from 38 mm down to 33 mm, therefore increasing resin mix-ability considerably. In the following discussion, we compare in-situ pull test and dynamic drop test results of MCB33 conebolts with different debonding agents. The goal is to understand how the conebolt performance can be affected by the method of debonding. 3.3. Comparison of static behaviour with different debonding agents Extensive in-situ pull tests were performed at five mines in Canada to evaluate the load-displacement response of MCB33 conebolts with grease and PVC sleeve as the debonding agent. The MCB33 conebolt total length is 1.7 m and the resin embedment length is roughly 1 m. A few test results are presented in Fig. 6. For comparison, load-displacement curves of a 20 mm rebar and a 16 mm South African conebolt are also shown in the same figure. The cement grouted conebolt (tested by GRC (Geomechanics Research Centre, Laurentian University, Canada) [1]) shows similar stiff behaviour as rebar until the load reaches 125 kN. After the peak load, there is an initial drop in load followed by an ever increasing load. The response of grease debonded MCB33 conebolts show a linear trend before the load reaches about 100 kN. When the load is further increased, the bolt debonds from the resin and the displacement rate increases drastically. Compared with the response of rebar and the South African conebolt, grease debonded conebolts are less stiff but the displacement that correspond to bond breakage (load deviation from linearity) is small, generally less than 5 mm. The reason that the grease debonded MCB33 conebolts are less stiff than the South African conebolt is that the bolts are not fully encapsulated with resin in the boreholes. For fully encapsulated bolts, the initial stiffness will be similar to that of rebar. The response of PVC sleeve debonded MCB33 conebolts is characterised with a gradual increase of load with deformation and a plateau is reached when the displacement is in the range of 20 to 40 mm. After the plateau, the load continues to increase. Yielding of the steel was noticed from the reduction of bolt diameter. The tests were stopped because the displacement reached the stroke limit of the ram. Compared with the behaviour of grease debonded MCB33, PVC sleeve debonded conebolts are less stiff before the plateau is reached. The ultimate loads of the bolts are not affected by the type of debonding agents used, but the ultimate deformations of PVC debonded conebolts are slightly larger (most likely due to some cone movements). The behaviour of the MCB33 conebolts with PVC debonding agent is attributed mostly to the cone resistance in resin without the interference of frictional forces between the bolt and resin. This feature may not be very desirable under slow loading rates, however under high loading conditions, as will be the case under seismic load, complete debonding of the shaft of the bar from the grout can enhance the cone plough performance and is demonstrated in the dynamic drop tests described next.
rockbursts. However, this difficulty does not stop researchers from finding practical means to test the dynamic behaviour of rock support elements. Drop test facilities, which can perform repeatable dynamic loading, have been built in Canada, South Africa, Sweden, and Australia to test the dynamic properties of reinforcing and holding elements. The drop test is characterised by its ease to repeat tests at a relatively low cost. The drop test facility employed in this study is shown in Fig. 7. The facility is used for rockbolt testing in steel tubes. The drop weight is lifted with an electromagnet, which in turn is lifted by a pair of cranes mounted in parallel on the top of the machine. By cutting the power to the magnet, the weight free falls onto the sample. The maximum weight capacity is 3 t, and the maximum input energy is 58.9 kJ at a drop height of 2 m. A target for measuring the plate displacement is installed at the end of the bolt, on the remaining threaded section. Displacements are measured using line-scan cameras. A rod is connected to the upper end of the bolt test specimen and a second target is fixed at the top of the rod to measure the cone displacement. To ensure that the bolts are properly installed into the steel tubes, CANMET also developed a special installation rig, which can control the insertion time and the number of bolt revolutions to minimise the difference in resin mixing quality between samples.
3.4.2. Dynamic behaviour of non-greased and grease debonded MCB conebolts One of the key factors, which influence conebolt performance is the type of debonding agent used. Another important factor that affects conebolt performance is the mixing of the resin during bolt installations. The normal installation procedure for MCB conebolt is to spin the bolt to mix the resin as soon as the bolt is inserted into the tube (or borehole in the field). Spinning is continued at a constant insertion rate until the bolt reaches the end of the tube (borehole). Immediately the bolt is spun another 30 to 40 revolutions to further mix the resin and the installation is completed. We also investigated an alternative installation method, i.e., to push the bolt into the tube as much as possible before spinning the resin for 30 to 40 revolutions. Sample autopsy by splitting the tube to examine the mixed resin indicated that
Steel tube
Drop weight
Plate
Test bolt
3.4. Comparison of dynamic behaviour with different debonding agents 3.4.1. Dynamic drop test method and facility It is almost impossible to have a testing system that can simulate the real dynamic loading to a bolt resulting from
Fig. 7. CANMET drop test facility and test setup.
M. Cai, D. Champaigne / International Journal of Rock Mechanics & Mining Sciences 49 (2012) 165–175
both the normal and the alternative installation methods can mix the resin satisfactorily. First, we conducted six drop tests of MCB33s, of which three were greased and three non-greased. The bolts were installed in the steel tubes using the alternative installation method. A mass of 1115 kg and a drop height of 1.5 m were used in this test series, and each impact produced 16 kJ of input energy. The test results are presented in Fig. 8, showing the plate displacement (in mm) and the relative cone displacement and steel stretching (in %) to the plate displacement. In the first impact loading (Drop-1), the cone plough, expressed as a percentage of the cone movement to the total deformation measured at the plate, varied from 11.2% to 97.5% (with an average of 57%) for non-greased bolts, showing a large variability. For greased debonded bolts, the cone plough varied from 58.1% to 95.4% with an average of 74.5%. In the second impact loading (Drop-2), for non-greased bolts, the cone plough varied from 1.1% to 91.1% with an average of 31.1%.
For greased debonded bolts, the cone plough varied from 43.1% to 72.0% with an average of 61.0%. One non-greased bolt and all three greased debonded bolts were dynamically loaded the third time (Drop-3). Again, the test results showed large variability in terms of cone plough behaviour. For one greased bolt, the cone was ‘‘locked’’ in place and 100% of the plate displacement came from steel stretching. As can be seen from the results, with respect to cone movement, grease debonded bolts seem to perform only marginally better on average (65% of total displacement is due to cone movement in all drops) than non greased bolts where cone movement accounts for 48% of total displacement. This test result confirms that grease is not an ideal debonding agent for MCB conebolts. In a separate dynamic drop tests conducted at CANMET, St-Pierre et al. [17] also found that the influence of grease on the bolt behaviour was not obvious. In two of their tests, the cones were cut off and a drop weight of 1016 kg from 0.5 m (5 kJ input energy) was used to test the dynamic capability of the
Non greased Drop-1
171
Greased
100%
16 kJ
80% 60% 40% 20%
0% Plate displ.(mm)
116
101
161
283
126
129
Steel stretch %
88.8%
37.6%
2.5%
4.6%
30.2%
41.9%
Cone plough %
11.2%
62.4%
97.5%
95.4%
69.8%
58.1%
Non greased Drop-2
Greased
100%
16 kJ
80% 60%
32 kJ (accumulative)
40% 20%
0% Plate displ.(mm)
79
89
93
130
122
150
Steel stretch %
8.9%
98.9%
98.9%
56.9%
32.0%
28.0%
Cone plough %
91.1%
1.1%
1.1%
43.1%
68.0%
72.0%
Non greased Drop-3 16 kJ
Greased
100% 80% 60%
48 kJ (accumulative) 40% 20% 0% Plate displ.(mm)
102
137
92
120
Steel stretch %
28.4%
46.7%
100.0%
37.5%
Cone plough %
71.6%
53.3%
0.0%
62.5%
Fig. 8. Comparison of dynamic behaviours of non-greased and grease debonded MCB33 conebolts, showing the percentage of cone plough and steel stretching at 16 kJ impact energy.
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energies of 16 kJ and the results are compared with the results of previous tests using grease as the debonding agent. All the tested bolts were the same type (MCB33), installed in the same steel tube using the normal installation procedure with the same type of resin. Dynamic drop test results showed that on average 99.6% of the plate displacement of PVC sleeve debonded bolts was from cone plough on the first drop as compared to 82.9% in the old design using grease as the debonding agent (Fig. 9). In the second drop, the cone plough percentages for the new and old designs are 100.0% and 29.2%, respectively. In the third drop, the cone plough percentages for the new and old designs are 99.6% and 4.7%,
smooth bars without cones. The resin embedment length was estimated at 1.2 m. The bar, which was greased sustained six drops with a total input energy of 35 kJ and the non-greased bolt accepted five drops with a total input energy of 30 kJ. Note that 5 kJ impact energy is large enough to break a fully bonded rebar at the threads.
3.4.3. Dynamic behaviour of PVC sleeve debonded MCB33 conebolts Dynamic drop tests were conducted to validate the new design of debonding agent using PVC sleeve. Firstly, plastic sleeve debonded conebolts (17.2 mm diameter) were tested at impact
PVC sleeve Drop-1
Greased
100%
16 kJ
80% 60% 40% 20% 0%
Plate displ.(mm)
334
385
484
201
127
153
150
161
101
Steel stretch %
0.3%
0.0%
0.0%
4.5%
22.0%
9.8%
9.3%
2.5%
54.5%
Cone plough %
99.7%
100.0%
100.0%
95.5%
78.0%
90.2%
90.7%
97.5%
45.5%
PVC sleeve Drop-2
Greased
100%
16 kJ
80% 60%
32 kJ (accumulative)
40% 20% 0%
Plate displ.(mm)
297
264
386
172
98
85
90
93
Steel stretch %
0.0%
0.0%
0.0%
0.0%
58.2%
100.0%
96.7%
98.9%
Cone plough %
100.0%
100.0%
100.0%
100.0%
41.8%
0.0%
3.3%
1.1%
PVC sleeve Drop-3
Greased
100%
16 kJ
80% 60%
48 kJ (accumulative)
40% 20% 0%
Plate displ.(mm)
350
359
85
Steel stretch %
0.6%
0.3%
95.3%
Cone plough %
99.4%
99.7%
4.7%
Fig. 9. Comparison of dynamic behaviours of PVC sleeve debonded and grease debonded MCB33 conebolts, showing the percentage of cone plough and steel stretching at 16 kJ impact energy.
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repeated shocks is significantly reduced. For example, a greased bolt resulted in 54.5% of steel stretching in the first drop. With this degree of steel stretching, the bolt did not survive the second drop. The new design of the conebolts eliminates this problem completely. In fact, two new bolts received a third energy input of 16 kJ and the cone ploughs were 99.4% and 99.7%, respectively. The test results of grease-debonded conebolts shown in Figs. 8 and 9 indicate that the installation method may have a minor influence on the effectiveness of grease as a debonding agent. Compared with the results using the alternative installation method shown in Fig. 8, the normal installation method leads to more steel stretching in the second drop as shown in Fig. 9. This is because the normal installation procedure causes more spinning of the bolt, thus has a larger potential of stripping more grease off the bolt during installation. When the debonding is not effective, consecutive loading can stiffen the bolt and the bolt relies on more steel stretching to absorb the impact energy. Three new MCB33 conebolts with plastic sleeves were also tested at 26 kJ impact energy (1784 kg weight, 1.5 m drop height). These tests resulted in 100% and 99.9% cone ploughs in the first and second drops, respectively (Fig. 10). This means that almost all of the impact energy was absorbed by cone plough. In our previous drop tests of grease debonded MCB conebolts, the highest impact energy tested was 22 kJ, which resulted in significant steel stretch. Close to perfect cone plough at 26 kJ was very encouraging and it was decided to use the last available bolt to test at an even higher energy level. The bolt was subjected to a 33 kJ impact load from a single drop (2229 kg weight, 1.5 m drop height) and the bolt survived. Cone plough constituted 76.3% of the total
respectively. It is seen that the new design allows the cone to plough through the resin very effectively while the old design relied increasingly on steel stretch to absorb the impact energy of repeated dynamic loading. It is of interest to explore why greased MCBs performed worse on the second drop. On the second drop one can clearly see more steel stretch and less cone plough in Fig. 9b. To explain this phenomenon, we need to refer to Fig. 4. As explained before, grease could be stripped off during the bolt installation, allowing the shaft of a MCB to come directly in contact with the resin. The smooth bar behaves more or less like the curves of ‘‘No wax’’ shown in Fig. 4. Although the curve was obtained in cement grout, we speculate that the response of a clean smooth bar in resin grout will be similar. Load resistance increases as the bars slide in the grout. The first drop test of greased MCB conebolts resulted in 149 mm plate displacement (average value). On the second drop, the frictional resistance along the bolt shaft is higher than the resistance in the first drop. With this effect, coupled with a change in cone resistance, the cone plough action became less effective in the second and subsequent drops. As mentioned above, greased MCBs could withstand 16 kJ of impact energy 3 to 4 times before the bolt broke. This means that the bolts become stiffer after repeated dynamic loading and the cause of this phenomenon can be attributed to the in-effectiveness of grease as a debonding agent. The implication of this observed phenomenon is that greased MCBs can become ‘‘stiff’’ after an initial moderate dynamic loading (e.g., 16 kJ). The same conclusion can be reached by examining the non-greased bolt test data (Fig. 8). With less cone plough and more steel stretching, their ability to withstand
26 kJ
33 kJ
100%
Drop-1
80% 60%
40% 20% 0% Plate displ. (mm)
602
597
729
531
Steel stretch %
0.0%
0.0%
0.0%
23.7%
Cone plough %
100.0%
100.0%
100.0%
76.3%
26 kJ Drop-2
100% 80% 60%
40% 20% 0% Plate displ. (mm)
173
393
636
Steel stretch %
0.3%
0.0%
Cone plough %
99.7%
100.0%
Fig. 10. Percentage of cone plough and steel stretching of MCB33 conebolts at 26 kJ and 33 kJ impact energy.
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displacement and steel stretch was 23.7%, which represents only 5.6% steel strain. This bolt did not survive the second drop at the same energy input level (33 kJ). It should be noted that the third bolt tested at 26 kJ did not fail in the second drop. The plate hit the floor so that displacement measurement could not be correctly conducted. Fig. 11 presents the relationship between bolt displacement and absorbed impact energy. Data other than MCBs are obtained from Player et al. [18] and Li and Charette [19]. It is seen that toe anchored thread bars fail at the thread and absorb very little energy ( o2 kJ). The fully bonded threadbars are tested using load transfer by separating the bolt in the middle of the installed pipe. The debonded threadbars have about 1.6 m debonded section, which allow the steel to stretch. All data points are from the first drop loading. Four threadbars failed but there were no conebolt failures on the first drop. Very good linear correlations between steel strain and energy absorption can be seen from the threadbar data and the D-bolt data. It is observed that the cement-grouted South Africa conebolts and at least two of the old MCB conebolts follow the trend line of threadbar, indicating that the dominant energy absorption mechanism of those conebolts had been steel stretch with very little cone plough. On the other hand, the new MCB design allows the cone to plough much more efficiently, and their capacity to absorb dynamic energy is thus substantially increased. Using MCB conebolt and threadbar data as an example, we identified in Fig. 11 three zones showing three distinct mechanisms of energy absorption by rockbolts. In Zone A, pure cone plough or friction resistance is responsible for energy dissipation. The energy dissipation capacity of most PVC sleeve debonded
conebolts fall into Zone A. The energy dissipation capacity of friction bolts such as Split Set will fall into Zone A as well, but at a much lower energy dissipation rate (indicated by the slope of the line). In Zone B, the energy is dissipated through a combination of steel stretching and cone plough (or other yielding mechanism). Most grease debonded MCB data points fall into this zone. In Zone C, including the boundary indicated by the blue solid line, energy absorption is achieved by pure steel stretching. The energy dissipation rate is hence dependent on the yielding mechanism. When steel stretching is involved in energy dissipation, high energy dissipation rates can be achieved. The highest energy dissipation rates can only be realized by deploying high strength steel (e.g., D-bolt) and/or using larger diameter bars, at an additional material cost. For other yielding bolts such as Roofex [20], their behaviour can be classified into Zone A or Zone B accordingly (the slope of the zones may be different than the ones shown in Fig. 11 because MCB data are used here). It is critical to differentiate the deformation source of the rockbolts shown in Fig. 11. For MCB conebolts, the deformation comes from the plate displacement, reflecting all the deformation due to cone sliding and steel stretching. For threadbar and D-bolt, the deformation comes from the steel stretching from a free deformation section between two internal anchor points/sections. Dynamic loads were directly applied to the plates for MCBs but not for threadbars and D-bolts. Hence, the former displacement is external and the later is internal. In an integrated support system where the yielding support elements are combined with surface retaining elements such as mesh and shotcrete, a large deformation capacity at the plate would be preferred to accommodate large wall deformation and load transferred from the surface
60 100% Steel stretching (high strength steel)
100% Steel stretching to absorb energy (normal rebar steel) 5.6% steel stretching (94.4% cone plough)
50
Dynamic input energy (kJ)
Increasing steel strength/bar diameter
Increasing steel stretching
0% Steel stretching (100% cone plough) to absorb energy
40
Zone C 30
Zone B 20
Pure cone plough or friction resistence
MCB33 (new, 17.2 mm) MCB33 (old, 17.2 mm) Fully bonded threadbar (20 mm)
Zone A
Fully debonded threadbar (20 mm)
10
Toe anchored threadbar (20 mm) SA Conebolt (22 mm, 40 GPa grout) D-bolt (22 mm, 0.846 m segment)
0 0
100
200
300 400 500 600 Displacement (mm)
700
800
Fig. 11. Energy dissipation of MCB33 and other bolts versus deformation. ‘‘new’’ and ‘‘old’’ stand for plastic sleeve and grease debonded conebolts, respectively. (For interpretation of the references to colour in this figure, the reader is referred to the web version of this article.)
M. Cai, D. Champaigne / International Journal of Rock Mechanics & Mining Sciences 49 (2012) 165–175
retaining system. A better understanding of the classification of energy absorption mechanism of rockbolts shown in Fig. 11 is important not only for selecting the right bolts for dynamic rock support, but also for guiding new development of yielding support products in the future.
4. Conclusion The use of yielding support is a key component when designing a rockburst support system. Modified conebolt based rockburst support systems have been proven to be very effective to mitigate/limit rockburst damage. To facilitate the conebolt’s yielding or ploughing through the grout, little or no bonding between the grout and the bolt is required. The effect of various debonding agents on MCB conebolt performance is investigated in this study. It is found that grease debonded conebolts perform only marginally better than nongreased ones, indicating that grease is not an ideal debonding agent for MCB conebolts. A new debonding agent in the form of PVC sleeve was developed for MCB conebolts. The new design has greatly improved the dynamic capability of the conebolt. Compared with the test results of grease debonded conebolts, the new design allows the cone to work in the fashion it is intended for – to crush the resin and absorb the dynamic energy while maintaining the load capacity of the bolt. By eliminating the bond between the bolt and resin, the bolt performance is now more consistent with less variability. Reducing the rock support installation time is important for the underground construction industry. To achieve the best support effect, an integrated rock support system should include reinforcing element such as rebars and holding/yielding elements such as conebolts. The previous design of the MCB conebolt requires 38 mm boreholes for bolt installation. Because most mines use 33 mm diameter boreholes to install rebars, yielding support system installation thus requires a second pass if other borehole sizes are used, adversely impacting the construction productivity. MCB33 conebolts, suitable for 33 mm diameter boreholes, can be used in a one-pass rock support system to facilitate rapid drift development in underground mines.
Acknowledgements The authors wish to acknowledge Xstrata Nickel, Xstrata Copper, Kirkland Lake Gold, Vale, and Goldcorp for their support during the field test of the development of the new MCB33 conebolt. The authors also wish to thank CANMET for executing the dynamic drop test. Comments and suggestions by Peter Kaiser are highly appreciated, and permission to publish the dynamic
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test results from Mansour Mining Technologies Inc. is gratefully acknowledged.
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