1

Backfill

σv

Span

Open Stope

Sill Mat Pillar

- Cohesion

- Friction Angle

-Tensile Strength

- Density

h

• Vertical Loading

• Stope Geometry Influences

• Failure Modes

Seismic

σh

β - Modulus, Poisson Ratio

Design Requirements

• Lateral Loading

• Seismic Effects

• Support Implications

• Strength Properties

• Fill Placement 

• Other (Air Gaps/Cold Joints)

• Stope Closure

Backfill

σv

Span

Open Stope

Sill Mat Pillar

- Cohesion

- Friction Angle

-Tensile Strength

- Density

h

• Vertical Loading

• Stope Geometry Influences

• Failure Modes

Seismic

σh

β - Modulus, Poisson Ratio

Design Requirements

• Lateral Loading

• Seismic Effects

• Support Implications

• Strength Properties

• Fill Placement 

• Other (Air Gaps/Cold Joints)

• Stope Closure

Figure 1: Mining underneath consolidated backfill. 

Design Spans – Underhand Cut and Fill Mining 

Presented at 107th CIM-AGM Toronto, April 2005 

Rimas Pakalnis, Cristian Caceres, Kathryn Clapp, Mario Morin 

University of British Columbia 

Tom Brady, Ted Williams 

Spokane Research Laboratory, NIOSH 

Wilson Blake 

Consultant 

Mary MacLaughlin 

Montana Tech, University of Montana 

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 is generally necessary either 

due to a weak rock mass comprising the immediate back and/or high induced back stresses. A major concern in the 

design of sill mats is the loading and strengths associated with the overlying sill mat. This paper reviews past practice 

coupled with present observations and measurements from over ten(10) mines throughout North America. It outlines 

areas of concern in terms of design requirements. 

1. Introduction 

Backfilling in North America has been practised since the turn of the century. Souza et al. (2003) has summarized the 

advancements in backfill with the introduction of hydraulic fills in the 1950’s and the addition of cement in the 1960’s. 

This coupled with cemented rock and paste fills being introduced in the 1980’ and 1990’s respectively resulted in the 

implementation of mining methods that require extraction under a consolidated back largely comprised of fill rather than 

timbered mats/cables (Marcinyshyn, 1996). The increased use of consolidated fills in the late 1990’s to present under 

engineered conditions with a high degree of reproducibility in terms of strengths and predicted behaviours has enabled 

man-entry methods such as underhand cut and fill to be implemented under greater controlled spans resulting in a safe 

and economic alternative to 

conventional cut and fill mining, (Mah, 

2003). A database of twelve (12) 

underhand cut and fill operations was 

compiled as part of this project through 

mine visits. The operations were 

located throughout North America and 

summarized in Table 1. 

The placement of consolidated fill 

either cemented rock fill or paste 

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 
2

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. This sill may be comprised of one large vertical height placed by bulk mining as 

shown in Figure 2a or by single lifts as placed by conventional drift and fill and/or underhand cut and fill as shown in 

Figure 2b. The major difference is that in Figure 2a one is largely operating remotely from the immediate filled back, 

whereas in Figure 2b one is mining by man-entry methods. This necessitates that the factor of safety for the man entry be 

substantially greater than the non-entry approach. 

Figure 2: Mining under a consolidated back. 

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, numerical assessment of span and applied 

loading conditions. 

2.1 Design Load 

A critical factor is estimating the design loads onto the sill mat. A recently completed MASC thesis by Caceres 

(2005) employing the Musselwhite mine of Placer Dome as a case study had looked at the loading conditions that exist 

on a cemented rock fill mat as shown schematically in Figure 2a. Knowing the loads is critical to determining the 

strength required of the sill mat for the given stope geometry. Under-estimating can cause a premature failure of the sill 

mat once mining exposes the mat whereas overestimating can result in unnecessary expense due to the cost of the cement 

in place. Knowing 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 value for “k” describes the ratio between the horizontal and vertical stresses in the 

fill and indirectly the ability of the degree of load transfer by arching. When fill is placed initially, very little shear 

resistance is mobilized through grain interaction, the coefficient of earth pressure at rest (Ko). Subsequent placement of 

fill results in the fill mass to settle and compact, increasing the shear resistance and transfer load to the abutments 

through arching. As mining underneath the sill pillar progresses, a void is created toward which the fill mass will tend to 

move if unconsolidated with transfer of vertical stresses laterally through arching. This condition is described by the 

passive earth pressure coefficient (Kp) where full shear resistance is mobilized. This is analogous to classical 

embankment theory where the walls are moving into the fill. The effect of K employing individual fill load formulaes is 

shown in Figure 3a. The analytical methods shown in Figure 3a assume vertical stope walls which is generally a 

conservative estimate of vertical loading as shown by Caceres (2005). 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 as shown in Figure 3b. 

CRF Sill Mat 

Pillar 

Backfill 10m + 

10m + 

a) Longhole mining under a cemented rock 

fill sill mat (Caceres, 2005) b) Underhand cut and fill under 

a paste back. 

PREVIOUSLY 

BACKFILLED ~3m LIFT 
3

1

2

3(low)

4 8

9

10 3(hi)

5

6

7

7 7

7

8

9

10 11

12

ROCKFILL

HYDRAULIC

PASTE

12% 14%

1

2

3(low)

4 8

9

10 3(hi)

5

6

7

7 7

7

8

9

10 11

12

ROCKFILL

HYDRAULIC

PASTE

12% 14%

Figure 3: Estimation of vertical loading onto sill mat. 

Eqn. 1: 

The above was derived for cemented rock fill, however, the analytical 

solution would be similar to that of 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 affects, 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 

tools to attempt to provide an initial tool for the operator for design. 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. Figure 4 shows 

the compiled database (Table 1) of backfill unconfined compressive strengths adapted from Souza et al.(2003). 

Vertical Stress Comparison

Analytical vs. Numerical Modeling

0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

0.40

0.45

45 50 55 60 65 70 75 80 85 90

Stope Dip Angle (deg)

Vertical Stress σyy (MPa)

Analytical Equation

Numerical Modeling

a) Vertical stress versus stope span 

using analytical solutions from different 

methods. 

b) Analytical versus numerical derivation 

for different stope inclinations. 

Vertical Stress Comparison

0.000

0.100

0.200

0.300

0.400

0.500

0.600

0.700

0.800

0.900

4. 6. 8. 10. 12. 14. 16. 18.

Stope Span (m) 

Ve

rti

ca

l 

St

re

ss

σ

yy 

(M

Pa

)

Flac Janssen

Reimbert

Assuming:

- Density of 2.0 t/m3

- 90deg stope dip

- 37deg friction angle

- ν = 0.29

Mitchell

Blight

⎥

⎦

⎤ ⎢

⎣

⎡

⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⋅

⋅ ⋅ ⋅ ⋅ ⋅ − − ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛

⋅ ⋅

⋅ = sin ( )

2 tan( ) sin ( ) 1 exp 2 tan( ) ( ) 2

2

β

φ β φγ σ

L

K z

K

L

z y

Where:

− L = Span of the stope

− z = Height of rockfill

− K = Coefficient of lateral earth pressure ~ 1.4Sin2( φ) - 2Sin ( φ) + 1

− γ = Rockfill’s unit weight

− φ = Rockfill’s friction angle

− β = Stope dip angle

Figure 4: Backfill strength vs binder 

content (Table 1). 
4

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

DESIGN OF UNDERCUT SILL SPANS

MINIMUM SPAN WIDTH (m)

UNCONFINED COMPRESSIVE STRENGTH (MPa)

UNDERHAND CUT AND FILL OPERATIONS

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

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

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

DESIGN OF UNDERCUT SILL SPANS

MINIMUM SPAN WIDTH (m)

UNCONFINED COMPRESSIVE STRENGTH (MPa)

UNDERHAND CUT AND FILL OPERATIONS

β

Uncemented

rockfill

Open Stope

Sill Mat

Caving

Rotation

Shear Shear

σv

σt τ τ

L

σ d c

σc

β

Uncemented

rockfill

Open Stope

Sill Mat

Caving

Rotation

Shear Shear

σv

σt τ τ

L

d

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

L ⋅ γ > 8 ⋅σ t π σ γ

σ σ

+ ⋅

⋅ + ⎟ > ⎠

⎞ ⎜

⎝

⎛

d d

L

v

t c 2 ( ) 2

⎟

⎠

⎞ ⎜

⎝

⎛

⋅ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ + ⋅ > ⋅ L

d d t

v sin ( ) ( ) 2 2 β

τ σ γ ( cot( )) sin ( ) ( ) 2

2

β β

σ σ γ ⋅ − ⋅ ⋅⋅ + ⋅ > L L dd 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

L

σ d c

σc

β

Uncemented

rockfill

Open Stope

Sill Mat

Caving

Rotation

Shear Shear

σv

σt τ τ

L

d

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

L ⋅ γ > 8 ⋅σ t π σ γ

σ σ

+ ⋅

⋅ + ⎟ > ⎠

⎞ ⎜

⎝

⎛

d d

L

v

t c 2 ( ) 2

⎟

⎠

⎞ ⎜

⎝

⎛

⋅ ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ + ⋅ > ⋅ L

d d t

v sin ( ) ( ) 2 2 β

τ σ γ ( cot( )) sin ( ) ( ) 2

2

β β

σ σ γ ⋅ − ⋅ ⋅⋅ + ⋅ > L L dd 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.

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 5. 

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 rock fills that crushing, 

caving, sliding are generally negated 

when the sill thickness exceeds 0.5 x 

span , absence of closure stresses and 

the unconfined compressive strength 

of the cemented rock fill is greater 

than 1.5MPa and that rotational 

instability where kinematically 

possible has to be analysed 

separately. Figure 6 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. 

The limit equilibrium approach shown in Figure 5 was simulated by Caceres(2005) employing FLAC2D models (finite 

difference code) for a given value of cohesion, span and stope dip. The cemented rock fill properties assigned are for a 

Mohr-Coulomb type of material with strain-softening behaviour where integrity is lost after 1.5% strain (Swan and 

Brummer, 2001). The resultant mode of failure was analysed for 90o

 to 75o dip stopes with cohesion on the hanging wall 

Figure 6: Stability chart for the design of undercut sills with vertical 

sidewalls with a FS of 2. Chart is based upon fixed beam bending failure 

Figure 5: Limit equilibrium criteria adapted from Mitchell, 1991. 
5

0.00 0.00 0.00 0.00

0.0

2.0

4.0

6.0

8.0

10.0

12.0

14.0

0.75 1.00 1.25 1.50 1.75 2.00 2.25 2.50 2.75

Cohesion (MPa)

Minimum Sill Mat Vertical Height (m)

FAIL

STABLE

6m

8m

10m

12m

S

S

S

S

S

S

S

S

S

S

Rc Rc Rc

Rc Rc

Rc

HW contact 

cohesion value

2b

4c 8

a)Sill mat stability for 75o – No strength on HW (τt = 0%)

σv

h

Span

τt=0%

75o

σv

h

Span

τt=0%

75o

0.50 0.75 1.00 1.25

0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

8.0

9.0

10.0

0.75 1.00 1.25 1.50 1.75 2.00 2.25 2.50 2.75

Cohesion (MPa)

Minimum Sill Mat Vertical Height (m)

FAIL

STABLE

6m

8m

10m

12m

Rc Rc Rc Rc

R Rc Rc

Rc Rc Rc

S Rc

S

S

S

R

HW contact 

cohesion value

2b

4c 8

σv

h

Span

τt=50%

75o

σv

h

Span

τt=50%

75o

b) Sill mat stability for 75o – 50% sill mat strength on HW (τt = 50%)

at different Stope Spans

1.00 1.50 2.00 2.50

0.0

1.0

2.0

3.0

4.0

5.0

6.0

0.75 1.00 1.25 1.50 1.75 2.00 2.25 2.50 2.75

Cohesion (MPa)

Minimum Sill Mat Vertical Height 

(m)

FAIL

STABLE

6m

8m

10m

12m

Rb

Rb Rb Rb

Rb Rb

Rb

Rb Rc

Rc Rc

Rc

S

Rc

Rc

Rc

HW contact 

cohesion value

2b

4c 8

c) Sill mat stability for 75o – 100% sill mat strength on HW (τt = 100%)

σv

h

Span

τt=100

%

75o

σv

h

Span

τt=100

%

75o

contact varying from zero to maximum (cemented rock fill cohesion) as shown in Figure 7 for the 75o stope dip. The 

analytical approach after Mitchell assumes no hanging wall cohesion for the rotational instability and this was found to 

result in a high degree of conservatism. The FLAC2D simulation shows the failure mode that results for a given stope 

span, sill height and cohesive strength for various assumptions on wall friction. The design curves have been developed 

for cemented rock fill operations, however, they are largely dependent upon material strength characteristics which are 

input into the model. This method 

does not have the constraints 

associated with the limit equilibrium 

techniques shown in Figure 5 as the 

relationships have been derived 

through numerical simulation of 

modeled conditions. The critical 

parameters identified in Figure 7 do 

not incorporate a factor of safety as 

they are identified in the numerical 

model by being either stable or 

unstable. Therefore one must apply a 

safety factor on the input parameters 

to ensure an adequate design factor of 

safety is incorporated. The cemented 

rock fill database shown in Table 1 

was overlain onto Figure 7. The type 

of failure occurring is indicated for 

each curve by either “S” for sliding, 

“Rc” for rotational-crushing, “Rb” 

for rotational-breaking, and “F” for 

flexural failure modes. The degree of 

cohesion at the wall contact allows 

flexibility in design depending upon 

the quality of fill placement. It must 

be recognized that based upon the 

cases identified in Table 1 and plotted 

onto Figure 7 that the depth of mat 

largely corresponds to 50%-100% 

cohesion on the hangingwall contact. 

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, 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). 

3.0 Seismic Case History – Managing Rockbursts 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). 

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 

Figure 7: Sill mat stability as derived from numerical modeling. Factor of 

Safety >1.0 
6

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 a “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 rockburst 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. 

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 cut-and-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 hangingwall/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. 

4.0 Other Observations 

The focus of the present research is towards mining under paste, however, a much larger database of information exists 

for mining under cemented rock fill as derived from the Nevada database of mines operating within a weak rock mass 

(Brady et al., 2005) and the database shown in Table 1 reflects this. The limit equilibrium methods are similar whether 

one is working with a cemented rock beam or a consolidated paste. The differences lie in the resultant strengths 

associated with each as shown in Figure 4 where the cemented rock fills exhibit strengths (UCS) generally in excess of 

two to three times that of the cemented pastes. This is largely due to cement being able to be more evenly mixed with the 

larger aggregate as compared to the paste which is generally between (60% passing) 20 microns (fine) to 100 microns 

(coarse) for Mine #1 and Mine #12 in Table 1. The typical cemented rock fill optimum aggregate size is 50 000 microns 

or 5cm (2inch) for the database analysed. 

The cohesion for paste fills was estimated as being 0.25 times the unconfined compressive strength based upon an 

internal angle of friction (f )of 30o

 and derived from the Mohr-Coulomb relationship where the “Unconfined 

compressive strength = 2*Cohesion* (cos f) /(1-sin f)”. A value of internal angle of friction of 35-40o

 was employed 

for the rock fills. 

Figure 8: Lucky Friday Mat 

Chainlink Fencing over 

plates 

Tie Wire attached to 

Dywidag bolt and to 

wall

6ft Dywidag

Bolt (typ) 

6” Prep muck 

maximum

Chainlink Fencing over 

plates 

Tie Wire attached to 

Dywidag bolt and to 

wall

6ft Dywidag

Bolt (typ) 

6” Prep muck 

maximum

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

Chainlink Fencing over 

plates 

Tie Wire attached to 

Dywidag bolt and to 

wall

6ft Dywidag

Bolt (typ) 

6” Prep muck 

maximum

Chainlink Fencing over 

plates 

Tie Wire attached to 

Dywidag bolt and to 

wall

6ft Dywidag

Bolt (typ) 

6” Prep muck 

maximum

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
7

The tensile strength was generally derived from the unconfined compressive strength for consolidated fills and a value of 

“0.1 x UCS =Tensile Strength” was employed (Jaeger et al., 1976). 

A Young’s modulus for paste ranged from 0.6GPa 

to 3MPa (laboratory) for 10% binder and these 

values were field calibrated (Williams et al., 2004) 

to reflect field data through a combination of earth 

pressure cells embedded within a paste stope at the 

Lucky Friday mine (Mine #12) and the closure 

recorded and related by the relationship “Stress = 

Modulus x Strain” where the strain was measured 

by means of closure meters divided by the stope 

width. This is shown in Figure 9 with the resultant 

closure/load history. It is interesting to note that 

the loads upon reaching 4MPa (UCS) showed 

yielding of the paste. The closure at the Lucky 

Friday was in excess of 25mm. The rock mass of 

the wall contact was under 50% (RMR76). Mine #1 

showed closure values of under 10mm at similar 

depths, with the wall contact having an RMR76 = 

75%. Both sites are considered to be burst prone. 

A further observation by Tesarik et al., 2005 was 

that 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. The vertical pressures that arise 

for a 3m pour height of paste fill would approach fill pressures of 0.02MPa/m of pour height (SG=2.0) or 60kPa of 

vertical pressure which is largely equivalent to the temperature correction on the earth pressure cell. The horizontal 

pressure would be a fraction of this. It is important to recognize the correction factor can be equivalent to the absolute 

value measured. 

5.0 Conclusions 

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. 

6.0 Acknowledgements 

The authors would like to thank Goldcorp, Barrick, Newmont, Placer Dome, Cameco, Hecla, Coeur d’Alene Mines, 

Stillwater and NIOSH for assisting in the overall project and providing direction. Particular thanks to NSERC, NIOSH 

and individual mining companies for providing funding of the graduate research that allows industry related research to 

be conducted. 

7.0 List of References 

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

Publication, 119pp. 

Figure 9: Instrumented underhand stope at Lucky Friday 

(Williams et al., 2001)

CLOSURE METER

a) Vertical section through instrumented 

stope at Lucky Friday Mine.

AIR GAP ~0.3m

EARTH PRESSURE CELL

b) Resultant load/closure 

measurements.

CLOSURE METER

a) Vertical section through instrumented 

stope at Lucky Friday Mine.

AIR GAP ~0.3m

EARTH PRESSURE CELL

b) Resultant load/closure 

measurements.

a) Vertical section through instrumented 

stope at Lucky Friday Mine.

AIR GAP ~0.3m

EARTH PRESSURE CELL

b) Resultant load/closure 

measurements.
8

BLIGHT, G. 1984. Soil Mechanic Principles in Underground Mining. Journal of Geotechnical Engineering, Vol. 110, 

No.5, 11pp. 

BRADY, T., PAKALNIS, R. and Clark, L. 2005. Design in Weak Rock Masses: Nevada Underground Mining 

Operations. SME-AGM, Salt Lake City, 9pp. 

CACERES, C. 2005. Effect of Backfill on Longhole Open Stoping, MASc Thesis, Mining Engineering, University of 

British Columbia, 139pp. 

GRAHAM S. and BRUMMER, R.K. 2001. Backfill Design for Deep, Underhand Drift-and-Fill Mining. Proceedings of 

the 7th International Symposium of Mining with Backfill. USA, 12pp. 

ITASCA, 2005. Software - FLAC (Fast Lagrangian Analysis of Continua), Minneapolis, Minnesota. 

JAEGAR, J.C. and COOK N.G.W. 1976. Fundamentals of Rock Mechanics. Chapman and Hall Publishers, p190-191. 

JANSSEN, H.A. 1895. Versuche uber getreidedruck in silozellen. Zeitschrift des Vereines Deutscher Ingenieure, Berlin, 

Germany, 39(35), 1045–1049. 

JORDAN, J., LANGSTON, R.KIRSTEN, H., MARJERISON, J, JACOBS, C., and STAHLBUSH, F. 2003. Underhand 

Cut and Fill Mining at the Stillwater Mine, 105th AGM-CIM, Montreal, 18pp. 

MAH, P., TESSIER,M. and CLELLAND, J. 2003. Mining Methods at Goldcorp’s Red Lake Mine, 16th CIM Mine 

Operators’ Conference, Saskatoon, 22pp. 

MARCINYSHYN, K. 1996. Sill Mat Design for Narrow Vein Underground Mines. MASc Thesis, Mining Engineering, 

University of British Columbia, 151pp. 

MITCHELL, R.J., 1991. Sill Mat Evaluation Using Centrifuge Models . Mining Science and Technology, Elsevier 

Science Publisher B.V., 13: pp. 301-313. 

REIMBERT, M. & A., 1976. SILOS Theory and Practice, Trans Tech Publications, Vol. 1, No 3, p. 250. 

SOUZA, E.De, ARCHIBALD, J.F., and DIRIGE,A.P. 2003. Economics and Perspectives of Underground Backfill 

Practices in Canadian Mines. 105th AGM-CIM, Montreal, 15pp. 

STONE, D.M.R. 1993. The Optimization of Mix Designs for Cemented Rockfill. Minefill 93. SAIMM, Johannesburg, 

pp. 249-253. 

TERAZGHI, K., PECK,R.B., and MESRI.,G. 1996. Soil Mechanics in Engineering Practice. Wiley and Sons, New 

York, 549pp. 

TESARIK, D.R., SEYMOUR, J.B., WILLIAMS, T.J., MARTIN, L.A. and JONES, F.M. 2005. Temperature Corrections 

to Earth Pressure Cells Embedded in Cemented Backfill, RI in Progress, 59pp. 

WILLIAMS, T.J., DENTON, D.K., LARSON, M.K., RAINS,R.L., SEYMOUR, J.B. and TESERIK, D.R. 

Geomechanics of Reinforced Cemented Backfill in an Underhand Stope at the Lucky Friday Mine. RI 9655, NIOSH 

(SRL), 18pp. 
9

SPAN SILL THICKNESS UCS COMMENTS

(m) (m) (MPa)

1 10 6.1 3 2 PASTE

2a 6.5 7.6 4.6 5.5 CRF

2b 8% 9.1 4.6 6.9 CRF Design

2c 8% 21 4.6 6.9 Mined Remote - No Cave

2" Minus Aggregate

Go Under a Minimum of 14days

Wall CRF 5-6% Binder

Jam Tight to Back/Steep

Go Under in 14Days(9MPa UCS)

3 7 3 3 4-11 CRF (4MPa Design)

UCS is 11MPa (28 DAY)

4a 9 13.7 4 8.3 CRF Test Panel

4b 9 3.7 3 8.3 CRF Drift & Fill

4c 9 7.3 3 8.3 CRF Panel

5 7 2.7 3 3.4 CRF

6 6.75 4.9 4.3 4.8 CRF

Go Under in 28 Days

 0.7

7a 10 1.8 2.7 0.3 PASTE (FS=1.5)

7b 2.4 2.7 0.5 Go Under in 7 Days - 28 Days

7c 3 2.7 0.7

7d 3.7 2.7 1 (5% binder - 0.5MPa UCS 28D)

7e 4.3 2.7 1.4 (7% binder - 0.7Mpa UCS 28D)

7f 4.9 2.7 1.8 (10% binder - 1MPa UCS 28D)

7g 5.5 2.7 2.3 (12% binder - 1.2MPa UCS 28D)

7h 6.1 2.7 2.9

8 7 4.6-6.1 4.6 5.5 CRF

9 10 5 5 4.45 CRF

10 12.8 6-9 6 2 High Density Slurry

(78% WT SOLIDS)

10% Cemented Hydraulic Fill

Go Under in 7days (2.5MPa UCS)

11 10 3 3 2.5 (73-75% Wt Solids)

(includes 0.9m air gap) (UCS after 7 days)

Go Under in 3days (2.4MPa UCS)

12 8 2.4-4.6 3 4.8 8% Paste (Coarse Tails)

(includes 0.6m air gap) (no free water)

(2% binder - 0.2MPa UCS 28D)

13 2-10% 0.2-2 (5% binder - 0.8Mpa UCS 28D)

(7% binder - 1.2MPa UCS 28D)

(10% binder - 2MPa UCS 28D)

MINE %BINDER

UNDERHAND CUT AND FILL MINING UNDER CEMENTED FILL

NO UNDERHAND AT THIS TIME

Table 1: Database of Underhand Cut and Fill Operations