1 

LONGWALL SHIELD AND 

STANDING GATEROAD 

SUPPORT DESIGNS – 

IS BIGGER BETTER?

Thomas M. Barczak, PhD 

Senior Research Engineer 

National Institute for Occupational Safety and Health 

Pittsburgh Research Laboratory 

Pittsburgh, PA 

Stephen C. Tadolini, PhD 

Chief, Rock Safety Engineering Branch 

National Institute for Occupational Safety and Health 

Pittsburgh Research Laboratory 

Pittsburgh, PA
2 

ABSTRACT 

Roof support systems, longwall shields in particular, have persistently grown in size and capacity 

for the past 25 years. Why is this? Is it a good thing or is it unnecessary? What are the 

consequences of this “bigger-the-better” design philosophy? Rarely do we see a “one size fits 

all” approach to roof support except when it comes to longwall shields, where the historical trend 

for new procurements is to simply purchase the highest capacity support you can buy. Does high 

capacity ensure failures will not occur? The answer is no! The highest capacity tailgate supports 

are typically the ones that fail the most often. What does this tell us? It tells us that there are 

fundamental aspects of support design and ground control that we still do not understand or 

recognize when designing a support system. This paper provides an interesting evaluation of 

these issues and proposes a different approach to support design that attempts to match the 

support performance characteristics to the ground behavior. Along the way, some of the old 

premises of support design will be discarded, and new “controversial” ideas developed. 

Thinking outside the box never hurt any industry. Mining has made it a common practice. 

Introduction 

The debate about whether bigger is better obviously refers to the support capacity not the size of 

the support, although in some cases the two are related. The debate is really about the degree of 

control that manmade supports have on ground behavior in underground mining. The key issue 

is whether the support system causes the ground to act in force or displacement control. The 

reason this is so important is that they have opposite effects in terms of support design 

requirements. In a force-controlled load environment, the goal is to provide enough support 

capacity to offset the rock load and provide equilibrium of the rock mass. Conversely, in a 

displacement-controlled load environment, the ground movement is not influenced or completely 

controlled by the support. The goal is to maintain support stability and roof support capability as 

convergence inevitably continues. 

With this background, this paper will evaluate the “bigger-is-better” design philosophy for 

longwall shields and standing roof support systems. The conventional support design approach 

based on simplistic models of supporting the full dead weight of detached rock masses is 

replaced by a ground reaction design approach. Here, the goal is to match the support 

characteristics to the ground response, and not to try and overpower the ground forces with some 

massive support capability. The ground reaction concept embodies both the force and 

displacement controlled loading aspects, and therefore provides a more accurate representation 

of the support loading requirements. 

Does anyone think that shield capacities can continue to increase forever? Obviously, they 

cannot. So at some point, more rationale design must prevail. Although standing supports have 

grown in capacities as well during the past 15 years, higher capacity standing supports can 

certainly still be designed. However, the highest capacity supports are typically the ones that fail 

prematurely. What can we learn from these failures? 

There are also design issues that are often a derivative of the pursuit of higher capacity support 

systems. For example, the desire to provide higher setting pressures is one of the reasons given 
3 

for wanting to have higher shield capacities. The impact of these issues will also be addressed 

and here too some surprises are revealed. 

Key Points in Current Shield Design 

The basic shield design has changed very little since its inception in the 1970's. The primary 

change has been a steady increase in shield capacity as shown in figure 1, resulting in larger 

hydraulic cylinders and support structures (U.S. Longwall Census, 2005). A few key points in 

shield design and behavior will facilitate the ensuing discussion on shield design issues. 

• Capacity increased by increasing leg diameter – The approach used to increase shield 

capacity has always been to increase the diameter of the leg cylinder while maintaining 

the same yield pressure, which limits the load carrying capacity of the shield. 

• Shield stiffness increases in direct proportion to the capacity – The approach used to 

increase shield capacity by increasing the leg diameter also results in a direct increase in 

shield (vertical) stiffness. This means that a higher capacity shield will develop more 

loading for the same roof-to-floor convergence than will a lower capacity shield. The 

caving shield-lemniscate assembly has no vertical stiffness and does not contribute 

significantly to the rated shield capacity. The stiffness of the shield is a function of the 

stage extensions (all current shields are two-stage designs). In general, the stiffness of 

the shield reduces as the mining height increases due to the cylinder extension.

• Setting pressure does not increase shield stiffness – The setting pressure provides an 

active force to the mine roof and floor, but does not change the stiffness of the shield. 

Figure 1 - Historical overview showing trend of increasing shield capacities during past 20 years.

500

600

700

800

900

1,000

1,100

1,200

1,300

1,400

1984 1986 1988 1990 1992 1994 1996 1998 2000 2002 2004 2006 2008

YEAR

SHIELD CAPACITY, tons

AVG SUPPORT CAPACITY

MAX SUPPORT CAPACITY
4 

Hence, regardless of what pressure the shield is set at, its reaction to convergence in 

terms of load development will always be the same at least for the same operating height 

and leg extensions.

• Canopy and base pressure distribution is not uniform – Although this assumption is 

often made in analyses of shield performance and used to justify higher capacity needs 

for larger supports, the pressure distribution is far from uniform. On the canopy, the 

pressure is focused near the leg connection, with very little and sometimes no significant 

pressure at the forward tip of the canopy. Conversely, the kinematics of two-leg shield 

designs tend to focus the pressure near the toe of the base. 

• Yielding lets fluid out of the leg causing it to lower – Yielding is accomplished by 

releasing hydraulic fluid from the leg cylinder, which causes the leg to lower a slight 

amount during each yield event. The resetting pressure of the yield valve will determine 

how much fluid is lost and how much closure will occur. Most yield valves have a 

resetting pressure of 90 pct or higher, meaning the pressure will decrease by 10 pct before 

the yield valve closes. 

The Ground Reaction Concept 

The ground reaction concept is a method that relates the support pressure to the convergence of 

the mine opening. The goal of any support design is to achieve equilibrium of the rock mass. 

The question is always the same, how much support is needed to accomplish this objective? The 

ground reaction concept can answer this question. The ground reaction concept is not a new 

idea. It has been around for several decades in the tunneling industry (Brown et al., 1983). The 

problem with the ground reaction concept is not that it's not useful; the problem is how you 

develop the curve. While in theory, it can be done by taking measurements of support load and 

convergence in the mine, in practice it is difficult to do because the support pressure must be 

varied to get the curve. For longwall shields, the curve (at least part of it) can be determined by 

lowering the setting pressures in a controlled series of steps and calculating the convergence 

from the resultant shield loading assuming the stiffness of the shield is known. 

Advancements made in numerical modeling are providing new opportunities to develop ground 

reaction information. The key issue is that the modeling is no longer constrained by elastic 

analysis and homogenous materials. Finite difference software such as FLAC (Itasca, 2005) can 

be used to develop meaningful ground response curves. The software can be used to realistically 

model rock behavior from the initial elastic response to the large displacements and deformations 

that are associated with rock failure. It has the capability to model strength anisotropy found in 

the bedded coal measures and can simulate strain related weakening of failed rock. The software 

also has a built-in programming language which allows the user to control loads and 

displacements in the model. By applying internal pressure to the entry excavation, such as a 

longwall face or tailgate entry, the ground response curve can be determined (Esterhuizen and 

Barczak, 2006). Once this is done, more appropriate design criteria for roof support can be 

developed. 
5 

The ground response curve plots the support pressure against the convergence, as shown 

conceptually in Figure 2. Prior to excavation, the excavation boundaries are subject to pressure 

equal to the field stresses (point A). After the excavation is created the boundaries converge and 

the pressure required to prevent further convergence reduces as arching and the self-supporting 

capacity of the ground develops (point B). A point is reached (point C) where loosening and 

failure of the rock occurs and the required support resistance begins to increase as selfsupporting capacity is lost and support of the dead weight of the failed ground is required 

(point D). 

The effect of the support system can also be plotted on figure 2. Equilibrium is achieved when 

the support curve intersects the ground reaction curve (point B). Using two hypothetical support 

responses as shown in figure 2, two key points can be made regarding support design. The 

second support has a higher ultimate capacity (point E) than the first support (point F), but both 

reach the ground reaction curve at the same spot. This shows that higher capacity does not 

necessarily ensure better ground control. It also shows that from the perspective of the available 

support capacity, there is a component of ground response (convergence) that cannot be 

prevented by the support system. Hence, a degree of “uncontrollable convergence” exists, and 

this must be accounted for in the support design. 

Applying the Ground Reaction Concept to Shield Design 

 Can the ground reaction concept be developed and applied to shield supports? Absolutely. It 

has been widely accepted that increased setting pressures provide improved ground control by 

reducing face convergence. Field measurements of setting pressures when plotted against 

Figure 2 - Conceptual illustration of ground reaction curve and support response. 

CONVERGENCE

SUPPORT PRESSURE

A

C

D B

E

F

P

Ground reaction 

curve

Support #1 curve

Support #2 curve

CONVERGENCE

SUPPORT PRESSURE

A

C

D B

E

F

P

Ground reaction 

curve

Support #1 curve

Support #2 curve
6 

convergence produce a trend curve similar to that shown in figure 3. This is a ground reaction 

curve. As seen from the curve, higher support pressure results in reduced convergence, but it 

also shows that an optimum setting pressure exists due to the rapidly diminishing impact of 

support pressure beyond the inflection point in the curve in figure 3. It is also evident from this 

graph that the response to the left of the inflection point is a clear indication that one is 

approaching uncontrollable convergence produced by the main roof activity, which is well 

beyond the capability of even modern day shield supports. 

Medhurst has incorporated the concept of the ground reaction curve into shield design analysis 

(Medhurst, 2005). As Medhurst correctly points out, the ground reaction curve is dynamic in the 

sense that it changes with changing conditions. For example, the curve moves to the right for 

increasing extraction heights as shown in figure 4, indicating the higher capacity shields will be 

needed to accommodate the increase in face convergence associated with the reduction in coal 

seam stiffness at the higher extraction height. Deeper cover will cause higher abutment loading, 

which increases the scale of the support resistance axis and essentially causes the ground reaction 

curve to shift to the right, again implying higher shield capacities are required. Periodic caving 

would do a similar thing; push the curve to the right and upward, causing higher shield loading. 

Therefore, it is important to realize that the ground reaction concept essentially consists of a 

family of curves, each representing a unique loading condition. 

Figure 3 – Impact of shield capacities (setting pressures) on convergence. 

SHIELD SUPPORT PRESSURES

CONVERGENCE

Phigh

(A) – Inflection point

Plow

Clow Chigh

PA

CA

UNCONTROLLABLE 

CONVERGENCE

PYield

SHIELD SUPPORT PRESSURES

CONVERGENCE

Phigh

(A) – Inflection point

Plow

Clow Chigh

PA

CA

UNCONTROLLABLE 

CONVERGENCE

PYield
7 

Examining Individual Loading Cycles from a Ground Reaction Perspective 

Locally, the ground response is also affected by the shearer cut and adjacent shield moves. If 

you consider a three-shield environment, the ground reaction curve might look like something 

shown in figure 5a. When a shield is lowered, the support pressure in that particular shield drops 

to zero. While this hardly affects the support resistance across the entire face, it does lower the 

support resistance locally, i.e. over a three-shield area. As this particular shield is reset against 

the mine roof (point A on the first curve), the setting pressure increases from zero at roof contact 

to the full setting pressure, normally the pump pressure, less frictional losses in the delivery 

system. As the setting pressure is increasing from the initial roof contact, the load path is up the 

first curve, indicating that the convergence is actually decreasing during the setting operation. 

Since the ground reaction curve is steep, the decrease in convergence is relatively small, but 

measurable nonetheless (point B on the curve). Physically, the support pressure is compacting 

debris on top of the canopy or under the shield base or compacting loose rock debris along 

bedding planes that have been damaged from shear stress along these laminations resulting in 

some bulking or dilation of rock material, or simply closing openings within the roof structure. 

At this point, the ground reaction curve will remain static until there is a change in load 

conditions. Since the adjacent shield is also being reset in short order, the ground reaction curve 

shifts to the right from this activity and the loading on the first shield increases in proportion to 

Figure 4 - Impact of depth of cover and extraction height on longwall ground reaction curve (from 

Medhurst, 2005). 

Convergence (in)

Onset of roof failure

Cutting height

90% 

set

Support Density Before the Cut (ton/ft2)

14

12

11

10

9

8

7

6

13

0 0.8 1.2 1.6 2.0 2.4 0.4

12.5 ft

13.8 ft

14.8 ft

12.5 ft

13.8 ft

14.8 ft

Depth = 650ft Depth = 820ft

Convergence (in)

Onset of roof failure

Cutting height

90% 

set

Support Density Before the Cut (ton/ft2)

14

12

11

10

9

8

7

6

13

0 0.8 1.2 1.6 2.0 2.4 0.4

12.5 ft

13.8 ft

14.8 ft

12.5 ft

13.8 ft

14.8 ft

Depth = 650ft Depth = 820ft
8 

the shield stiffness (point C). The conditions remain nearly static once the shearer leaves the 

vicinity of the local area. Then, as the shearer reapproaches the local area and takes a cut from 

the face, the ground reaction curve again shifts to the right and the first shield again sees an 

increase in loading in proportion to its stiffness (point D). Release of the other adjacent shield 

again causes a slight shift in the ground reaction curve, again increasing pressure on the 

monitored shield (point E). The support is now at peak loading for the cycle, and as it is 

lowered, it follows down the ground reaction curve that it is currently located on until it loses 

contact with the mine roof. At this point (F), the shield is advanced and reset and the conditions 

are restored to the initial ground reaction curve (point A). 

The individual shield will then follow this cyclic load path from cycle to cycle 

(A→B→C→D→E→F). The ground will be stable as long as the shield has a load path (ground 

reaction curve) to follow that can complete the load cycle. If the shield fails to provide adequate 

support resistance to stay on a ground reaction curve that can complete the cycle as illustrated in 

figure 5b, equilibrium of the ground is not achieved and unstable conditions are likely to prevail. 

 This could be caused by inadequate set pressures, leaking leg cylinders, or dramatic shifts in the 

ground reaction curve by such things as periodic weighting induced by the intermediate roof 

structure. 

Figure 5a - Cycle loading pattern on ground reaction curve for individual or small group of shields.

CONVERGENCE

SHIELD LOAD

Support 

advance

Support 

lowering

Support 

loading

Support 

setting

Condition at 

setting

Shearer cut 

GRC Curve

Adjacent shield 

advance GRC Curve

Adjacent shield 

advance GRC Curve

A

B

E

F

C

D

A

B

E

F

+

+

-

-

C

D

CONVERGENCE

SHIELD LOAD

Support 

advance

Support 

lowering

Support 

loading

Support 

setting

Condition at 

setting

Shearer cut 

GRC Curve

Adjacent shield 

advance GRC Curve

Adjacent shield 

advance GRC Curve

A

B

E

F

C

D

A

B

E

F

+

+

-

-

C

D
9 

Comparison of the Ground Reaction Concept to Other 

Models Used for Shield Design 

Wilson was the first to promote the detached block-loading concept back in the early 1970's 

(Wilson, 1975). A caving zone occurs when a portion of the immediate roof is fractured by the 

overburden pressure, causing that portion to separate from the overlying rock mass and the full 

gravity weight of this block must be carried by the support (see figure 6). The size and therefore 

weight of the block that determined the required shield capacity was dependent upon the bulking 

factor of the immediate roof rock. This is a strictly force-controlled loading concept and as such 

does not fit well within the confines of the ground reaction concept. Early on when longwall 

mining first began in the U.S. with powered supports, it was concluded that the more competent 

strata found in U.S. mines caved in a more blocky fashion with less of a bulking factor than the 

weaker, more friable roof geologies found in most European mines. This immediately prompted 

a demand for higher shield capacities for the U.S. market. However, the requirement did not 

stop there, and shield capacities continued to grow beyond what bulking factors would justify. 

One reason was that the shields were also getting bigger (at first mostly longer and then wider as 

well). Hence, as the shields got bigger (longer and wider), they would need more capacity to 

carry the weight of the detached roof block. 

CONVERGENCE

SHIELD LOAD

Condition at 

setting

Shearer cut 

GRC Curve

Adjacent shield 

advance GRC Curve

Adjacent shield 

advance GRC Curve

A

B

E

F

C

D

Equilibrium not established –

ground conditions unstable

+

-

- CONVERGENCE +

SHIELD LOAD

Condition at 

setting

Shearer cut 

GRC Curve

Adjacent shield 

advance GRC Curve

Adjacent shield 

advance GRC Curve

A

B

E

F

C

D

Equilibrium not established –

ground conditions unstable

+

-

- +

Support 

lowering

Support 

loading

A

B C D

E

F

Support 

setting

Support 

lowering

Support 

loading

A

B C D

E

F

Support 

setting

Figure 5b - Equilibrium is not established in this shield loading cycle due to inadequate shield setting 

force. 
10 

A more realistic model of longwall rock mechanics is the stiffness model promoted by Smart 

(Smart, 1986). Smart suggested that the powered support was one of three yielding foundations, 

which support the immediate cantilevered roof and bridging main roof strata. The other two 

elements being the coal ahead of the face and waste (gob) behind the face (figure 7). The 

process, as Smart saw it, was that the powered support requirement is to control the subsidence 

(convergence) of the roof strata along the longwall face sufficiently to prevent fractures, which 

are generated ahead of the face line in the upper regions of the immediate roof, from migrating 

downward to the roof horizon at the longwall face. This is necessary to prevent the formation of 

the detached block condition promoted by Wilson. The key to Smart’s concept is that the 

stiffness of the coal largely controls the shield loading and design requirements. In particular, 

Smart hypothesized that the reason U.S. supports were required to be higher capacity than their 

European counter parts was that the more competent immediate roof structures failed in what he 

called “controlled parting-plane-caving” as opposed to the bulking-factor-caving of the weaker 

European geologies. The result being that the gob material provided very little support and 

created an even larger cantilevered roof beam that must be supported by the longwall shields. 

This model fits nicely with the ground reaction concept for shield design. First, it uses 

convergence as an outcome measure of the ground response based on the stiffness of the various 

supporting elements. Moreover, it allows for a component of uncontrollable convergence from 

the main roof loading relative to practical shield capacities. 

Figure 6 - Detached block model of longwall roof behavior. 
11 

Implications Toward Shield Design Requirements 

Stiffness Considerations 

The ideal situation would be one in which the mine roof does not know that the coal has been 

mined out from beneath it. In the global scale of the full panel, this is obviously quite impossible 

since the single row of shield supports would have to completely offset the full panel load of the 

overburden to prevent caving. On the more local scale of the supported advancing longwall face, 

the goal would be that the roof does not know the coal is mined out from beneath it at least to the 

point of the cave behind the shield line. To achieve this goal, the support design requirement 

should be to match the shield stiffness to that of the coal seam stiffness to provide a balanced 

stress distribution in the face area. Some researchers suggest that the reason why modern high 

capacity shields provide better ground control is that the increased stiffness of the higher 

capacity shields more closely matches that of the coalface structural stiffness. This would be one 

justification for increasing shield capacity under the premise that lower capacity shields were 

considerably less stiff than the coal structure, such that there was a large imbalance in stress 

acting on the immediate roof due to the degraded shield capacity compared to that provided by 

coal structure. 

Evidence to support this claim is provided by Medhurst (2005). Assuming that the coal seam 

yields at about 0.5% strain (based on laboratory testing of coal samples), the required shield 

stiffness can be determined. With similar stiffness, this will help ensure a uniform vertical 

compression profile and minimize mining-induced shear stresses in the immediate roof. For 

Figure 7 - Concept for evaluating roof shield response based on the stiffness of the shield, coal, and 

gob foundation. 
12 

example, a 10-ft seam would produce about 0.5 inches of convergence from the abutment 

loading before it fails. This indicates that a shield should accommodate about 0.5 inches of 

convergence beyond the setting load to reach the yield load and this is consistent with modern 

shield stiffness measures and setting practices where the set pressure is about 70 pct of the yield 

pressure. 

This reasoning would also imply that the shield stiffness, by design, should be adjusted as a 

function the extraction thickness (lower stiffness required at higher seam extractions). Here too, 

the shield response is favorable in that the shield stiffness is reduced with operating height. 

Therefore, assuming the shield stiffness is correctly matched, one shield design should function 

reasonably well through the full range of operating heights. 

Is Yielding a Good or Bad Thing? 

Generally, yielding is good from the perspective of the shield and bad from the perspective of the 

roof. Yielding protects the shield from developing excessive pressures in the leg cylinder that 

can damage the leg or overstress the structural components of the shield. Yielding is 

accomplished by letting fluid out of the leg, which invariably results in the lowering of the 

support canopy. Simply translated, whenever the shield is yielding the roof is moving down. 

Downward displacement of the roof is generally an indication of reduced stability of the rock 

mass. 

Now, how much the roof moves down or how much the support lowers is also relevant to this 

discussion. Yield valves will typically have a 10 pct pressure resetting rating, meaning after a 10 

pct reduction in leg pressure caused by the loss of fluid, the valve will close and allow pressure 

to increase in response to convergence once again (Barczak and Gearhart, 1998). It would take 

several hundred to over a 1,000 yield cycles to fully eliminate the hydraulic stroke of the 

cylinder. However, that is hardly the issue. Lowering of the roof by one inch can be significant 

from a ground control viewpoint. It would take only 8 yielding events to cause one inch of 

downward roof movement or 44 yield events to produce as much as 6 inches of downward roof 

movement in a typical shield design, or you can become iron bound if there is not enough stroke 

left in the cylinders. 

Another controversial issue needs to be raised in this discussion. Why does the yielding stop? 

Since the lowering of the shield increases the leg inclination, the capacity of the shield also 

decreases slightly after each yielding event. If the roof or load condition is strictly acting in 

force control and the worst case scenario is that the shield is supporting the weight of a detached 

block of roof that remains isolated (detached), then the yielding would not cease to occur until 

the shield went solid since the load acting on the shield would not decrease with convergence. 

Since this rarely happens, it suggests that this simplistic model of roof behavior as promoted by 

Wilson is not accurate. This means that, for some reason, the support resistance required to 

produce equilibrium decreases as a result of the yielding or with increasing convergence, which 

is consistent with a ground reaction approach to shield design. It does not necessarily imply that 

the shield is responsible for this condition, it may well be that the shield had nothing to do with it 

and the load conditions changed because of stress relief in the ground due to deformation or 

macro movements of rock masses in the more global sense. However, from a ground reaction 
13 

curve perspective, support yielding is moving down the ground reaction curve toward a condition 

of decreased stability, hence, it must be concluded that ultimately yielding is not a good thing. 

This ground reaction analysis is also useful to reexamine the impact of yielding. Each yield 

event drops the support pressure, and equilibrium is reestablished through additional 

convergence, i.e. moving down the ground reaction curve to a condition of less stability. As you 

move down the ground reaction curve through successive yielding events, the convergence 

required to reestablish equilibrium grows. In addition, the premise thus far has been that you 

remain on the same ground reaction curve. If the yielding degrades the immediate roof structure, 

then the ground reaction curve will be shifting to the right, and thereby cause a dramatic increase 

in the convergence required to reestablish equilibrium and substantially reduce the number of 

yielding events that will create poor ground conditions (Barczak, 2006). 

Set Pressures 

The desire to increase set pressure is another justification given for increasing shield capacity. 

The primary function of set pressure is to strengthen the immediate roof beam by creating a more 

laminated structure. The immediate roof beam does provide some control over the behavior of 

the intermediate roof. Perhaps, more correctly, failure of the immediate roof beam can lead to 

the stress fractures of the intermediate roof. In the face area, this failure can lead to a detached 

block that cannot be controlled by the support, almost regardless of the shield capacity. The 

deflection of the immediate roof beam from the loading of the intermediate roof is the 

mechanism that produces the reduction in convergence associated with increased set pressures. 

However, as previously discussed in reference to figure 3, the inflection point of the ground 

reaction curve indicates that beyond a certain level of set pressure or shield resistance, relatively 

little benefit is gained in terms of reduced face convergence. Essentially, if the convergence is 

produced by mechanisms beyond the control of the shield, such as deformation and yielding of 

the coal face, then increased setting pressures beyond the inflection point could cause premature 

yielding of the support. In other words, if the goal is to prevent yielding of the support, lowering 

the set pressure to accommodate the convergence (that cannot be controlled by the shield) can, at 

times, improve ground control. 

The Impact of Leaking Cylinders 

 This one is obvious to anyone who has ever looked at a ground reaction curve. “Leaking 

cylinders” essentially means you are moving down the curve, allowing more convergence. You 

get to the trough of the ground reaction curve or miss the curve, and poor ground conditions are 

likely to occur. 

Reading Too Much into Leg Pressure Data 

The ground reaction concept suggests that much of shield loading is due to ground responses that 

are beyond the control of the shield. For example, the shield cannot change the stiffness of the 

coal face or the gob structure. The overburden and abutment pressures will produce 

deformations of these structures that cannot be prevented or significantly altered by the shield. 
14 

Likewise, the increase in span, as a result of the shearer cut, is going to cause the majority of 

shield loading in most cases; and here too the shield has limited control. High setting forces and 

stiff shields will help to maintain better control of the immediate roof beam. This can influence 

both the intermediate roof response and the span problem associated with the shearer cut, but no 

matter what the shield capacity, this change in state (movement of the ground reaction curve), 

will produce shield loading in proportion to the shield stiffness. Ultimately, the shield force is 

there to control the immediate roof response, build a good roof beam and keep its response 

coupled to the main roof response. Beyond this, the shield is simply reacting to the main roof 

activity, which is not controlled at all by the shield. 

Since the load environment is largely displacement or convergence-controlled (i.e., not directly 

influenced by the shield capacity), then shield load (i.e. maximum leg pressure) is not a good 

indicator of required shield capacity. Just because the shield is heavily loaded or even if it is 

yielding, does not necessarily mean that the support has inadequate capacity. If the immediate 

roof stays coupled to the main roof, then the shield response can be dominated by the main roof 

response and the coal face behavior. In this case, a higher capacity shield will exhibit the same 

behavior, i.e. heavy loading. In fact, since the higher capacity shield is also stiffer, it may show 

even higher levels of loading. Without a full understanding of the ground response, it is easy to 

see how one could jump to the wrong conclusion by focusing only on the shield loading as a 

measure of the required shield capacity. 

Unsupported Roof Span Issues for Shield Design 

A study by Langosch showed the relevance of the canopy tip to face distance to the probability 

of roof falls or cavity formation in front of shields (Langosch et al., 2003). A critical 

unsupported span was determined based on the thickness and strength of the immediate roof 

layer or composite roof beam. When the unsupported span stays less than this critical distance, 

cavity formation is highly unlikely. Underground assessments were made on 14 German 

longwall faces and an empirical formulation was developed that relates the probability of cavity 

formation to the following parameters when the unsupported span is exceeded: (1) the shield 

resistance, (2) abutment stress, (3) directional orientation of stress induced roof fractures, which 

are likely to be influenced by the disparity in rock strength of the immediate roof and the coal in 

the longwall panel, and (4) the distance by which the critical roof span was exceeded. Figure 8 

shows the relationship for support (load density) capacity for favorable and unfavorable fraction 

orientation; where favorable fraction orientation is such that the coalface helps to prevent 

movement along fracture planes leading to the dislodgment of the rock segments. Here too, the 

trend is the same as that observed with shield resistance and face convergence. The figure 

indicates a diminishing return in terms of reduction in cavity formation as the shield capacities 

are increased. For the favorable fracture orientation, shield load densities as low as 3.5 tons/ft2

reduce the probability of roof falls to less than 10 pct and doubling of the load density to 7.0 

tons/ft2

 lowers the probability of roof fall to only a 3 pct. 
15 

A Discussion of Active Horizontal Loading 

in Two-Legged Shield Design 

Active horizontal roof loading is claimed to be one of the major advantages of the two-leg shield 

design. In theory, the forward inclination of the leg cylinders provides a component of the leg 

force that acts to transfer compressive force into the immediate roof structure acting in the 

horizontal plane parallel to bedding, which can help to maintain stability of a disjointed roof 

beam. The question is, “Does this really happen or to what extent does it happen in the mine?” 

The mechanics of the shield relative to the development of an active horizontal force is 

illustrated in figure 9. The key issue is how much horizontal displacement of the canopy occurs, 

which can be ascertained from the reaction or loading of the lemniscate links. 

The lemniscate links develop stress in response to the horizontal displacement of the canopy 

relative to the base. Without the horizontal displacement, there is very little stress developed in 

the caving shield-lemniscate assembly since it has no vertical stiffness (Barczak and 

Schwemmer, 1988). Using the canopy-caving shield hinge pin as a frame of reference, a 

horizontal displacement of this joint toward the face will produce compression in the top links 

and tension in the bottom links. When the shield is set against the mine roof, if the canopy does 

not slide forward (i.e., the friction prevents slippage of the canopy along the roof line), then there 

Figure 8 - Impact of shield load density on the probability of roof falls in front of the shields when the 

critical unsupported span is exceeded (after Langosch, et al., 2003). 

0

5

10

15

20

25

30

2.0 4.0 6.0 8.0 10.0 12.0

SHIELD LOAD DENSITY, tons/ft2

PROBABLITY OF ROOF FALL, pct .

Favorable orientation of 

fissures (roof fractures)

Unfavorable orientation of 

fissures (roof fractures)

Modern shields

(1000+ ton capacities)
16 

is very little stress developed in the lemniscates links and the horizontal component of the leg 

force is transferred to the mine roof as a compressive acting stress. Roof-to-floor closure 

produces no significant change in the stress in the lemniscate links since the caving shieldlemniscate assembly has no vertical stiffness. Conversely, the leg pressures increase in direct 

proportion to the roof-to-floor closure. Therefore, if the link stresses increase proportionally 

with increasing leg pressures, then the canopy is moving forward causing displacement of the 

canopy-caving shield hinge pin toward the coal face. If the canopy is not moving and instead the 

horizontal force is transferred to the mine roof through the friction between the canopy and the 

roof, then the link stress would not increase in proportion to the leg pressure. Likewise, if the 

canopy were displaced toward the gob, then the previously developed link stresses would 

decrease. In a similar manner, if at any time the canopy lost frictional contact with the roof, then 

the canopy would slide forward and the horizontal force from the leg component would be 

transferred to the caving shield-lemnicsate assembly. This force is then not recoverable in terms 

of active roof loading. It is lost forever. Any additional leg pressure development and active 

horizontal roof loading during a face cutting cycle will be relatively small. 

Unfortunately, very little link stress data has been obtained, particularly from active longwall 

mining shields. The U.S. Bureau of Mines collected some link strain data in the 1980’s. This 

data shows that the link strains or link forces were generally correlated to the leg pressures; again 

suggesting that the horizontal component of the leg force was being transferred into the caving 

shield-lemniscate assembly and not into the mine roof (Barczak and Garson, 1986). Further, the 

magnitude of the link forces provide a caving shield reaction force that is close to the horizontal 

component of the leg force. It is also clear from the data that the link strains were occurring 

immediately upon setting the shield, indicating the horizontal leg force was being absorbed by 

the caving shield and not going into the roof. A recent study of link loading on a modern (1,000 

ton capacity) longwall shield confirms this conclusion, as shown in figure 10, by showing the 

strong correlation between the shield leg pressures and lemniscate link strains. 

Figure 9 - Shield mechanics showing that caving shield/ lemniscate assembly 

reaction to forward canopy displacement reduces active horizontal roof loading 

caused by the horizontal component of the leg force. 

δh

LF CF

Active Horizontal Force (AHF) δh

LF CF

Active Horizontal Force (AHF)

AHF = LF – CF

AHF = Active Horizontal Force
17 

Figure 10 - Comparison of shield leg pressures and lemnsicate link strain. 

y = 0.4699x + 2427

R2 = 0.96

y = 0.5262x + 2356.8

R2 = 0.99

2,500

2,520

2,540

2,560

2,580

2,600

2,620

265 285 305 325 345 365 385 405

LEG PRESSURE, bars

LEMNISCATE LINK STAIN, microstrain .

3300

3400

3500

3600

3700

3800

3900

12/14/06

0:00:00

12/14/06

2:24:00

12/14/06

4:48:00

12/14/06

7:12:00

12/14/06

9:36:00

12/14/06

12:00:00

12/14/06

14:24:00

12/14/06

16:48:00

12/14/06

19:12:00

12/14/06

21:36:00

12/15/06

0:00:00

DATE AND TIME

LEMNISCATE LINK STRAIN, microstrain

0

1200

2400

3600

4800

6000

7200

LEG PRESSURE, psi

FRONT LEMNISCATE LINK STRAINS SHIELD LEG PRESSURE
18 

Dr. Syd Peng directly measured the horizontal displacement of the canopy during the setting 

operation in one underground study of shield supports (Peng, 1990). He found that as about 0.4 

inches of horizontal movement of the canopy was occurring from the time the canopy first 

contacted the roof to the time full set pressure was reached. The horizontal component of the leg 

force was computed as 150 tons. A caving shield/lemnsicate assembly stiffness of 384 tons/in 

would completely eliminate the active horizontal loading in this situation. Measurements of 

caving shield/lemnsicate assembly stiffness measured through full scale tests in the Mine Roof 

Simulator on a 360-ton support ranged from 162 to 340 tons/in (Barczak and Schwemmer, 

1988). Hence, it is likely that most of the force was consumed in link force loading. Dr. Peng 

also showed that when a steel plate was placed above the canopy with rollers sandwiched 

between the plate and top surface of the canopy to create a “zero friction” condition, the 

measured displacement increased to 0.5 inches. In theory, a true “zero friction” condition would 

produce zero active horizontal force. Using a stiffness of 384 tons/in, this difference in 

displacement would indicate an active horizontal force of about 20 tons, an 80 pct reduction in 

the full active horizontal load potential had no slippage occurred (Barczak, 2006). 

Why do the Higher Capacity, Modern, Shields Last Longer? 

Many people believe modern high capacity shields last longer than their lower capacity 

predecessors because of their higher capacity. As previously indicated, the higher capacity 

shields are also stiffer than their lower capacity predecessors. Therefore, from the convergencecontrolled roof behavior and the uncontrollable convergence produced by the main roof loading, 

they are just as likely to be as fully loaded relative to their yield capacity, if not more fully 

loaded, than their predecessors. The manufacturers achieve the longer life from better 

management of the stress in the shield components through improved design. The advancement 

of numerical modeling has matured to the point where prototype shields for new procurements 

are designed to better manage the stress and avoid stress concentrations that have been problems 

in the past. Fabrication methods have improved quality control as well. 

Shield Design – Thinking Outside The Box 

The modern, two-legged, lemniscate shield design has performed very well and some would 

argue why change it. Although its performance record supports its effective design, there are 

design issues which still should be evaluated if further improvements are to be realized. These 

include: 

• High toe pressure – Two-leg shields inherently have high toe pressure due to the 

kinematics of the design. The forward inclination of the leg cylinder, coupled with the 

reactions of the lemniscate links, produces a resultant force acting on the canopy and base 

that induce a moment causing high toe pressures. 

• High link forces – As described in the analysis of horizontal shield loading, the forward 

inclination of the leg cylinder and potential loss of frictional control on the canopy or 

base makes the two-leg shield design prone to high link loading. 
19 

• Low tip loading – Solid canopy designs have prevailed and the canopy ratios have 

grown to the point where tip loading is getting less and less. Canopy ratios have 

exceeded the 2:1 limit which is required to provide some roof pressure over the full 

length of the canopy. Attempts to generate additional tip loading by having the canopy 

tip project upward structurally leads to additional toe loading. 

• Reduced setting forces at full bottom stage leg extension – The setting forces can be 

reduced by as much as 50 pct when the bottom stage is fully extended during the setting 

operation. The condition is created primarily by having the bottom stage extend and 

retract first. A design where the bottom stage extend and retract last would solve the 

problem. 

These deficiencies in support performance suggest that alternative designs might offer some 

additional improvements in shield performance. For example, a more vertical leg orientation 

would reduce the toe loading and caving-shield lemniscate assembly stress, but would require a 

longer canopy to preserve full canopy contact pressure. This would be detrimental if the canopy 

is carrying the full weight of the immediate roof, but would be less of a factor if the roof is acting 

only as a cantilevered beam. Four-leg shield designs provide more uniform base and canopy 

loading, but provide multiple load paths that can cause significant reductions in shield capacity if 

one set of legs act in tension. Another approach might be to include vertical stiffness in the 

lemniscate design. This could be including a hydraulic cylinder between the canopy connection 

and the front lemniscate connection. The caving shield lemniscate assembly could also be 

inverted or reversed. The benefit of this configuration is that it either reverses or repositions the 

state of stress in the links and loading to the canopy and base to a pattern which reduces the 

moment that promotes toe loading. 

Standing Supports – Clearly Bigger is Not Always Better 

Perhaps the misconception that bigger is not always better is best exemplified in standing 

supports for longwall tailgates. Several examples support this claim. The transformation from 

wood to concrete crib structures in the 1980’s provided the first indication that high capacity is 

not always the answer. Concrete, with a compressive strength and material modulus an order of 

magnitude higher than wood provided a unit support capacity of 500 tons compared to only 50 

tons for wood cribbing. “This had to provide superior roof support” was the thinking at that 

time. This hypothesis was completely wrong. These super high capacity concrete cribs were 

crushed in several longwall tailgates (see figure 11). Again, the ground reaction curve explains 

why. Much of the loading outby (before the approach of) the longwall face is uncontrollable. 

The supports simply do not have enough capacity to offset or control the global forces of the 

overburden in creating the abutment pressure on the longwall tailgate. Unless the supports can 

survive the convergence associated with the abutment pressures during advancement of the 

longwall face, the supports will fail prematurely within the realm of practical standing support 

capacities. This example also clearly indicates that the stiffness of the support must be 

considered in determining the required capacity. The reason is that the convergence in not 

controlled by the capacity. A stiffer support will require higher capacity than a less stiff support 

design. 
20 

Premature failures of concrete cribs in pre-driven longwall recovery rooms also support the need 

to match the support stiffness to the ground response to achieve an acceptable support design. 

The yielding of the panel fender, which is carrying a tremendous amount of load, causes 

uncontrollable convergence in the longwall recovery room. Here too, concrete cribs as shown in 

figure 12 with capacities of over 1 million pounds, will fail unless the support can be sufficiently 

softened to absorb the ground deformations without inducing loading beyond the peak capacity 

of the support. 

Another example of this was a trial of two different pumpable support systems on a longwall 

tailgate. The two support systems had similar loading stiffness and residual load capacity, but 

the maximum capacity of the one support was lower than the other due to a weaker grout 

material. During trials of the two support systems on the same longwall panel, the lower 

capacity support yielded outby the longwall face while the higher capacity one did not. Again, 

the stiffness of the support is the relevant design issue. The convergence is largely driven by the 

Figure 11 - Failure of concrete cribs in longwall tailgate due to convergence beyond the control of the 

support system. 
21 

ground behavior, not the support capacity. With similar loading stiffness, the lower capacity 

support failed first, since the capacity of the support did not stop the tailgate convergence. The 

higher capacity support was able to continue loading as the tailgate convergence continued to 

increase as the face advanced toward the supports. This support also yielded just inby the 

longwall face, as its capacity too was exceeded by the face convergence beyond the shield line. 

Numerical modeling can be used to develop ground reaction responses for various mining 

conditions. The FLAC 2D finite difference software was used in a recent investigation of 

standing support in a three-entry longwall tailgate in the Pittsburgh coal seam. FLAC is well 

suited to modeling the layered coal measure rocks, since the bedding layers can be described as 

strain softening ubiquitous joints, while failure of the rock matrix can be simulated as a strain 

softening Coulomb material. The Pittsburgh coal seam is overlain by a weak stack rock followed 

by interlayered black shale and thin coal rider beds providing a weak immediate roof structure. 

A thick limestone structure provides a strong structure that can deform and bridge to the rear 

abutment. Figure 13 shows a ground reaction curve developed from the model for tailgate 

conditions just outby the longwall face due to front abutment pressures. Also shown on the 

Figure 12 - Failure of pre-driven longwall recovery room support due to convergence beyond the 

control of the support during panel fender yielding. 
22 

curve are support responses for a 4-point wood crib, pumpable support crib, and a concrete donut 

crib. As seen by the intersection points of the support loading with the ground reaction curve 

which indicates a condition of ground equilibrium in the model, none of the supports have much 

impact on the resulting convergence in the longwall tailgate despite large differences in support 

capacities, although the wood crib is on a section of the curve, near where convergence does 

increase quickly when capacities are reduced below that of wood cribbing. Figure 14 shows the 

impact of these support systems when abutment loading increases, causing damage and strain 

softening of the rock mass. The roof structure remains intact with the pumpable supports, while 

separations occur with both the concrete cribs and wood crib. The results show that extensive 

yield capability, coupled with adequate support capacity, is required to maintain control over the 

roof deformation. The wood cribs, while having a large yield capacity, do not have sufficient 

resistance to prevent separation in the roof. The donut cribs, while having a significant initial 

capacity, do not have sufficient residual yield capacity to maintain control during extended 

closure. It was concluded from these studies that the stiffness of the support will have a positive 

impact on controlling the ground and minimizing the convergence of the tailgate, provided they 

can sustain their load carrying capacity. The benefit is small until the near seam rock has strain 

softened and is depending on its residual load capacity for stability. After this stage of behavior, 

a support that is too soft or sheds load after yielding can allow too much deformation to occur 

increasing the risk of roof instability and falls. 

Figure 13 - Ground reaction for longwall tailgate just outby the face developed from FLAC model with 

comparisons of wood crib, concrete donut crib, and pumpable roof support systems. 

0.00

0.50

1.00

1.50

2.00

2.50

3.00

3.50

4.00

4.50

0 0.5 1 1.5 2 2.5 3

CONVERGENCE, in

SUPPORT LOAD RESISTANCE, ton/ ft 2 ..

Convergence prior to 

support installation

Ground Reaction Curve 

(Front abutment loading outby the face)

Concrete (donut) crib

4-point wood crib

Pumpable support
23 

CONCLUSIONS 

The goal of any support design is to match the support performance to the ground response. 

From a ground reaction perspective, “bigger is always better” in the sense that higher support 

resistance results in reduced convergence, which ultimately is a good thing. The question can 

also be addressed in another manner. Is “bigger-the-better” justified? If operating on the steep 

part of the ground reaction curve, then the reductions in convergence may be so small, that they 

produce negligible benefits in ground control and at what cost are these convergence reductions 

obtained? 

The bigger-the-better design philosophy is most contentious when it is applied to longwall 

shields. The ground reaction concept is essentially addressing the entire support system, which 

in longwall mining includes the coal face and gob along with the shield support. The shield does 

not control the other two; so in fact, there is an upper limit to shield capability. This was the 

premise behind the design philosophy of matching the shield stiffness to the coal structure 

stiffness. Essentially, there is no need to have the shield do more work than the coal in the 

longwall panel ahead of the shields. Therefore, from that perspective, there is a logical upper 

limit to shield capacity. 

As the discussion on active horizontal loading indicated, it is likely that much of the horizontal 

component of the leg force is consumed internally in the caving shield-lemniscate assembly, 

particularly in weak roof geologies. The larger the shield capacity, the more load that is being 

transferred to these components, with essentially no direct benefit in ground control. Moreover, 

since the joints in the caving shield-lemniscate assembly are the first structural components to 

Figure 14 - Roof deformation in longwall tailgate from FLAC model for wood crib, concrete crib and 

pumpable roof support systems. 

0

2

4

6

8

10

12

14

16

18

20

0 0.2 0.4 0.6 0.8 1 1.2

ROOF DEFORMATION, in

DEPTH INTO THE ROOF, ft .

Wood cribs

Concrete cribs

Pumpable supports
24 

wear out on modern longwall shields, is the “bigger the better” design philosophy improving this 

problem or actually making it worse? Higher capacity also means that canopy and base 

structures must be able to transfer higher stress through the structure. Toe loading on two-legged 

shield designs becomes more of a problem as the capacities increase in greater proportion than 

the size of the support, resulting in higher toe pressures. 

The fact that the shield stiffness increases with capacity results in a self propagating rational for 

increasing support capacity. If the load environment is largely convergence or displacement 

controlled and the shield is simply reacting to the rock mechanics, then the shields will have to 

have higher capacity because they are stiffer and will develop more loading from the 

“uncontrollable convergence” component. Simply stated, a stiffer support will develop more 

loading than a softer support under the same load conditions. This phenomenon is observed 

throughout mining. For example, higher modulus or stiffer rock structures develop more stress 

than lower modulus materials due to strain-induced loading of the rock mass. The higher 

modulus rock must have sufficient strength to handle the high stress or it will fail prematurely, 

while a lower modulus rock can have lower strength and survive. 

The shield’s primary job is to control the immediate roof. It can do this by strengthening the 

roof beam in the same manner that roof bolts help to facilitate beam building, except of course 

the loading capacity and active loading far exceeds that of roof bolts. None-the-less, the theory 

is essentially the same; the moment of inertia of the beam is increased by increasing the frictional 

resistance along laminations. The increased moment of inertia enhances the bending stiffness of 

the beam and reduces beam deflections under loading. There is a limit to beam building 

capability, which occurs when there is no slippage. Beyond this, the rock structure is loaded in 

compression and the immediate roof beam is essentially pressed against the intermediate roof 

structure. As long as the immediate roof beam stays coupled to the intermediate roof structure as 

face convergence occurs, there is no significant benefit developed from additional compressive 

loading of the immediate roof by the shield. 

This logic also indicates that yielding does not indicate inadequate shield capacity. If the 

immediate roof stays coupled with the intermediate roof behavior, and the yielding is produced 

from the coal panel and main roof activity, then the shield is simply reacting to a load condition 

it cannot change. In fact, in this situation, a lower capacity shield may be less likely to yield 

since it is less stiff. 

Obviously, a “bigger-the-better” design philosophy provides insurance against conditions that 

are abnormal. However, history clearly shows us that all of the capacity is not needed all of the 

time, as many longwall panels have been successfully mined with no significant problems at far 

less capacity than is currently provided by modern longwall shields. The argument against using 

all of the capacity all the time is that it unnecessarily causes loading of the shield components, 

particularly the caving shield-lemniscate assembly and high toe loading in two-leg shields. An 

alternative approach would be to develop a “smart loading technology” design that uses the 

capacity only when it is needed. This could be done by lowering the set pressure to some 

acceptable threshold that provides good ground control under nominal conditions, and increases 

the setting pressure when the convergence rate increases due to worsening conditions. The 

convergence could be assessed from the shield loading profile through the leg pressure 
25 

measurements, or convergence instrumentation could be included in the next generation shield 

design. This type of control logic is not beyond the capability of today’s control system 

technology. 

Clearly, concrete supports taught us that not all convergence can be stopped by practical standing 

support systems, and that this uncontrollable convergence must be considered as part of the 

design process. Here too, the stiffness of the support is an essential element of the support 

design. Stiffer supports will require higher capacity because there is always a component of 

convergence that cannot be prevented. As such, yielding supports will always provide the best 

ground control capability. The goal of standing support design is again to match the support 

response to the ground conditions, and prevent failure caused by inadequate yielding of the 

support. 

In conclusion, is bigger better? No, bigger is not always better, but is often required because of 

the support design. The ground reaction curve holds the key to optimizing support design. 

NIOSH is continuing to explore the application of ground reaction design for roof support. Do 

you really think shield capacities can continue to increase forever? 

REFERENCES 

1. Barczak, T.M. and Garson R.C. (1986). Shield Mechanics and Resultant Load Vector 

Studies. U.S. Department of Interior, Bureau of Mines RI 9027, 43 pp. 

2. Barczak, T.M. and Gearhart, D.F. (1998). Performance and Safety Considerations of 

Hydraulic Roof Support Systems. Proceedings: 17th International Conference on Ground 

Control in Mining, Morgantown ,WV, pp 176-186. 

3. Barczak, T.M. and Schwemmer, D.E. (1988). Stiffness Characteristics of Longwall 

Shields. U.S. Department of Interior, Bureau of Mines RI 9154, 14 pp. 

4. Barczak, T.M. and Schwemmer, D.E. (1988). Horizontal and Vertical Load Transferring 

Mechanisms in Longwall Supports. U.S. Department of Interior, Bureau of Mines RI 9188, 

24 pp. 

5. Barczak, T.M. (2006). A Retrospective Assessment of Longwall Roof Support with a 

Focus on Challenging Accepted Roof Support Concepts and Design Premises. Proceedings: 

25th International Conference on Ground Control in Mining, Morgantown, WV, Aug. 1-3, 

2006, pp. 232-244. 

6. Brown E.T., Bray J.W., Ladanyi, B. and Hoek, E. (1983). Ground Response curves for rock 

tunnels. J. Geotech. Eng., v 109, pp. 15-39. 

7. Esterhuizen, G.S. and Barczak, T.M. (2006). Development of Ground Response Curves for 

Longwall Tailgate Support Design. GoldenRocks-2006. 

8. Itasca Consulting Group, Inc. (2005). FLAC Version 5.0, User’s Guide, Minnesota. 

9. Langosch, U.L., Ruppel, U. and Witthaus, H. (2003). Longwall Roof Fall Prediction and 

Shield Support Recommendations. Proceedings: 22nd International Conference on Ground 

Control in Mining, Morgantown ,WV, pp 27-33. 

10. Medhurst, T.P. and Reed, K. (2005). Ground Response Curves for Longwall Support 

Assessment, Trans. Inst. Min. Metall. A, Mining Technology, V114, pp A81-88. 

11. Peng S.S. (1990), Design of Active Horizontal Force for Shield Supports for Controlling 

Roof Falls. The Mining Engineer, pp 457-462. 
26 

12. Smart, B.G.D. (1986). The Evaluation of Powered Support Performance from Geological 

and Mining Practice Information. Proceedings: 27th U.S. Rock Mechanics Symposium, 

Alabama, pp. 367-377. 

13. U.S. Longwall Census – Population Pushes Past 50 (2005). Coal Age, February, pp. 26-30. 

14. Wilson, A.H. (1975). Support Load Requirements on Longwall Faces. The Mining 

Engineer, pp. 479-491.