Backfill

σv

Span

Open Stope

Sill Mat Pillar

- Cohesion

- Friction Angle

-Tensile Strength

- Density

Numerical Modeling of Paste Sills in Underhand Cut & 

Fill Stopes 

Paul Hughesa, Rimas Pakalnisa, & Cristian Caceresa 

Wilson Blakeb 

Tom Bradyc 

a

Geomechanics Centre, Norman Keevil Institute of Mining, University British Columbia.

b

Consultant
 c

Spokane Research Lab (NIOSH)
 

ABSTRACT 

This paper describes a focus of work presently being conducted at the Rock Mechanics Research 

Group at the University of British Columbia. The underhand method under consolidated fill ensures 

a high recovery under an engineered back that is comprised of cemented rock fill and/or cemented 

paste fill. This method of mining has been employed successfully in mines throughout North 

America as a method of mitigating exposure to the operator having to work under a seismically 

active “back”. This paper reviews design methodology in the placement and analysis of sill mats 

with reference to site observation coupled with on going numerical and analytically derived 

solutions. 

Corresponding Author 

E-mail address: phughes@mining.ubc.ca 

1 INTRODUCTION 

With increasing demand on minerals and escalating market prices, the need to mine deeper depths 

in high stress mining environments is becoming more common. Within these high stress 

environments the frequency of rock bursts have increased. In order to provide a safe working 

environment underhand cut and fill with paste has become a more widely used mining method. The 

placement of consolidated backfill requires one to understand the overall factors affecting design. 

Figure 1 graphically summarizes some of the parameters that are being investigated in terms of their 

implication on developing a design span enabling man entry access. A sill for this study is defined 

as a consolidated layer of previously placed fill immediately above the mine opening that is being 

excavated. 

Through numerical modelling, the UBC Geomechanics Centre has done a sensitivity analysis of 

paste unconfined compressive strength (UCS) versus sill width. A further comparison has been 

with modelled UCS strengths compared to existing literature for determining span width based on 

UCS. 

Design Requirements 

Page 1 

Seismic

Backfill 

σv

Span 

Open Stope 

Seismic

• Vertical Loading 

• Stope Geometry Influences 

• Lateral Loadingal Loadingal Loading 

• Failure Modes h 

σ Sill Mat Pillar h • Strength Properties - Cohesion

• Fill Placement - Friction Angle 

-Tensile Strength • Seismic Effects 

- Density

• Support Implicatio β ns - Modulus, Poisson Ratio 

• Stope Closure 

• Other (Air Gaps/Cold Joint/Cold Joint/Cold Joints) 
 

 

Figure 1: Mining Under Consolidated Backfill 

2. DESIGN CONSTRAINTS 

Figure 1 shows the factors that have to be accounted for in terms of mining under an engineered 

back. These will be outlined in this paper from a general perspective with focus on the analytical 

and numerical assessment of span and applied loading conditions. 

2.1 Design Load 

A critical factor is estimating the design loads onto the sill mat. Caceres (2005) employs an existing 

database of a Canadian mine as a case study that looks at the loading conditions that exist on a 

cemented rockfill sill mat. Design loads are critical to determine (or “in determining”) the strength 

required of the sill mat for the given stope geometry as under-estimating can cause a premature 

failure of the sill mat once mining exposes the mat, whereas overestimating can result in 

unnecessary expenses due to the cost of the cement in place. Estimating the vertical loading is not a 

trivial solution as many factors affect the overall loading conditions as evident from the many 

theoretical derivations that are available as per Janssen (1895), Terzaghi et al.(1996), Reimbert 

(1976) and Blight (1984) all of which have significant assumptions in terms of coefficient of lateral 

earth pressure “K” as detailed by Marcinyshyn (1996). The typical geometry was modelled 

employing FLAC2D (Itasca, 2005) which did not have the constraints the analytical methods had in 

terms of ‘K” and stope inclination. An analytical approximation, as shown by Eqn. 1, was derived 

by Caceres (2005) relating the numerical simulation to an equivalent relationship. 

Eqn. 1: 

⎛ γ ⋅ L ⎞ ⎡ ⎛ 2 ⋅ K ⋅ tan(φ) ⋅ z ⎞⎤ 2 ⎢

⎣ ⎜

⎜

⎝

−−⋅ ⎟

⎟⋅

⎠ ⎜

⎜

⎝ ⋅⋅ = sin ( ) 1 exp tan( )2 ( ) β φ

σ

K

z y ⋅sin ( ) 2 L β ⎥

⎦

⎟

⎟

⎠

Where: 

-L = Span of stope 

-z = Height of fill 

-К = Coefficient of lateral earth pressure 

-γ = Unit weight of fill 

-Φ = Friction Angle of Fill 

-Β = Stope dip angle 

As mentioned, the above was derived for unconsolidated or uncemented rock fill, however, the 

analytical solution would be similar to that of unconsolidated paste as the input parameters would 

define the loading conditions. 

2.3 Failure Mechanism 

The methodology of span design under consolidated fill is complex as many factors control the 

overall stability as shown in Figure 1. The failure modes and combination thereof must be analysed 

with respect to the placed fill, stope geometry, loading conditions, seismic effects, stope closure, 

and support placement as well as other factors that are due to filling practises such as cold joints and 

gaps between successive lifts among others. This paper employs analytical, numerical and empirical 
 

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

2b 2c

a

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

2b 2c

a

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

2b 2c

a

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

2b 2c

a

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

2b 2c

a

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

3m THICK SILL

5m THICK SILL

10m THICK SILL

15m THICK SILL

20m THICK SILL

L

d

L

d

 

 

tools to attempt to provide an initial tool for design by the operator. The database of underhand 

stopes observed by the author is shown in Table 1, which is comprised of twelve(12) operations 

which include seven cemented rock fill and five having paste within the immediate back. 

The unconfined compressive strength is typically the parameter employed to benchmark the overall 

stability of the immediate back. The compiled database (Table 1, found at the end paper) of backfill 

unconfined compressive strengths was adapted from Souza et al.(2003). 

β

Uncemented

rockfill

Open Stope

Sill Mat

Caving

Rotation

Shear Shear

σv

σt τ τ σc

σc

β

Uncemented

rockfill

Open Stope

Sill Mat

Caving

Rotation

Shear Shear

σv

σt τ τ

β

Uncemented

rockfill

Open Stope

Sill Mat

Caving

Rotation

Shear Shear

σv

σt τ τ σc

σc

Where:

− L = Span of the stope

− γ = Rockfill’s unit weight

− st = Tensile strength of the cemented sill

− d = Thickness of sill

− sc = Horizontal confinement (assumed zero – conservative)

− sv = Vertical stress due loading above sill mat

− Tf = Shear strength along fill/wall contact

− β = Stope dip angle

πσγ L t 8 ⋅>⋅ γσ

σσ

⋅+

+⋅ ⎟ > ⎠

⎞ ⎜

⎝

⎛

d d

L

v

ct )(2 2

⎟

⎠

⎞ ⎜

⎝

⎛

⋅ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⋅>⋅+ L

d d t

v sin )( 2)( 2 β

τ γσ ( cot( )) sin )( )( 2

2

ββ

σ γσ ⋅⋅−⋅⋅ >⋅+ dLLd d t

v

Caving Failure Flexural Failure

Sliding Failure Rotational Failure

a) Schematic showing typical failure modes 

after Mitchell, 1991.

b) Limit equilibrium analysis of typical failure modes adapted from Mitchell, 1991.

β

Uncemented

rockfill

Open Stope

Sill Mat

Caving

Rotation

Shear Shear

σv

σt τ τ σc

σc

β

Uncemented

rockfill

Open Stope

Sill Mat

Caving

Rotation

Shear Shear

σv

σt τ τ

β

Uncemented 

rockfill 

Open Stope 

Sill Mat 

Caving 

Rotation 

Shear Shear

σv 

σt τ τ

L 

σ d c 

σc 

Where: 

− L = Span of the stope 

− γ = Rockfill’s unit weight 

− st = Tensile strength of the cemented sill 

− d = Thickness of sill 

− sc = Horizontal confinement (assumed zero – conservative) 

− sv = Vertical stress due loading above sill mat 

− Tf = Shear strength along fill/wall contact 

− β = Stope dip angle 

πσγ L t 8 ⋅>⋅ γσ

σσ

⋅+

+⋅ ⎟ > ⎠

⎞ ⎜

⎝

⎛

d d

L 

v 

ct )(2 2 

⎟

⎠

⎞ ⎜

⎝

⎛

⋅ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⋅>⋅+ L 

d d t 

v sin )( 2)( 2 β

τ γσ ( cot( )) sin )( )( 2 

2 

ββ

σ γσ ⋅⋅−⋅

⋅ >⋅+ dLL d d t 

v 

Caving Failure Flexural Failure 

Sliding Failure Rotational Failure 

a) Schematic showing typical failure modes 

after Mitchell, 1991. 

b) Limit equilibrium analysis of typical failure modes adapted from Mitchell, 1991. 

Figure 2: Limit equilibrium criteria adapted from Mitchell, 1991. 

The design methods (Table 1) all employed a form of limit equilibrium analysis coupled with 

modelling. The failure modes are summarized by Mitchell (1991) and shown in Figure 2. 

DESIGN OF UNDERCUT SILL SPANS 

MINIMUM SPAN WIDTH (m)

UNCONFINED COMPRESSIVE STRENGTH (MPa)

3m THICK SILL 

5m THICK SILL 

10m THICK SILL 

15m THICK SILL 

20m THICK SILL 

2b 2c

a 

MINIMUM SPAN WIDTH (m) 

UNCONFINED COMPRESSIVE STRENGTH (MPa) 

UNDERHAND CUT AND FILL OPERATIONS 

Figure 3: Stability chart for the design of undercut sills with vertical sidewalls with a FS of 2. Chart is based 

upon fixed beam bending failure with surcharge, adapted from Stone, 1993. (Pakalnis, 2005) 

Flexural instability was found to be most critical in the absence of rotational instability and closure 

stresses (sc) which have to be evaluated separately. Stone (1993) had concluded that, for cemented 
 

Chainl nk Fenc ng over 

plates 

Tie W re attached to 

Dyw dag bo t and to 

wa

6ft Dyw dag

Bo t (typ) 

6” Prep muck 

max mum

Chainl nk Fenc ng over 

plates 

Tie W re attached to 

Dyw dag bo t and to 

wa

6ft Dyw dag

Bo t (typ) 

6” Prep muck 

max mum

Chainl nk Fenc ng over 

plates 

Tie W re attached to 

Dyw dag bo t and to 

wa

6ft Dyw dag

Bo t (typ) 

6” Prep muck 

max mum

PLAN

VERTICAL VIEW

3D VIEW

Chainl nk Fenc ng over 

plates 

Tie W re attached to 

Dyw dag bo t and to 

wa

6ft Dyw dag

Bo t (typ) 

6” Prep muck 

max mum

Chainl nk Fenc ng over 

plates 

Tie W re attached to 

Dyw dag bo t and to 

wa

6ft Dyw dag

Bo t (typ) 

6” Prep muck 

max mum

 

rock fills, crushing, caving and sliding are generally negated when the sill thickness exceeds 0.5 x 

span. In the absence of closure stresses, when the unconfined compressive strength of the cemented 

rock fill is greater than 1.5MPa, the kinematically possible rotational instability has to be analysed 

separately. Figure 3 shows the database that has been compiled in Table 1 and plotted onto a 

stability chart adapted from Stone(1993) and developed for the design of sills with vertical 

sidewalls with a Factor of Safety of two. The chart is based upon flexural instability employing 

fixed beam analysis with surcharge loading after Eqn. 1. It shows the unconfined compressive 

strength required (FS=2) for a given sill thickness and span exposed and related to actual field 

observations. Generally the mine data was found to be more conservative than the required for a 

Factor of Safety of 2.0. This may reflect the quality control requirements at individual operations, 

along with other factors such as seismicity and stope geometry among others as shown in Figure 1. 

Caceres (2005) simulated the limit equilibrium approach shown in Figure 2 employing FLAC2D 

models (finite difference code) for a given value of cohesion, friction angle, vertical surcharge, span 

and stope dip. The backfill properties assigned are for a Mohr-Coulomb type of material with strainsoftening behaviour where integrity is lost after 1.5% strain (Swan and Brummer, 2001). The 

resultant mode of failure was analysed for various stope dips with cohesion on the hanging wall 

contact varying from zero to maximum value (equal to the cohesion value of sill). From the 

simulations, it was found that the analytical approach after Mitchell,(1991) which assumes no 

hanging wall cohesion for the rotational instability and this resulted in a high degree of 

conservatism. 

2.3 Other Factors 

The above attempts to outline a methodology for span design. It is critical that the method be 

calibrated for individual sites, incorporating critical factors such as seismic conditions, installed 

support, wall closure and methods of fill placement as these all play a significant role in ensuring a 

safe exposed operating span. A major benefit of mining under paste is the mitigation of the hazards 

posed by bursting (Blake et al, 2004). 

3 SEISMIC CASE HISTORY – MANAGING ROCKBURSTS BY EMPLOYING 

UNDERHAND PASTE FILL AT HECLA’S LUCKY FRIDAY MINE, MULLAN IDAHO 

The following have been compiled by Blake and Hedley, 2003. Its importance is that the underhand 

mining as practised at Lucky Friday (Mine #12 in Table 1), is the first to incorporate paste to 

mitigate burst damage and the method has been adopted at mines throughout North America such as 

the Red Lake Mine in Ontario (Mah et al., 2003) and the Stillwater Mine in Montana (Jordan et al., 

2003). 

i i

i

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ll

i

l

i

i i

i

i l

ll

i

l

i

i i

i

i l

ll

i

l

i

i i

i

i l

ll

i

l

i

i i

i

i l

ll

i

l

i

Chainlink Fencing over 

plates 

Tie Wire attached to 

Dywidag bolt and to 

wall 

6ft Dywidag 

Bolt (typ) 

6” Prep muck 

maximum 

PLAN 

VERTICAL VIEW 

3D VIEW 


 

 

 

Figure 4: Lucky Friday Mat 

Hecla initiated overhand cut-and-fill mining on the Silver Vein at the Lucky Friday Mine in the late 

1950’s. By the mid 1960’s mining had progressed down to the 3050 level (~930m below surface), 

and the mining geometry consisted of long, flat-backed stopes, all at the same elevation, being 

carried up from two or more levels simultaneously. A burst prone sill pillar was formed when 

mining from below would approach the overlying mined out level. As a result of a double rockburst 

fatality in 1969, the mining front was changed to“centre lead stope” geometry. In 1973 the first 

computer controlled seismic monitoring system was installed, and pillar distressing was routinely 

carried out when a sill pillar was mined to approximately 12m (thickness). 

This rockburst strategy allowed mining to proceed safely down to below the 4660 level (~1420m 

below surface). In 1982 the mining front entered a highly burst prone formation, and serious rock 

burst problems were encountered. As a result of rockburst fatalities in 1984 and 1985, Hecla 

initiated an experimental underhand cut-and-fill stope along the east abutment of the mine. After 

another rockburst fatality in March 1986 Hecla realized that it was not possible to manage their 

rockburst problem with overhand cut-and-fill mining. Production mining at the Lucky Friday was 

stopped in April 1986, and plans were made to convert the entire mine to mechanized underhand 

cut-and-fill mining geometry, which they named LFUL – Lucky Friday underhand longwall. The 

key features of this mining method were that pillars would never be formed, and the mining would 

be carried out under a stable, engineered, paste type fill back as shown in Figure 4. 

Production mining at Lucky Friday resumed in October 1987 incorporating the above changes. 

Despite increased rates of rockbursting, as well as larger magnitude bursts (Ml 4.1), underhand cutand-fill mining at the Lucky Friday has been carried out without any serious rockburst injuries or 

fatalities, and with greatly increased productivity at significantly reduced costs. Underhand mining 

has allowed Hecla to very effectively manage their rockburst problem. The miners have a higher 

sense of security working below an engineered back. Management has said that the mine would 

likely have never reopened after 1986 had it not been for the all the benefits of LFUL mining. 

Finally, the paste backfill is only very rarely damaged by the effects of nearby rockbursts. The only 

burst induced fill failure at the mine occurred in 1991 during mining of a remnant pillar where a 

3.5Ml burst caused the wall to fail and in turn undercutting the past back which collapsed. The 

peak particle velocity at the hanging wall/fill mat was approximately 1m/s. Despite closure from 

ongoing mining, as well as closure and shock loading from the burst, the fill was not rubbilized as 

might have been expected. 

4 ONGOING RESEARCH 

The following is a summary of current work being completed at the University of British 

Columbia’s Norman B. Keevil Institute of Mining on paste backfill. 

4.1 Numerical Model Description 

A FLAC2D was constructed to represent a high angle ore body with competent host rock in which 

underhand cut and fill mining is being implemented. The model was constructed in such a way that 

span widths, span heights, rock properties, paste properties and wall closure can be specified. The 

model was set up with far field stress equal to a plane strain, σz = σv =0.027MPa/m depth, σx = σy = 

2/5(σv) based on gravity loading (Figure 5) with a 42o

 friction angle. The footwall and 
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

hanging wall have an angle of 70o. 

JOB TITLE : Vertical and Horizontal Stresses - Gravity Loading 

 FLAC (Version 5.00) 

 0.000 

(10 ) 01 

LEGEND 

27-Jul-06 13:21 

step 1874 -0.100

HISTORY PLOT

 Y-axis : -0.200 

 1 Ave. SYY ( 51, 21)

 2 Ave. SXX ( 51, 21)

 X-axis : -0.400 

-0.300 

Number of steps

-0.500 

-0.700 

-0.600 

-0.800 

-0.900 

2 4 6 8 10 12 14 16 18 

(10 ) 02 

UBC Geomechanics Center 

Figure 5: Plot of Horizontal and Vertical Stress used within FLAC2D model 

The rock properties were equivalent to a competent limestone with properties shown in Table 2. 

The cohesion and tensile strength was equivalent to 1⁄4 and 1/10 of the UCS (Jaeger et al., 1976) 

respectively and Elastic properties values were obtained from various limestone tests performed at 

the UBC Geomechanics Centre. The rock behaved according to a Mohr-Coulomb constitutive 

model. The amount of wall closure is user specified and has initially been tested for 0 and 10mm of 

wall closure. 

Table 2: Rock: Properties 

Properties Value 

UCS 22Mpa 

Friction Angle 42o 

Elastic Modulus 25GPa 

Poisson’s Ratio 0.22 

The paste fill properties were user defined and were typically based on values obtained in literature 

from Pakalnis (2005) and Caceres (2005). The elastic modulus was based on the UCS of the paste 

and was determined according to Eqn 2.. This equation serves only the purpose of allowing the 

elastic modulus to reflect a change in UCS during coding. It should be noted that Eqn 2. is a best fit 

line of limited test samples. 

⎧0.15 GPa UCS < 0.4 MPa 

E = ⎨ (2) − 371.6 ⋅UCS 2

⎩ +1317.9 ⋅UCS − 274.4 0.4 < UCS < 1.75MPa

The paste behaved according to a Mohr-Coulomb strain softening constitutive model with cohesion 

friction angle and tension decreasing in value as the strain increases as seen in Figure 5 a,b,c. 
 

Parameter Cut-offs 

0 

0.1 

0.2 

0.3 

0.4 

0.5 

0.6 

0.0% 0.2% 0.4% 0.6% 0.8% 1.0% 1.2% 

Percent Strain 

MPa 

0 

5 

10 

15 

20 

25 

30 

35 

40 

45 

Friction Angle 

Cohesion Tension Friction Angle 

FIGURE 6: (a)Cohesion vs. Strain cut-off; (b) Friction Angle vs. Strain cut-off (c) Tensile Strength vs. Strain 

cut-off 

The following geometries were modelled: 

• 3m Span with sill thickness of 3, 3.5, 4 and 23.5 m thickness 

• 6 m Span with sill thickness of 3, 3.5, 4 and 23.5 m thickness 

• 8 m span with sill thickness of 3, 3.5, 4 and 23.5 m thickness 

The air gaps of 0.5m were assigned between subsequent lifts with exception of the 23.5m thickness 

which is supposed to represent a near infinite span. Models were run with only the UCS of the 

paste being modified. This had a trickle down effect of altering the cohesion, tension and elastic 

modulus of the material. Span width and UCS paste combinations were marked as stable or 

unstable. An unstable scenario was considered a model with no numerical convergence or sill mat 

failure. 

4.2 Initial Findings 

Figure 7 shows the initial findings of the model for a span of 23.5 m height. Note that the wall 

closure consist of 5mm of closure on the hanging wall and footwall respectively. It can be seen that 

as the span width decreases the effect of wall closure affects the necessary strength of paste required 

for a stable span. This is due to the increase in strain within the paste and is considered a crushing 

failure. The zero convergence line follows a similar trend as the 10mm of wall closure line for large 

spans yet a deviation occurs as the spans become smaller than 4 meters. As the wall closure is mine 

sensitive and is independent of span width, it is more beneficial to real world applications to plot 

wall closure rather than strain within the paste. 
 

UCS vs. Span Width 

0 

0.2 

0.4 

0.6 

0.8 

1 

1.2 

0 2 4 6 8 10 12 14 Span Width (m) 

UCS (MPa) 

10mm Wall Closure 0 mm Wall Closure 

Figure 7: Span vs. Paste UCS 

4.3 Future Work 

Further research will be done with respect to Figure 6 to determine the effect of wall closure on the 

UCS of paste required for a safe working environment. This will consists of refining the mesh 

along the contacts of the paste with the host rock and by measuring total strain across the paste 

span. 

Currently the work is in the theoretical stage and is not calibrated to a specific mine. Future work 

would be to calibrate the model to a mine currently employing underhand paste fill. From this 

verification, an optimization of span width vs. UCS of paste can be derived. 

4 OTHER OBSERVATIONS 

A recent report from NIOSH (Tesarik et al., 2005) discussed the earth pressure cells are 

significantly affected by the paste cure temperature which can reach 40o C with stresses measured 

in excess of 69kPa due to the temperature difference. Figure 8 shows the uncorrected values and 

corrected values based on the temperature of the paste curing. The applied corrections are on top of 

the corrections supplied by the pressure cell manufacturers, the manufacturer’s correction factor 

applies only to the transducer and not to the entire earth pressure instrument body. By inspecting 

Figure 7, it can be seen that not only are the uncorrected stress greatly lower than the true stress but 

the uncorrected stress show a negative trend while the corrected stress show a positive increase of 

stress. The great difference in uncorrected stress vs. corrected stress is very important in the case of 

paste as the UCS of paste is very low and that without the proper correction an increase in loading 

causing failure could go undetected. Furthermore, without the correction over time, the pressure 

within the fill has a negative trend when in actual fact the fill is being loaded. 
 

Figure 8: Corrected vs. Uncorrected Earth Pressure Cell readings. Tsearik, (2005). 

Current work is underway on lateral loading mechanisms of fill fences. It is the goal of the research 

to understand the loading mechanisms, determine fill fence construction guidelines and understand 

paste/ fill fence interaction. The research will consist of field work inspecting current fences, 

placing instruments on fences to determine loading mechanisms and, hopefully, a destructive test of 

an instrumented fill fence. Once the field has been completed a model will be constructed based on 

recorded values and further analysis will be carried out. 

6 CONCLUSION 

Mining under consolidated fills is becoming competitive to conventional cut and fill mining as 

increased spans and productivities are realized through reduced placement of ground support and 

more control on the mine cycle due to working under an engineered back. This requires a thorough 

understanding of the mechanism of support that one is relying upon which is the consolidated fill 

immediately above. The fill may be supported in terms of conventional bolts and screen in order to 

counter “cold joints” that may develop in the fill, account for variability in fill quality control and/or 

increase the overall factor of safety required due to seismic events in the close proximity. This 

requires an understanding of the stabilization affect of the consolidated fill and the mine 

environment that it is placed within. Through the gathering of site data, modelling of behaviour 

either analytically and/or numerically coupled with observation and measurement one will be able 

to advance the overall design criteria to provide a safe and cost effective workplace. 
 

 

 

 

 

ACKNOWLEDGEMENTS 

The authors would like to thank NSERC, NIOSH and individual mining companies for providing 

funding of the graduate research that allows industry related research to be conducted. 

REFERENCES 

BLAKE, W. and HEDLEY, D.G.F. 2003. Rockbursts: Case Studies from North American Hard-Rock Mines. 

SME Publication, 119pp. 

BLIGHT, G. 1984. Soil Mechanic Principles in Underground Mining. Journal of Geotechnical Engineering, 

Vol. 110, No.5, 11pp. 

BRADY, T., PAKALNIS, R., CACERES and HUGHES. 2006. Design in Weak Rock Masses: North 

America Underground Mining Operations. AGM-CIM, Vancouver 

CACERES, C, 2005. Effect of Backfill on Open Stope Mining Methods, MASc Thesis, Mining Engineering, 

University of British Columbia, 139pp. 

Pakalnis R, Caceres C, Clapp K, Morin M, Brady T, Williams T, Blake W, MacLaughlin M 2005, Design 

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