錢營孜煤礦1.8 Mta新井設(shè)計含5張CAD圖.zip,錢營孜煤礦1.8,Mta新井設(shè)計含5張CAD圖,錢營孜,煤礦,1.8,Mta,設(shè)計,CAD 英文原文 Geotechnical considerations in mine backfilling in Australia N. Sivakugan a,*,R.M. Rankine b, K.J. Rankine a, K.S. Rankine a a School of Engineering, James Cook University, Townsville 4811, Australia b Cannington Mine, BHP Billiton, P.O. Box 5874, Townsville 4810, Australia Abstract :Mine backfilling can play a significant role in the overall operation of a mine operation. In the Australian mining industry, where safety is a prime consideration, hydraulic systems are the most common backfills deployed.Many accidents reported at hydraulic fill mines worldwide have mainly been attributed to a lack of understanding of their behaviour and barricade bricks.This paper describes the findings from an extensive laboratory test programme carried out in Australia on more than 20 different hydraulic fills and several barricade bricks. A limited description of paste backfills is also provided, and the usefulness of numerical modelling as an investigative tool is highlighted. Keywords: Hydraulic fills; Mining; Backfills; Paste fills; Geotechnical 1．Introduction In the mining industry, when underground ore bodies are extracted, very large voids are created, which must be backfilled. The backfilling strategies deployed often make use of the waste rock or tailings that are considered by-products of the mining operation. This is an effective means of tailing disposal because it negates the need for constructing large tailing dams at the surface. The backfilling of underground voids also improves local and regional stability, enabling safer and more efficient mining of the surrounding areas. The need for backfilling is a major issue in Australia, where 10 million cubic metres of underground voids are generated annually as a result of mining [1]. There are two basic types of backfilling strategies. The first, uncemented backfilling, does not make use of binding agents such as cement, and their characteristics can be studied using soil mechanics theories. A typical example of uncemented backfilling is the use of hydraulic fills that are placed in the form of slurry into the underground voids. The second category, cemented backfilling, makes use of a small percentage of binder such as Portland cement or a blend of Portland cement with another pozzolan such as fly ash, gypsum or blast furnace slag. The purpose of this paper is to analyse the findings from an extensive laboratory test programme carried out in Australia on hydraulic fills and several barricade bricks. Hydraulic fills are uncemented techniques, and are one of the most widely used backfilling strategies in Australia. More than 20 different hydraulic fills, representing a wide range of mines in Australia, were studied at James Cook University (JCU). The grain sizer distributions for all of these fills lie within a narrow band as shown in Fig. 1. Along with them, the grain size distribution curves for a paste fill and a cemented hydraulic fill are also shown. It can be seen that the cemented hydraulic fill falls within the same band as the hydraulic fill. The addition of a very small percentage of cement has a limited effect on grain size distribution. Paste fills generally have a much larger fine fraction than hydraulic fills or cemented hydraulic fills, but have negligible colloidal fraction finer than 2 μm. Fig. 1. Typical grain size distribution curves for hydraulic fills,cemented hydraulic fills and paste fills. 2．Hydraulic backfills Hydraulic fills are simply silty sands or sandy silts without clay fraction, and are classified as ML or SM under the Unified Soil Classification System. The clay fraction is removed through a process known as desliming, whereby the entire fill material is circulated through hydrocyclones and the fine fraction is removed and then sent to the tailings dam. The remaining hydraulic fill fraction is reticulated in the form of slurry through pipelines to underground voids. Over the past decade there has been a steady increase in the solid content of the hydraulic fill slurry placed in mines in an attempt to reduce the quantity of water that must be drained and increase the proportion of solids. The challenge posed by a high solid content is that it becomes difficult to transport the slurry through the pipelines due to rheological considerations. Currently, solid contents of 75---80% are common, although even at 75% solid content, assuming a specific gravity of 3.00 for the solid grains, 50% of slurry volume is water. Therefore, there is opportunity for a substantial amount of water to be drained from the hydraulic fill stope. To contain the fill, the horizontal access drives created during mining are generally blocked by barricades constructed from specially made porous bricks (Fig. 2). The access drives, which are made large enough to permit the entry of machinery during mining, are blocked by the barricades during filling. The drives are often located at more than one level. Initially, the drives located at upper levels act as exit points for the decanted water, and also serve as drains when the hydraulic fill rises in the stope. Fig. 2. An idealised stope with two sublevel drains. 2.1 Drainage considerations Drainage is the most important issue that must be considered when designing hydraulic fill stopes. There have been several accidents (namely, trapped miners and machinery) worldwide caused by wet hydraulic fill rushing through horizontal access drives. Several reasons, including poor quality barricade bricks, liquefaction, and piping within the hydraulic fill are attributed to such failures [2]. Therefore, permeability of the hydraulic fill in the stope is a critical parameter in the design; continuous effort is made during mining to ensure that it is kept above a threshold limit in the vicinity of 100 mm/h [3]. Larger permeability leads to quicker removal of water from the stope, thus improving the stability of the fill contained within the stope. Permeability tests for mine fills and barricade bricks are discussed by Rankine et al. [4]. The constant head and falling head permeability tests carried out on the hydraulic fill samples give permeability values in the range of 7---35 mm/h. In spite of having permeability values much less than the 100 mm threshold suggested by Herget and De Korompay [3], each of these hydraulic fills has performed satisfactorily. Anecdotal evidences and back calculations using the measured flow in the mine stopes suggest that the permeability of the hydraulic fill in the mine is often larger than what is measured in the laboratory under controlled conditions. Kuganathan[5] and Brady and Brown[6] proposed permeability values in the range of 30---50 mm/h, which are significantly larger than those measured in the laboratory for similar fills. These values are much less than the threshold limit prescribed by Herget and De Korompay[3], suggesting that it is a conservative recommendation. 2.2 Stability considerations The stability of the hydraulic fill stope during and after the drainage period depends on several parameters that determine the strength and the stiffness of the hydraulic fill mass. These parameters can be measured in the laboratory using reconstituted samples or in the mine using in situ testing devices. Due to the difficulties and high costs associated with carrying the in situ testing rigs into the underground openings, laboratory tests are the preferred alternatives. Strength and stiffness are directly related to the relative density of the fill. When the hydraulic fill is denser, the relative density and friction angle are higher, and thus the fill is more stable. In geotechnical engineering, there are several empirical correlations relating relative density to the Young’s modulus and friction angle of a granular soil. 2.2.1 Maximum and minimum dry density tests A larger void ratio does not always mean a looser granular soil. Relative density is a good measure of the density of the grain packing, and depends on the maximum and minimum possible void ratios for the soil whilst still maintaining intergranular contact. The minimum void ratio is generally determined by pouring the dry tailings from a fixed height so that the grains are placed at a very loose state [7]. The maximum void ratio is generally achieved by saturating the tailings and vibrating them to attain dense packing [8]. These two extreme void ratios provide the lower and upper bound for the void ratios, and, depending on where the current void ratio of the hydraulic fill is, the relative density is defined as: (1) Laboratory sedimentation exercises at JCU laboratories, during which hydraulic filling processes were simulated, showed consistently that when slurry settles under its self-weight, the relative density of the fill is in the range of 40---70% (Fig. 3). Fig. 3. Relative density of the hydraulic fills sedimented in the laboratory. Similar observations were made by Pettibone and Kealy [9] at selected mines in the United States. The in situ measurements showed relative density values ranging from 44 to 66% at four different mines. The laboratory exercise also showed that the hydraulic fill slurry settles to a dry density (g/cm3) of 0.6 times the specific gravity (Gs) for a wide range of tailings with specific gravity values ranging from 2.8 to 4.4. Dry density (rd) and void ratio (e) are related by: (2) This implies that all the hydraulic fills settle to a void ratio of 0.67 and porosity of 40%. The laboratory sedimentation exercise verifies this. 2.2.2 Oedometer tests Oedometer tests are carried out on hydraulic fills to determine the constitutive modelling parameters for the Cam Clay model e one of the constitutive models that can be adapted for hydraulic fills when analysed using numerical modelling packages such as FLAC, FLAC3D or ABAQUS. In addition, oedometer tests are useful in determining the constrained modulus (D) from which, Young’s modulus (E ) can be estimated for an assumed value of Poisson’s ratio using the following equation. (3) Young’s modulus is a crucial parameter in deformation calculations using most constitutive models. The oedometer tests on the hydraulic fills showed significant creep settlements that took place on the completion of consolidation settlements. This has yet to be verified quantitatively and on a full-scale stope. 2.2.3 Direct shear test Direct shear tests are carried out to determine the peak and residual friction angle of the hydraulic fill. The tests are carried out on reconstituted hydraulic fills representing the in situ grain packing in the stope, which can be at relative densities of 40--70%. Since there is no clay fraction, cohesion is zero. Direct shear tests conducted at JCU reveal that the friction angles determined from direct shear tests are significantly higher than those determined for common granular soils. This can be attributed to the very angular grains that result from crushing the rock Fig. 4. Scanning electron micrograph of a hydraulic fill sample. waste, which interlock more than the common granular soils. The angular grains can be seen in the scanning electron micrographs of the hydraulic fill samples (Fig. 4). 2.2.4 Placement property test A placement property test for hydraulic fills was proposed by Clark[10]. This is essentially a compaction test, where the compactive effort is applied through 5 min of vibration on a vibrating table. Porosity at the end of vibration is plotted against the water content. Alternatively, dry density can be plotted against water content, as shown in Fig. 5. Here a is the air content, and the contours of a=0, 3, 10, 20 and 30% are shown in the figure. The shaded region is where the hydraulic fill can exist whilst maintaining intergranular contact. The slurry follows a saturation line when settling under its self-weight, with the density increasing with some vibratory loading. One of the main applications of the placement property test, as in a compaction test, is to determine optimum water content. In Fig. 5, the optimum water content of the fill is 14%, with the maximum dry density of 2.42 t/m3. This water content can also be estimated from a maximum dry density test and the saturation line as 12%. These curves are useful in assessing the contractive or dilative behaviour of hydraulic fills at various water contents. For example, when the fill in Fig. 5 is subjected to vibratory loading (e.g., due to blasting) at 14% water content and a dry density of 2.0 t/ m3, it will densify, whilst the same fill at 8% water content and dry density of 2.2 t/m3 will become looser. Fig. 5. Placement property curve of a hydraulic fill sample. 3. Barricade bricks for hydraulic fill mines Barricade failure in underground mining operations is a primary safety concern because of the potential consequences of failure. Between 1980 and 1997, 11 barricade failures were recorded at Mount Isa Mines in both hydraulic and cemented hydraulic fills[5]. In 2000, a barricade failure at the Normandy Bronzewing Mine in Western Australia resulted in a triple fatality, and two permeable brick failures were reported later that same year as a result of hydraulic fill containment at the Osborne Mine in Queensland [1]. The specialized barricade bricks often used for the containment of hydraulic fill in underground mines are generally constructed of a mortar composed of mixture of gravel, sand, cement and water at the approximate ratio of 40:40:5:1, espectively. Fig. 6 shows a photograph of (a), a barricade brick and (b), an underground containment wall constructed from bricks. Traditionally, the walls have been constructed in a vertical plane, but the recent industry trend has been to increase wall strength by constructing them in a curved manner, with the convex toward the hydraulic fill as shown in Fig. 6b. (a) (b) Fig. 6. Porous brick barricade. (a) A brick, (b) brick barricade under construction in a mine. Although it is known within the mining industry that the porous bricks used in underground barricade construction are prone to variability in strength properties [5], the manufacturers often guarantee a minimum value for uniaxial compressive strength for the bricks in the order of 10 MPa[11]. Kuganathan[5] and Duffield et al. [11] have reported uniaxial compressive strength values from 5 MPa to over 26 MPa. A series of uniaxial compressive strength tests undertaken on a large sample of brick cores have demonstrated the scatter of results, but more importantly, have highlighted a distinct variation in brick performance when saturated, as it would occur in the mines. Two identical cylindrical cores were cut from 29 porous barricade bricks. One of the brick cores from each of the individual bricks was tested dry, and the other core was tested after having been saturated for either 7 or 90 days. The strength and deformation parameters (namely, the uniaxial strength, Young’s modulus, and the axial failure strain) for the wet and dry cores are shown in Figs. 7--9. Fig. 7. Uniaxial strength of dry and wet bricks. Fig. 8. Young’s modulus of dry and wet bricks. Firstly, the extreme scatter between all results reiterates the significant deviation in brick quality. Fig. 7 shows the average uniaxial compressive strength of dry bricks to fall between 6 and 10 MPa, when the brick manufacturers guarantee minimum of 10 MPa. It can also be seen from this figure that there is a distinct loss of compressive strength as a result of wetting the brick. There was no significant difference between 7 and 90 days soaking, implying that the strength loss occurs immediately upon wetting. This loss appears to be in the order of approximately 25%, which is notable considering that bricks are generally exposed to saturated conditions when placed underground, and all manufacturer strength specifications are based on bricks that are tested dry. The stiffness also appears to be reduced by wetting (Fig. 8). The Young’s modulus of the dry cores ranged between 1 and 3.5 MPa. The length of time the bricks were wetted did not have a significant impact on the magnitude of the reduction in stiffness. The peak failure axial strain was not reduced by wetting (Fig. 9). The cores in general failed under an axial strain of less than 1%. The porous bricks are designed to be free draining and therefore, their permeability is at least an order of magnitude greater than that of hydraulic fill. The barricade bricks have proven, over time, to satisfy the free-draining situation, and the reduction of permeability through mitigation of fines has not been recorded. Rankine et al. [4] carried out constant head and falling head permeability tests on several barricade bricks and reported permeability values in the order of 3500 mm/h, three orders of magnitude greater than the permeability of the tailings. Fig. 9. Axial failure strains of dry and wet bricks. 4. Paste fill Like hydraulic fill, paste fill falls into the category of thickened tailings. A conceptual framework to describe thickened tailings in terms of concentration and strength is shown in Fig. 10 [12,13]. Fig. 10. Thickened tailings continuum [13]. Paste fill is comprised of full mill tailings with a typical effective grain size of 5 mm, mixed with a small percentage of binder, in the order of 3--6% by weight, and water. It is the densest form of backfill in the spectrum of thickened tailings placed underground as a backfill material. The acceptance of paste backfill as a viable alternative to hydraulic slurry and rock fill did not truly occur until the mid- to late- 1990s with the construction and successful operation of several paste backfill systems in Canada and the BHP Billiton Cannington Mine in Australia. Since a desliming of the tailings is not undertaken, there is a substantial fine content in paste fills (Fig. 1). A generic ‘‘rule of thumb’’ for the grain size distribution is for a minimum of 15% of the material to be finer than 20 mm, which ensures that the surface area of the grains is large enough to provide adequate surface tension to ensure that the water is held to the solid particles and to provide a very thin, permanent lubricating film. Paste fill typically shows non-Newtoniane Bingham plastic flow characteristics, resulting in plug flow (batches flow in solid slugs) characteristics of the paste. As most of the early research performed on paste fills was on the transportation and deposition of the paste, the majority of the definitions of the paste are based on its rheological characteristics. Table 1 summarises some common characteristics of the thickened tailings continuum shown in Fig. 10 [14]. Hydraulic fills fall into the thickened tailings profile. A significant difference to note is that the water content in paste fill is retained on placement, through the large surface area of the grains, eliminating the need for the design of drainage of the fill or barricades. The design requirements for paste filled stopes are then reduced to static and dynamic stability requirements. By designing the fill masses with sufficient strength to ensure the vertical faces of the back filled stopes remain stable throughout the mining of the adjacent stopes, the static stability requirements are satisfied. If the paste becomes unstable, the adjacent faces may relax and displace into the open stope, causing high levels of dilution and loss of mining economies. The required strength of the backfills is typically calculated using analytical solution techniques. More recently, numerical modelling solutions have been used to determine backfill stability throughout the entire mining sequence. The dynamic stability of the paste fill stopes is addressed by designing the backfill mass to resist liquefaction or