451-470_Schellman 45

The South African Institute of Mining and Metallurgy International Symposium on Stability of Rock Slopes in Open Pit Mining and Civil Engineering Manuel Schellman

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SLOPE DESIGN FOR THE EAST WALL OF MANTOVERDE MINE, CHAARAL, CHILE

Mr Manuel Schellman Senior Geotechnical Engineer, Anglo American Chile

Mr Ricardo Sepulveda Geotechnical Consultant, A. Karzulovic & Assoc. Ltd.

Mr Antonio Karzulovic Geotechnical Consultant, A. Karzulovic & Assoc. Ltd.

ABSTRACT

Mantoverde Division of Empresa Minera Mantos Blancos S.A., Anglo American Chile, developed an open pit mine to exploit copper oxide ore. Mantoverde mine is located in the Province of Chaaral, Northern Chile, about 1,020 Km far from Santiago. Cur-rently, the East wall of the mine is reaching the condition of final pit. To evaluate the possibility of a pit optimization on this wall a geotechnical study was performed, in-cluding slope stability analyses by limit equilibrium and numerical methods. As a result of this geotechnical evaluation, it was decided to allow a probability of failure of 12% for instabilities affecting up to 600,000 tonnes, and use steeper slopes on the East Wall of the mine. This paper describes this geotechnical evaluation and how the results of the stability analyses were used to define a flexible slope design, which allows modification if the observed slope behaviour deviated from the expected one.

INTRODUCTION

Mantoverde Division of Empresa Minera Mantos Blancos S.A., Anglo American Chile, developed an open pit mine to exploit copper oxide ore. Mantoverde mine (see Figure 1) is located in the Province of Copiap, Northern Chile, about 1,020 Km far from San-tiago, at an elevation of 900 masl. The open pit operation began in 1995 and no slope instabilities have occurred, except some minor wedge instabilities at bench scale. Hence, Mantoverde Division decided to study the possibility of a slope steepening on the East wall of the pit, where the final pit condition is being reached. This study included an updated geotechnical characterization of the East sector, and slope stability analyses by limit equilibrium and numerical methods. The results of these analyses indicated that accepting a probability of failure of 12% for slope insta-bilities affecting up to 600,000 tonnes, the East wall interramp slopes could be steep-ened locally by up to 7o. Mantoverde Division accepted this more aggressive acceptability criterion, modified the slope design at the East wall, and initiated a continuous geotechnical monitoring of the East wall to maintain a safe operation.


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ENGINEERING GEOLOGY

The main rock types in the East wall of Mantoverde mine are: andesites (AND), brec-ciated andesites or transition zone (ZT), hydrothermal breccia (BXH), tectonic brec-cia (BXT), green breccia (BXV), diorite and other intrusive rocks (INT). Figure 2 shows the geology of a typical cross section through the East wall. These rocks present hydrothermal alteration of the silica-chlorite type, with potassic feldspar in veinlets and accretion veins, with a weak sericitic pervasive alteration. In some sectors, the rocks also present argillic alteration. At Mantoverde the mineralization corresponds to copper oxides. The main structural feature at the mine is the Mantoverde Fault, a major geological fault with an extent of about 14 Km, a strike N20W and a dip of 50E. The structural data was obtained from bench mapping. There are three main structural trends: subvertical structures with N30oE and N30oW strikes, and NS structures dipping 35o to 55o. The geological faults mapped on the East wall are shown in Figure 3. The groundwater level is deeper than the bottom of the final pit.

GEOTECHNICAL CHARACTERIZATION

Considering the rock and alteration types, thirteen geotechnical units were defined in the East wall. The rock mass properties for these geotechnical units were evaluated with the Hoek-Brown criterion (Hoek et al, 2002), using the results of laboratory testing (unconfined compression and triaxial tests), and the geotechnical quality of the rock mass defined from bench mappings and drill core logging.

Figure 1: Overall view of Mantoverde mine. The East wall can be seen on the centre.


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The geotechnical quality of the rock mass was rated using the Rock Mass rating (RMR) defined by Bieniawski (1979), for drill core and the Geological Strength Index (GSI) defined by Hoek (1994), for outcrop mapping. The effect of the uncertainty on the geotechnical data was included as suggested by Hoek (1998), using an Excel spreadsheet and the Excel add-on program @RISK (available from www.palisade.com).

Cross Section N 100,300

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Figure 2: Cross section N 100,300 showing the typical lithology of the East wall of Man-toverde mine, the current and the final pit.

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Figure 3: Map showing the geological faults mapped on Mantoverde mines East wall.


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Table 1 GEOTECHNICAL PROPERTIES OF THE ROCK MASS

Geotechnical Unit ci c B G Lithology Alteration Code (ton/m3) (MPa)

mi GSI mb (kPa) (degrees) (GPa) (GPa)

Silica-Chlorite AND 70 35 to 60 0.70 700 37 2.8 1.7 Andesite

Argillic AND-ARG 2.66

62 19.6

25 to 50 0.37 500 31 1.7 0.9

Silica-Chlorite ZT 64 35 to 60 0.42 600 32 2.7 1.7 Transition Zone Argillic ZT-ARG

2.68 58

12.0 25 to 50 0.22 400 26 1.6 0.9

Silica-Chlorite BXH 62 40 to 60 0.53 675 34 3.0 1.8 Hydrothermal Breccia Argillic BXH-ARG

2.88 56

13.0 30 to 50 0.28 450 28 1.7 1.0

Silica-Chlorite BXT 30 30 to 55 0.47 450 28 1.6 0.9 Tectonic Breccia Argillic BXT-ARG

2.46 26

15.7 20 to 45 0.21 275 21 0.9 0.4

Silica-Chlorite BXT 80 35 to 60 0.39 625 34 3.0 1.9 Green Breccia Argillic BXT-ARG

2.65 72

10.8 25 to 50 0.21 425 27 1.8 1.0

Silica-Chlorite INT 52 35 to 55 0.30 475 28 2.1 1.3 Intrusive Rocks Argillic INT-ARG

2.68 46

10.5 25 to 45 0.16 325 22 1.2 0.7

Fault Zones 2.20 60 24 0.5 0.3

Unit weight mi Hoek-Browns m parameter for the intact rock ci Intact rocks unconfined compressive strength mb Hoek-Browns m parameter for the rock mass GSI Geological strength index (Hoek, 1994) c Cohesion of the rock mass B Bulk modulus of the rock mass Friction angle of the rock mass

G Shear modulus of the rock mass

Equivalent values for the rock mass cohesion and friction angle were computed from the Hoek-Brown strength envelope, for confining pressures ranging from 0 to 2.5 MPa. The rock mass modulus (E) was estimated as suggested by Hoek et al (2002):

( )( )40/10101002

1 GSIciD E

=

(1)

where the units of E are MPa, ci is the unconfined compressive strength of the intact rock (in MPa), and D is a factor which depends upon the degree of disturbance to which the rock mass has been subjected by blast damage and stress relaxation (in this case D ranges from 0.8 to 1.0). The Poisson ratio of the rock mass, , was estimated as:

7.001.040.0 GSI = (2) The values of E and were used to compute the bulk (B) and shear (G) modulus of the rock mass. These geotechnical properties are summarized in Table 1. Laboratory direct shear tests were performed on samples containing joints. From the results of these tests and the experience from other Chilean open pits, it was considered that: Major structures such as large faults and shears are continuous (ie they could affect

the stability of interramp and overall slopes). Their shear strength is defined by a co-hesion of 60 20 kPa and a friction angle of 24o 4o (values from back analyses of planar slides on faults with clay gouge in some Chilean open pits).


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Minor faults, whose lengths usually range from 20 to 40 m, are continuous at bench scale (including double benches). Their shear strength is defined by a cohesion of 105 20 kPa and a friction angle of 26o 3o (estimated values).

Joints are continuous at bench scale (including double benches). Their shear strength is defined by a cohesion of 95 20 kPa and a friction angle of 36o 3o (values from results of direct shear tests and back analyses of planar and wedge slides at bench scale in other Chilean open pits).

Bench-scale structures such as minor faults and joints, are discon-tinuous and contain rock bridges at the scale of interramp and overall slopes. To include the effect of these rock bridges, probabilistic analyses were performed to define the strength of equivalent struc-tures like the one illustrated in Figure 4. These probabilistic analy-ses consider the mean and the stan-dard deviation for the strength of the rock and the structures, and for the length, dip and spacing of the structures, in a form similar to the one described by Baczynski (2000). This computation was done only for those structures subparallel to the slopes (ie the angle between the slope and the structures strike is equal or smaller than 20o).

Isotropic rock mass strength can be assumed if: (a) the rock mass is very massive and does not have joints, (b) the rock mass is fractured but there are no structural sets with a strike subparallel to the slopes strike (ie the angle between them is greater than 20o), or (c) the rock mass is highly fractured and there are so many structures that there are no predominant joint orientations. In these cases the Hoek-Brown criterion is applied straightforward manner, and the strength properties of the rock mass are those of Table 1 in a straightforward manner. Isotropic rock mass strength cannot be assumed if: (a) the rock mass is layered/foliated, or (b) the rock mass is fractured, and contains one ore more predominant joint sets with a strike subparallel to the slopes orientation. In these cases the directional effect of the planes of weakness, defined by the structures or equivalent structures if they contain rock bridges, is included in the following form: The apparent dip of the structures (a) and its possible variation (a) are defined in

the cross section to be analyzed. The strength of the structure (Sj) and the rock mass (Srm) are defined (note that the

term strength is generic and could mean cohesion or friction).

Rock bridge

Equivalent structure

Joint Set 1

Joint Set 2

Rupture surface through joints and rock bridges, defining an equivalent structure

Rock bridge

Equivalent structure

Joint Set 1

Joint Set 2

Rupture surface through joints and rock bridges, defining an equivalent structure

Figure 4: Equivalent structure defined by a failure through joints and rock bridges, in a pit slope much higher than the mean length of the structures dipping towards the pit.


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The strengths of the rock mass strength and the structure are overlapped defining a directional strength, as shown in Figure 5.

If the rock mass contain two predominant joint sets with strike subparallel to the slope the same procedure is repeated for the second joint set, as shown in Figure 6.

ACCEPTABILITY CRITERIA FOR SLOPE DESIGN

At Mantoverde mine the following acceptability criteria were defined for slope design: Benches: Double benches with a total height of 20 m and a bench face inclination of 80o. The berm or bench width should contain the debris from a planar/wedge slide from the upper bench whose volume has a probability of exceedance of 20% (ie the prob-ability of a larger instability is 20%). Final Pit Slopes: There isnt any in pit infrastructure. The surface infrastructure is far away from the East wall, and cannot be affected by subsidence problems. Hence, the following criteria are applied in terms of the tonnage affected by the potential failure (W, defined in tonnes per meter because stability analyses are bidimensional), the factor of safety (FS), and the probability of failure (PF, defined as the probability of FS being smaller than one): - If W 15,000 tons/m, then FS 1.20 and PF 12% - If 15,000 tons/m < W 30,000 tons/m, then FS 1.25 and PF 10% - If 30,000 tons/m < W, then FS 1.30 and PF 8%

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Predominant joint orientationwith strike subparallel to the slope

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Figure 5: Definition of directional strength for a rock mass with one joint set subparallel to the slope. The radial distance from the centre to the thick curve defines the strength (eg cohesion), which varies with the inclination of a potential failure plane.

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Figure 6: Definition of directional strength for a rock mass with two joint sets subparallel to the slope. The radial distance from the centre to the thick curve defines the strength (eg cohesion), which varies with the inclination of a potential failure plane.


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The following must be pointed out regarding these acceptability criteria: As shown in Figures 7 and 8, the current blasting practices at Mantoverde mine

achieve well groomed double benches with bench face inclinations of 80o. Hence, this is a realistic design assumption.

Figure 7: Mantoverde mines East wall (looking SE). The well groomed double benches in the upper were achieved using presplitting (20 m vertical blastholes, = 6).

Figure 8: Detail of the well groomed double benches achieved using presplitting (20 m vertical blastholes, = 6).


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The limit of 15,000 ton/m for the maximum probability of failure is due to the fact that an eventual interramp slope failure with a lateral extent of about 40 m (expected lateral extent for 120 to 150 m high interramp slopes), would affect 600,000 tonnes. This tonnage can be removed in one week by the current mining equipment (two front loaders Caterpillar 994). This would have no effect on the plant operation because there is a stockpile of about 800,000 tons that could keep on feeding the plant.

Considering that the typical pit slope design means FS 1.3, which is more or less equivalent to the condition PF 8%, these acceptability criteria are approximately equivalent to increase 25% and 50% the permissible probability for a relatively large failure (up to 1,200,000 tons) and a manageable failure (up to 600,000 tons), respectively.

The seismic condition is considered an eventual loading condition that has a low probability of occurrence during the life of the pit. Hence, for seismic loading it was considered that FS 1.1 was acceptable for pseudostatic analyses with a horizontal seismic coefficient (kh) equals to 0.12g.

A more aggressive slope design requires good slope management. This means good operational practices (good quality blasting, a good fulfilment of the design slope geometry, etc.), good geotechnical monitoring and slope control. Indeed, Man-toverde Division began a training program on slope performance and control aimed to general foremen and mine operators. Currently the East wall is under continuous monitoring by a robotic Leica APS-Win system.

BENCH-BERM DESIGN

A bench height (hB) of 10 m was defined from operational considerations regarding ef-ficiency of the loading equipment, and a double bench design was selected (20 m total height). The expected bench face inclination () was defined as 80o, in agreement with the cur-rent blasting practices and observed bench performance (see Figures 7 and 8), con-sidering the fact that presplitting will be used. Monte Carlo simulations were carried out to define the cumulative distribution of the tonnages associated with structurally controlled bench instabilities (ie planar and wedge slides). From these cumulative curves the tonnage with a 20% probability of ex-ceedance was selected, and a swelled volume was computed assuming a 2.00 tons/m3 unit weight for the debris or broken rock. The basal length of this volume (ld) was computed assuming that the debris is cohe-sionless and its friction angle is 39o. The berm or bench width (wb) required was computed as ld plus one meter. According to this, the minimum berm or bench width required for the benches of the East wall ranges from 8.0 to 11.6 m.


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SLOPE STABILITY ANALYSES, EVALUATION AND DESIGN

Limit equilibrium was used to analyze potential stability problems structurally controlled, and defined by geological faults at interramp and/or overall slopes (see Figure 3). These analyses were performed using the softwares DIPS, ROCPLANE and SWEDGE (available from www.rocscience.com), and the results obtained indicated that the geological faults do not define potentially unstable blocks/wedges on the slopes of the East wall. To assess the stability of interramp and overall slopes in Mantoverde mines East wall, six vertical cross sections were defined as shown in Figure 9.

Each section was analyzed using the Generalized Limit Equilibrium (GLE) method available in the software SLIDE (available from www.rocscience.com), which is equivalent to the Morgenstern-Price method (Morgenstern & Price, 1965). The stability analyses were done according to the following: In each section all the possible failures were considered: one interramp, more than

one interramp, overall slope, and local failures associated to lithological contacts and/or geological faults.

The path search technique and four segment lengths were used to find the most critical failure surface in each case. To do this 5,000 to 10,000 failure surfaces were generated and once the critical surface was found, the optimization option of SLIDE was used to find the critical FS.

For the most critical failure surface the probability of failure was computed using the Taylor series method suggested by Duncan (2000), considering a coefficient of variation equals to 17% and 25% for the angle of friction and the cohesion, respectively and a variation of 5o for the apparent dip of those structures that were subparallel to the slope.

Tension cracks were assumed to reach depths up to 10% of the slope height being analyzed.

If required (eg due to a joint set with a strike subparallel to the slope orientation) a directional rock mass strength was used, as previously discussed.

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Figure 9: Vertical cross sections for slope stability analyses.


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As the groundwater level is below the bottom of the final pit, the slopes are in a dry condition.

The current East walls slopes were analyzed and, as expected, the results obtained indicated that these slopes are very stable (FS 1.7 and PF 1% in all cases). Figure 10 shows a typical result from these slope stability analyses.

Then the original slope design for the final pit on the East wall was analyzed. The results obtained indicated that: The design can be optimized in the sectors of sections N-99,700, N-99,950 and N-

101,100 (ie in the Northern and Southern parts of the East wall), steepening the slopes.

In the sectors of sections N-100,525 and N-100,700 the design can be maintained because it fulfils the acceptability criteria previously defined.

In the sector of section N-100,300 (in the central part of the East wall), the original slope design did not fulfil the acceptability criteria previously described.

According to these results, the slope design was modified by steepening the slopes in those sectors where the design can be optimized, and by modifying the slope geometry at section N-100,300 in order to fulfil the acceptability criteria (two options were considered). The results of the slope stability analyses for the original and modified slope designs are summarized in Table 2.

TensionCracks

Critical Failure SurfaceFS = 1.90

MANTOVERDE MINEEAST WALLSECTION N-100,525CURRENT SLOPE

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Figure 10: Slope stability analysis of East walls current slopes. This example corresponds to section N-100,525, where the critical failure surface for the 220 m overall slope has a FS of 1.90 and a PF smaller than 1%.


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Table 2 RESULTS OF THE SLOPE STABILITY ANALYSES

N-99,700 N-99,950 N-100,300 N-100,525 N-100,700 N-101,000 Parameters

OD MD OD MD OD MD1 MD2 OD MD OD MD OD MD

ho (m) 110 110 155 155 280 280 280 275 230 110 110 hf (m) 100 100 155 110 280 280 280 275 230 110 90 r (degrees) 53o 54o 53o 59o 53o 53o 53o 53o 53o 53o 60o r (degrees) +1o +6o 0o 0o +7o o (degrees) 51o 51o 45o 45o 52o 49o 45o 50o 46o 38o 38o o (degrees) 0o 0o -3o -7o 0o W (kton/m) 6.8 6.8 25.0 10.6 57.5 62.6 63.3 66.5 43.8 16.8 7.1 FS 1.39 1.38 1.40 1.27 1.21 1.31 1.32 1.30 1.35 2.28 1.95

FSEQ 1.20 1.20 1.17 1.11 1.03 1.11 1.11 1.10 1.11 1.92 1.88

PF (%) 4 4 3 10 14 7 7 8 3 < 1 < 1 ho Overall slope height r Interramp angle (measured from toe to toe) hf Slope height affect by critical failure surface r Interramp angle increment (+) / decrement (-) W Tonnage affected by critical failure surface o Overall slope angle (measured from toe to crest) FS Factor of safety (critical failure surface) o Overall slope angle increment (+) / decrement (-) FSEQ Seismic factor of safety (kh = 0.12g) OD Original slope design PF Probability of failure (critical failure surface) MD Modified slope design MD1 Modified slope design - Option 1 MD2 Modified slope design - Option 2

Hence, the East wall slopes can be steepened in some sectors, increasing the interramp angles an average of 3o in the Southern sector, and 7o in the Northern sector. On the other hand, the geotechnical evaluation showed that the slope in the central part of the East wall (section N-100,300) did not fulfil the acceptability criteria (FS = 1.21 < 1.30 and PF = 14% > 8%, see Figure 11). Two options were considered to modify the slope geometry in this section and fulfil the acceptability criteria: Option 1: The benches in the lower part of the slope (at elevations 870 and 850, being

810 the elevation of the final pit bottom) have wider berms or bench widths (23.1 m), defining an abutment toe. This modified slope geometry fulfils the acceptability criteria (W = 62.6 kton/m, FS = 1.31 > 1.30 and PF = 7% < 8%, see Figure 12).

Option 2: The geometry of the upper part of the slope is modified, leaving a 60 m wide platform at elevation 1030, and at elevation 850 the bench has a wider berm or bench width (23.1 m). This modified slope geometry fulfils the acceptability criteria (W = 63.3 kton/m, FS = 1.32 > 1.30 and PF = 7% < 8%, see Figure 13).


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TensionCracks


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MANTOVERDE MINEEAST WALLSECTION N-100,300FINAL PIT CONDITIONORIGINAL SLOPE DESIGN

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MANTOVERDE MINEEAST WALLSECTION N-100,300FINAL PIT CONDITIONORIGINAL SLOPE DESIGN

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Figure 11: Slope stability analysis of section N-100,300, original slope design (final pit con-dition). Acceptability criteria are not fulfilled (W = 57.5 kton/m, FS = 1.21 and PF = 14%).

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MANTOVERDE MINEEAST WALLSECTION N-100,300FINAL PIT CONDITIONMODIFIED SLOPE DESIGNOPTION 1

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OriginalSlope Design


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Figure 12: Slope stability analysis of section N-100,300, modified slope design - Option 1. Acceptability criteria are fulfilled (W = 62.6 kton/m, FS = 1.31 and PF = 7%).


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Option 1 leaves a larger abutment toe in ore, but it seems preferable to Option 2 which requires a relatively large removal of waste rock. On the other hand, the goal was to extract as much ore as possible and, in spite of the fact that the original design for the final pit condition did not fulfil the acceptability criteria, some questions arise: Is it really so bad to have FS = 1.21 instead of FS = 1.30? Could we afford a probability of failure of 14% for a relatively large slope failure? Which is the best course of action considering the mining business? Hence, it was decided to compare the original slope design with Option 1 by means of numerical analyses using the software FLAC (available from www.itscacg.com), including the geological structures (or equivalent structures to consider the effect of rock bridges) as ubiquitous joints when they were subparallel to the slope orientation. These numerical models are shown in Figures 14 and 15. The results of these numerical analyses indicated that: In both cases, original and modified slope geometries, the slopes are stable (ie the

numerical models reached an equilibrium condition). The plasticity indicators (failure through the rock mass by shear and/or tension, and

slippage along ubiquitous joints) show the following:

TensionCracks

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Critical Failure SurfaceFS = 1.32280 m

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Critical Failure SurfaceFS = 1.32280 m

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Platform

Figure 13: Slope stability analysis of section N-100,300, modified slope design - Option 2. Acceptability criteria are fulfilled (W = 63.3 kton/m, FS = 1.32 and PF = 7%).


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FLAC (Version 5.00)

LEGEND

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- Original slope geometry (final pit condition): Tensile failures occur in the upper part of the slope and also, locally, at elevation 870. Shear failures occur mainly in the bottom of the pit and in the zones adjacent to geological faults. Slippage along ubiquitous joints occurs in the upper mid-part of the slope, especially at the contact between intrusive rocks and green breccias. The plasticity indicators do not define a continuous rupture surface through the overall slope (see Figure 16).

- Modified slope geometry Option 1 (final pit condition): Tensile failures occur in the upper part of the slope. Shear failures occur mainly in the bottom of the pit and in the zones adjacent to geological faults. Slippage along ubiquitous joints occurs in the upper mid-part of the slope, especially at the contact between intrusive rocks and green breccias. The plasticity indicators do not define a continuous surface of rupture through the overall slope (see Figure 17).

In both cases the slopes maximum horizontal displacements occur locally, at those benches affected by geological faults (see Figures 18 and 19). In the original slope geometry these displacements do not exceed 90 cm (see Figures 18); and in the case of the modified slope design Option 1, they do not exceed 80 cm (see Figure 19). Hence, in both cases the ratio between the slope crests horizontal displacement and the slope height is smaller than 0.5%, which means that the slopes do not show signs of instability (Hoek & Karzulovic (2001) suggested that slopes begin to show signs of instability when this ratio is about 1%).

In both cases the zones of maximum shear strain increment are associated to the presence of geological faults, and do not define a continuous surface of rupture through the slope (see Figures 20 and 21).

Considering all these results Mantoverde Division decided to optimize the mining business using what Terzaghi called the observational method (eg see Peck, 1969). Hence, the slopes of Section N-100,300 will be developed according with the original slope design, considering this a flexible design that can be modified to Option 1 if the slope monitoring indicates a sudden increase in the slope displacement rate when the wall toe reaches elevation 870 (Contingency Plan A), or modified to Option 2 if these sudden increase is detected when the wall toe reaches elevation 850.

CONCLUSIONS

The use of more aggressive acceptability criteria for slope design (considering that the typical case FS 1.3 is more or less equivalent to PF 8%, these acceptability criteria are approximately equivalent to increase 25% and 50% the permissible probability for a 1,200,000 tons failure and a 600,000 tons failure, respectively), the Mantoverde mines East wall interramp slopes were steepened locally up to 7o. The definition of these acceptability criteria considered the operational practices, and the fact that the eventual occurrence of a 600,000 tonnes can be cleaned-up in one week by the current mining equipment (two front loaders Caterpillar 994), having no effect on the plant operation because there is a stockpile of about 800,000 tons that could keep on feeding the plant.


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This optimized slope design requires good operational and blasting practices, and also an efficient slope instrumentation and monitoring system. Currently the quality of the blast allow to achieve well groomed 20 m double benches, with 80o bench face inclina-tions, and the East wall is under continuous monitoring by a robotic Leica APS-Win system. Mantoverde Division also began a training program on slope performance and control aimed to general foremen and mine operators. Limit equilibrium and numerical methods were used to assess slope stability, and the slope design was optimized by steepening the slopes in those sectors where it was possible to do so. In the central sector of the East wall the observational method was used to define an aggressive slope design, but flexible enough to allow modifications according to well defined contingency plans if the observed slope behaviour deviates from the expected one. These contingency plans were defined considering the results of limit equilibrium and, especially, numerical analyses of slope stability.

REFERENCES

BACZYNSKI, N R P. STEPSIM4 step-path method for slope risks, GEOENG2000: An International Conference on Geomechanical & Geological Engineering, Vol 2, pp 86. Technomic Publishing Co: Lancaster, Pennsylvania, USA, 2000. BIENIAWSKI, Z T. The geomechanics classification in rock engineering applications, Proc 4th Congress Int Soc Rock Mech, Montreux, Switzerland, Vol 2, pp 41-48, Balkema: Rotterdam, 1979. DUNCAN, J M. Factors of safety and reliability in geotechnical engineering, J. Geotechnical and Geoenvironmental Engineering, ASCE, 2000, 126(4):307-316. HOEK, E. Strength of rock and rock masses, ISRM NewsJ., 1994, 2(2):4-16. HOEK, E. Reliability of Hoek-brown estimates of rock mass properties and their impact on design, Technical Note, Int. J. Rock Mech. Min. Sci., 1998, 35(1): 63-68. HOEK, E, CARRANZA-TORRES, C & CORKUN, B. Hoek-Brown Failure Criterion - 2002 Edition, Mining and Tunnelling Innovation and Opportunity, Proc. 5th North Am Rock Mech Symp & 17th Tunn Assn Can Conf, Toronto (eds R Hammah, W Bawden, J Curran and M Telesnicki), Vol 1, pp 267-273. University of Toronto Press: Toronto, 2002. HOEK, E & KARZULOVIC, A. Rock-mass properties for surface mines, Slope Stability in Surface Mining (eds W A Hustrulid, M K McCarter and D J A Van Zyl), pp 59-67. University of Toronto Press: Toronto, 2002. MORGENSTERN, N R & PRICE, V E. The analysis of the stability of general slop surfaces, Geotechnique, 1965, 15(1):77-93. PECK, R B. Advantages and limitations of the observational method in applied soil mechanics, Ninth Rankine Lecture, Geotechnique, 19(2):171-187, 1969.


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