A RETROSPECTIVE ASSESSMENT OF LONGWALL ROOF SUPPORT WITH 

A FOCUS ON CHALLENGING ACCEPTED ROOF SUPPORT CONCEPTS 

AND DESIGN PREMISES 

Thomas M. Barczak, Mining Engineer 

NIOSH-Pittsburgh Research Laboratory 

Pittsburgh, PA, USA 

1

Mention of company name or product does not constitute endorsement by 

the National Institute for Occupational Safety and Health. 

ABSTRACT 

The era of the shield-supported longwall face began in the 

United States a little over 25 years ago. The most fundamental 

development of the shield in the past 25 years has been an increase 

in size and capacity and a progression toward two-legged designs 

in favor over four-legged shields. Is the “bigger-the-better” design 

philosophy justified? The advantage of the two-legged design is 

largely attributed to its active horizontal force caused by the 

forward orientation of the leg cylinders. How do we know this? 

You might be surprised to learn that shield loading data may 

suggest otherwise. In addition, what about setting pressures? 

Should they be set as high as possible? Some people have 

advocated that setting pressures should equal the yield pressure. 

Does this make any sense? Is yielding a good or bad thing? These 

are some questions still worthy of investigation that are addressed 

in this paper. 

The first major change in gateroad support was the use of cable 

bolts to replace conventional wood cribbing particularly in Western 

mines. Standing support changed very little until the mid-1990s, 

when Strata Products 1 introduced the Hercules and Propsetter 

supports in Eastern mines and Burrell Mining Products introduced 

the Can support in Western mines. Since then, there has been a 

dramatic increase in new roof support technologies for longwall 

gateroad support. Historically, standing support, including shield 

supports, have been designed based on a simplistic model requiring 

the support to have sufficient capacity to support the gravity 

loading of a perceived detached rock mass. This approach ignores 

the stiffness of the support and the resulting ground deformation 

that is known to be critical to roof stability. Recently, a new 

approach based on the ground reaction concept has been promoted 

by the National Institute for Occupational Safety and Health 

(NIOSH) to account for the interaction of the support with the rock 

mass behavior in achieving ground stability. In addition to 

gateroad support design, this concept has ramifications for shield 

capacities, setting pressures, and recovery room support design that 

challenge current practices in these areas. 

Although longwall mining has matured into the safest and most 

productive underground mining method used today, traditional 

concepts of roof support design and practice should still be 

evaluated and challenged. Ultimately, advancements in the science 

of ground control are made by recognizing deficiencies in current 

theories and thinking outside the box. 

PRELUDE 

Throughout the history of longwall mining, the design of the 

roof support system has been critical to the success of the mining 

operation. Early forms of longwall mining used wood props for 

face support and packwalls made from roof and floor rock to 

control caving of the immediate roof (1). These systems were 

replaced by powered roof supports that could be advanced easily 

while allowing the strata to cave behind them. The first powered 

roof supports were simple frame and chock structures. These 

designs were poor in their ability to resist horizontal displacements 

and the resulting moment loading caused by the strata dynamics 

during the caving process. They often experienced difficulty and 

failure (2). The shield greatly improved kinematic stability by 

providing a horizontal stiffness by structurally connecting the 

canopy to the base. The success of the shield support promoted the 

application of the longwall mining method in highly faulted and 

massive strata conditions where caving was difficult to control and 

where chock and frame supports were inadequate. 

The basic shield design has changed little since its inception in 

the 1970's. The primary change has been a steady increase in 

shield capacity, resulting in larger hydraulic cylinders and support 

structures. The consequences of a “bigger-the-better” design 

philosophy are examined in the paper. However, more importantly, 

it is surprisingly interesting to see how the large capacities are still 

rationalized on the misconception that the shield loading provides 

full control of the ground, when in fact a large portion of the shield 

response is from ground behavior that is independent of the shield 

capacity or design. Misconceptions also exist about setting 

pressures, shield stiffness characteristics, and what happens during 

yielding that further challenge the “bigger-the-better” design 

philosophy and the rationale for the need for higher capacity to 

extend the service life of the shield. 

There has also been a definite trend in the United States and 

world wide toward two-leg shield designs. With the leg cylinder 

inclined toward the coal face, this provides the opportunity for the 

shield to induce an active horizontal force toward the coal face that 

helps maintain stability of a disjointed immediate roof structure. In 

addition, the caving shield lemniscate assembly provides a passive 
 

 

 

 

 

 

 

 

 

 

restraint to face-to-waste strata movements. While these claims are 

true in theory when certain assumptions are made, is this what 

actually happens in practice? Here too, the evidence may be 

surprising. 

What about standing supports? The transformation in passive 

gateroad support has been much more varied than in the powered 

face supports. Prior to 1980, timber was the dominant support 

material with wood cribs and timber posts used exclusively for 

tailgate and bleeder support. In the 1980’s, various trials of 

concrete supports were conducted. These also revealed 

misconceptions about ground control regarding the “bigger-thebetter” design philosophy. Concrete, with a compressive strength 

and material modulus an order of magnitude higher than wood, had 

to provide superior roof support was the thinking at that time. 

Wrong again! Fill a tailgate or a longwall recovery room full of 

concrete cribs and everything will be fine. I don’t think so! Why, 

because fundamental rock mechanics issues were misunderstood or 

ignored. 

Well, if all of this is true, then how can we have had so much 

success? Record longwall productions and downward trends in 

blocked tailgates cannot be denied. True, but I would propose that 

these successes are still largely a product of trial-and-error with a 

little science thrown in. Things have evolved because we have 

learned from our mistakes. However, I think the science still lags 

practice and to some extent, successes have been achieved in spite 

of the roof support design. 

Sounds controversial, I know. However, I think it should make 

for an interesting discussion in this 25th anniversary of the 

International Conference on Ground Control in Mining. 

POWERED ROOF SUPPORTS 

Historical Overview of Shield Design 

Two fundamental changes in shield design have been made 

since the introduction of the shield in 1975: (1) the caliper design 

was replaced with a lemniscate-guided caving shield that maintains 

a constant tip-to-face distance throughout its operating range; and 

(2) electrohydraulic control systems have replaced manual systems 

to permit remote and automated operation of the shield. 

Material Specifications -- In order to accommodate everincreasing capacities, while still maintaining component cross 

sectional dimensions small enough to provide adequate working 

space and manageable shield weights, the trend has been toward the 

use of high strength steels (100,000+ psi yield). One consequence 

of some high strength steel applications is that special welding 

practices are required (heat control, etc.), making underground 

repairs more difficult. 

Capacity -- Support capacity has continued to increase throughout 

the history of longwall mining. This trend has continued during the 

past two decades as shown in figure 1 (3). Average support 

capacities in the United States have increased by 70 pct since 1985 

and by 32 pct since 1995 to an average support capacity at yield of 

879 tons for the 52 operating longwalls in the U.S. in 2005. Ten 

installations (19 pct) employ shields with capacities greater than 

1,000 tons, and 37 installations (71 pct) have capacities equal to or 

greater than 800 tons. The highest capacity shield is 1,280 tons (3). 

As shown in figure 1, maximum shield capacities have evolved 

from 800-ton shields, which were common from 1985 to 1990 to 

980-ton shields, which were frequently installed from 1993 to 1998, 

to the 1,170-ton designs from 1999 to 2001, and then the 1,280-ton 

shields first installed in 2002. 

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 

YEAR 

SHIELD CAPACITY, tons

AVG SUPPORT CAPACITY 

MAX SUPPORT CAPACITY 

Figure 1 – Trend in shield capacities. 

Type of Shield -- Beginning in 1985, there has been a steady 

increase in the use of two-leg shields. Today, all shields in the 

United States are two-leg designs. The application of high-capacity, 

two-leg shield designs was primarily controlled by the development 

of larger size hydraulic cylinders. A 1,280-ton shield requires a 

bore diameter of approximately 16.2 inches (412 mm). 

Size -- Longwall shields have increased in length to accommodate 

one-web-back operations and larger face conveyors and deeper 

shearer webs, making these units over 20 feet in length. The 

standard canopy width is now 1.75 m, replacing the 1.5 m designs 

that were prevalent 10 years ago. The first 2.0-meter wide shield 

was installed in Foundation Coal’s Cumberland mine in 2003. This 

trend of increasing canopy width to 2.0 m is likely to continue. The 

2.0-m width may represent an upper limit with current shield 

construction materials since wider shields weigh more, thereby 

requiring more effort to transport during face moves. The potential 

benefit of wider shields is less cost since much of the cost is due to 

the machining of the hydraulic leg cylinders, and fewer legs are 

required with wider supports. However, this cost advantage is 

being offset by higher capacity systems. The wider supports may 

also be more stable in thick-seam operations. A decrease in move 

time might also be realized, since 15 pct and 30 pct fewer supports 

are employed on a face with the 1.75-m wide and 2.0-m wide 

designs, respectively, compared to the 1.5-m designs of the initial 

generation of shield supports. 

Setting Forces -- Setting forces have increased in proportion to the 

increase in yield capacity because the size (diameter) of the leg 

cylinders has increased to accommodate the higher yield capacities, 

while the hydraulic setting pressures have remained constant in the 

4,200 to 4,400 psi range. Most hydraulic power supply systems 

and electrohydraulic control technologies still try to maintain a setto-yield ratio in the 0.6 to 0.8 range. Optimum setting forces as a 

function of overall support loading and geological conditions are 

not pursued by current control technologies. 

Component Constructions -- The component that has changed the 

most is the canopy. Extensible canopy designs were occasionally 

used in Europe, where friable roof geologies had to be maintained 

(4). The trend in the United States has been towards rigid canopies 

without any extension or flipper arrangements to control face 

sloughage, except in thick seam (> 10 ft) operations. Canopy tips 

have been curved upward to promote tip contact. However, the 

force generated at the tip is typically about 10 pct of the leg force 

and is mostly a function of the distance of the tip from the leg 

cylinders. The distribution of loading on the canopy and base is 
also dependent upon the stiffness of these components and the 

deformation characteristics of the immediate strata. There has been 

increased usage of solid base configurations to alleviate splaying 

that occurs with individual base pontoons and to improve load 

distribution on the mine floor. 

Control Systems -- Improvements in the electrohydraulic control 

systems continue to be made. Positive-setting algorithms are more 

sophisticated with pressure monitoring at programmable thresholds 

to more intelligently achieve full setting, or deactivation if full 

setting is incapable of being achieved due to some sort of hydraulic 

failure on a particular leg cylinder. Improvements have been made 

in face-end control logic to allow for full automation of bidirectional cutting. The cycle times for shield advance continue to 

be minimized. Combinations of multiple functions pertaining to 

the powered lowering, base lift, and conveyor push of individual 

shields are now routinely performed at the same time to compress 

the support advance time requirements. 

Hydraulic Components -- Solenoid-operated valving systems are 

now standard. Spool valves have been shown to be superior to ball 

and seat designs that are prone to contamination problems. In 

addition, these systems allow the solenoid to be activated upon 

demand, unlike previous systems that required the hydraulic feed to 

be interrupted by a control solenoid. This leads to both quicker and 

smoother control of support functions. 

Is the “Bigger-the-Better” Design Philosophy Justified? 

 Perhaps, but not for the reasons you might think. First, the 

question itself is somewhat misleading. The shields were forced to 

get bigger to accommodate wider web extractions and wider face 

conveyors to convey the coal from the face at ever-increasing 

production rates. Therefore, the bigger part is understandable, 

although it is interesting to note the trend for higher capacities 

occurred before the increase in shield size. Why? The question 

can be analyzed from the perspective of rock mechanics models for 

longwall mining. 

Figure 2 – Detached block model of longwall roof behavior. 

 Wilson was the first to promote the detached block-loading 

concept back in the early 1970's (5). 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 2). 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. 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 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. 

These factors started the ball rolling along the path of higher 

capacity, and I personally believe this model is an 

oversimplification of the rock behavior. I believe that elements of 

this thinking persist in roof support design even today, and they 

continue to promote the belief in the “bigger-the-better” design 

philosophy. 

 Another motivating factor for the continued increase in support 

capacity was the desire to have higher and higher setting forces. 

Measurements taken during the early growth of the longwall shield 

capacities from the mid-1970’s to the early 1980’s indicated that 

increasing setting force resulted in less face convergence 

suggesting a force-controlled model of rock mechanics for longwall 

mining. This is conceptually illustrated in figure 3 which shows 

that some convergence cannot be controlled by the shields due to 

their limited capacity relative to the ground stress and that the 

benefit in reducing the convergence is diminished once a certain 

setting pressure is achieved (6). Further, these measurements were 

generally taken when shield capacities and thus setting forces were 

small compared to today’s standards. Even these measurements 

show the benefit from increased setting forces diminished with 

relatively small increases in setting forces, suggesting that a 

practical optimum setting pressure exists, well below that of 

today’s standards. For example, Gupta made the claim in his 1982 

Ph.D. dissertation that in British coal mines increasing the setting 

load density from 1.15 to 4.39 tons/ft2

 completely converted the 

tensile stress zone above the shields into a compressive zone and all 

but eliminated ground control problems (7). Today’s shields 

sustain setting load densities of over 8 tons/ft2

. 

Figure 3 – Impact of shield capacities (setting forces) on 

convergence. 

 From a rock mechanics perspective, it is important to 

understand what the increased setting pressures are trying to 

accomplish. The benefit of the setting pressure, as in all active 

support loading, is to “precondition” the immediate roof in order to 

enhance its structural stability. Can this be done? The setting force 

will act to increase the bending stiffness of the immediate roof by 
 

 

 

 

 

 

 

 

 

 

 

 

 

 

developing increased shear resistance along laminations (8). In 

theory, the stiffer cantilevered roof beam will result in reduced 

deflections, thereby reducing stress development in the beam and 

lower shield loading beyond the setting load. Better ground control 

and less shield loading, just what we are looking for, right? 

However, that is not the whole story. In addition to increasing the 

bending stiffness of the roof beam, the setting force may also 

increase the stiffness of both the roof and floor structure. The 

consequence of this is higher shield loading since the equivalent 

stiffness of the roof-shield-floor structure increases. Therefore, the 

increase in setting pressure produces contradicting benefits from 

the perspective of shield loading and roof control. 

As more realistic models of longwall rock mechanics were 

developed, shield capacities continued to grow. Of these, I believe 

the cantilevered beam model integrating the stiffness of the coal 

seam and gob material into the deformation response of the roof 

beam was the most valid. Smart was one of the first to promote a 

stiffness model to determine longwall support requirements (9). As 

illustrated in figure 4, Smart suggested that the powered support 

was one of three yielding foundations which support the immediate 

roof and bridging beds. The other two elements being the coal 

ahead of the face and waste (gob) behind the face. 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 ensure that the immediate roof does not 

become unstable, resulting in excessive support loading leading to 

ground falls in front of the supports. The key to Smart’s concept is 

that the stiffness of the coal and gob largely control the shield 

loading and design requirements. In particular, Smart hypothesized 

that the reason American 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-factorcaving 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. 

Figure 4 – Concept for evaluating roof shield response based on 

the stiffness of the shield, coal, and gob foundations. 

perspective of shield capacity. I do not give the shields as much 

credit as Smart implies in controlling the impact of the abutment 

pressures on the intermediate or immediate roof structure ahead of 

the face. I believe the shields are much more “reactive” as opposed 

to “controlling” in this phase of the rock mechanics. Most 

importantly, I believe that Smart’s model was much more 

conducive to displacement or convergence-controlled rock 

mechanics behavior than the detached block concepts widely 

accepted before this. It is my belief that a significant portion of the 

shield loading is dictated by the reaction of the shield based on its 

stiffness to the uncontrollable convergence that occurs mostly from 

the main roof behavior and subsidence loading that Smart is 

addressing. 

Therefore, is a bigger-the-better design philosophy justified? 

The field data from the 1970’s and early 1980’s clearly shows that 

there is a benefit, to increased capacity to a point. However, it also 

shows that the benefit at least from a convergence control 

perspective diminishes with increasing capacity. I do not think 

there is sufficient data from modern shields to quantify the 

optimum, and it is likely to be different for various loading 

conditions including depth of cover, extraction height, and roof 

geology. However, based on the data from the 1970’s and 1980’s, 

today’s shields are well beyond the 2-3 tons per sq ft that produced 

inflections in the load density vs. convergence curves of that era. 

Another argument I do not believe holds merit is that the higher 

capacities are needed to maintain constant load densities over the 

supported roof area. The whole concept of analysis based on 

normalized pressure is misleading. The load distribution acting on 

the canopy is not even close to being uniform. Normalizing to the 

canopy or support area tends to diminish the true increase in load 

density of today’s shields because the capacities have grown at a 

larger rate than the size of the shields. The one element that does 

lend some credibility to the “bigger-the-better” design philosophy 

has been the fact that the shield stiffness has increased in direct 

proportion to the increase in capacity. Why? Historically, the 

design practice has been to increase shield capacity by increasing 

the diameter of the leg, while keeping the yield pressure constant at 

about 6,200 psi. Because of the increase in leg area, the stiffness of 

the shield also increased in direct proportion to the increase in 

shield capacity (10). It may well be that the increased stiffness may 

be more responsible for improved ground control than the increased 

capacity. In other words, if the capacities we have today would 

have been achieved by reducing the stiffness of the shield, would 

we have the same improvements in ground control that are reported 

now. Personally, I do not think so. However, it is double-edge 

sword so to speak. The consequence of the increased stiffness is 

also greater shield loading from the uncontrollable convergence 

component of the rock mass behavior. In other words, as the 

shields grew in capacity and the demand for increased setting 

forces persisted as part of this growth, they had to have higher 

capacity just to accommodate the higher stiffness and resultant 

loading in a largely convergence-controlled load environment. 

That is the strongest justification I can give for a “bigger-thebetter” design philosophy, and yes, it too has diminishing returns 

suggesting that an optimum capacity exists. 

A Discussion of Active Horizontal Loading in TwoLegged Shield Design 

Active horizontal roof loading is claimed to be one of 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, which can help to maintain stability of a disjointed roof 

I believe this to be a realistic model of longwall support and 

strata interaction with one missing component, that being that the 

subsidence factor that Smart was talking about, in particular the 

elastic response of the overburden is uncontrollable from the 
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 5. The key issue is how much horizontal displacement of the 

canopy occurs, which can be ascertained from the reaction of the 

lemniscate links. 

Active Horizontal Force (AHF)

 

Figure 5 – Shield mechanics showing that caving shield/ 

lemniscate assembly reaction to forward canopy displacement 

δh

LF CF

Active Horizontal Force (AHF) 

AHF = LF – CF 

AHF = Active Horizontal Force 

reduces active horizontal roof loading caused by horizontal 

component of the leg force. 

The lemniscate links develop stress in response to 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 

(11). 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 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 shield-lemniscate 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 canopycaving 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 shieldlemnicsate 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 (see figures 6 and 7); 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 

(12). Further, the magnitude of the link forces provide a caving 

shield reaction force that is close to the horizontal component of the 

leg force. On occasion, data would show that the link strains 

decreased during the shield cycle indicating that the roof may have 

been moving toward the gob and pushing the canopy back toward 

the gob in the process (12), but these events were far less frequent 

than the trend of increasing link strains in proportion to the leg 

pressures. 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/ 

lemniscate assembly and not going into the roof. 

Left front link

Right front link

-350 

-100

-150

-200

-300

-100 

-150 

-200 

-300 

Left front link 

Right front link 

4000 5000 

LINK STRLINK STRAINS, microstraincrostrain 

LEG PRESSURES, psi ESSURES, psi 

Figure 6 – Direct correlation found between link strain and leg 

pressures suggesting leg force is being absorbed by caving shieldlemniscate assembly and not induced into the mine roof. 

6000 

Figure 7 – Link strain and leg force data for selected shield 

loading cycles. 
 

 

 

 

 

 

 

 

 

 

 

 

 

 

Peng directly measured the horizontal displacement of the 

canopy during the setting operation in one underground study of 

shield supports (13). He found that about 0.4 inches of horizontal 

movement of the canopy was occurring from the time the canopy 

first contacted the roof to the time when full set pressure was 

reached (see figure 8). The horizontal component of the leg force 

was computed as 150 tons. A caving shield/lemnsicate assembly 

stiffness of 192 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 (14). Hence, it is likely that most of the force 

was consumed in link force loading, although Professor Peng 

estimated that 100 tons of active load could have occurred. Peng 

also showed that when a steel plate was place 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 192 tons/in, this difference in displacement 

would indicate an active horizontal force of about 20 tons. Peng 

also found examples where the difference in the “zero-friction” and 

“normal” canopy contact produced larger variations in measured 

horizontal displacement, suggesting that higher active horizontal 

forces were developed. 

 

 

“Zero friction” canopy 

contact condition 

Normal canopy contact 

condition 

Displacement before 

roof contact 

Figure 8 – Measurements of horizontal canopy displacement 

during setting operation of longwall shield (from Peng). 

Is Yielding a Good or Bad Thing? 

I think this one is easy. Yielding is good from the perspective 

of the shield and bad from the perspective of the roof. Let me 

explain. 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 lowering 

of the support canopy. Simply translated, whenever the shield is 

yielding the roof is moving down. Downward displacement of the 

roof is generally not a good thing. 

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 (15). It would take several hundred to 

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

the cylinder. 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 6 inches of downward roof 

movement on a typical modern shield. 

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 is acting in 

force control and the worst case scenario is 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 curve perspective, 

yielding is moving down the ground reaction curve toward a 

condition of decreased stability, hence, it must be concluded that 

ultimately it is not a good thing. 

Why do the higher capacity, modern, shields last longer? 

Many people believe the modern high capacity shields last 

longer than their lower capacity predecessors because of their 

higher capacity. I do not believe this to be true. 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. I believe the manufacturers achieve the 

longer life from better management of the stress in the shield 

components through improved design. I think 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. I would suspect that fabrication methods have also 

improved quality control as well. 

In summary, shields have proven to be effective roof supports 

providing adequate face control in a variety of mining conditions. 

However, fundamental issues regarding design requirements and 

their intervention in ground behavior are still not well understood. 

The “bigger-the-better” design philosophy has worked, but does not 

address persistent controversial issues that need to be considered to 

pursue optimization of the support system with the realization that 

capacities cannot continue to increase forever. 

GATEROAD SUPPORT 

Historical Perspective of Gateroad Support Technologies 

There has been a dramatic change in standing roof support 

products for longwall gateroad support. Timber supports in the 

form of wood cribs and posts were all that was used prior to mid1980, when trials of concrete supports were common. Then, the 
 

 

 

1990’s saw a revolution in new support technologies, providing a 

wide range of support capabilities. Figure 9 shows a distribution of 

several types of tailgate supports over the past 14 years. Many 

lessons were learned during the past 25 years, both in terms of 

support design and design requirements for longwall mining. 

50 

60 

70 

90 

80 

40 

30 

20 

2005 10 2001 

0 1996 

1993 

SUPPORT USAGE, pct 

Wood Cribs

Timber Posts

Concrete Cribs

Eng Timber Cribs

Cables

Yld Timber Props

Burrell Can

3-C Support

Steel supports

Pumpable cribs 

Figure 9 – Distribution of tailgate support since 1993 showing 

change in support practices from conventional wood cribs and 

posts into a variety of alternative support technologies. 

What Concrete Supports Taught Us 

about Ground Control 

As indicated in the introductory comments, the thinking at the 

time when concrete supports were introduced to longwall mining 

was that with a compressive strength and material modulus an 

order of magnitude higher than wood, concrete had to provide 

superior roof supports. This hypothesis was completely wrong. 

Why, because fundamental rock mechanics issues were ignored or 

misunderstood. Plain and simple, you cannot prevent all of the 

closure of a mine opening with any practical and economical 

support system, and if the support cannot survive this 

uncontrollable convergence, it will fail prematurely and not provide 

adequate roof support. Any standing support or a longwall shield 

for that matter cannot control the behavior of the overburden. 

These super high capacity concrete cribs were crushed in several 

longwall tailgates (see figure 10). 

Figure 10 – Failure of high capacity concrete crib in longwall 

tailgate. 

A similar argument can be made in longwall recovery 

operations. One study, based on empirical data and qualitative 

assessments of predriven recovery rooms, suggested that the design 

requirement for successful roof support was 150 psi of support 

pressure (16). This too is misguided in that the stiffness and yield 

characteristics of the support are ignored. Concrete cribs fail at less 

than 1 inch of convergence. The uncontrollable component of 

convergence in these recovery room operations can exceed one 

inch causing premature failure of concrete cribs since they are not 

able to sustain any useable load capacity after reaching their peak 

load capacity. Here too, failures of high capacity concrete cribs 

were observed in practice (figure 11). 

Figure 11 – Failure of high capacity concrete crib in longwall 

recovery room. 

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. As such, the simplistic model of detached block 

loading was no longer completely valid for support design. 

Improvements in Material Handling 

and Support Installation 

Part of the revolution in standing support design has been the 

development of support systems that can be installed more easily 

with less manual effort and in less time. The Can support, which 

actually began the revolution out of a need to find an alternative to 

conventional wood cribbing in western mines, was the first system 

to be installed with a machine. The installation of this system 

overcomes the bulkiness and heavy weight of the support by using 

an articulated claw attached to a scoop or load-haul dump vehicle 

to pick the support up and place it in the desired spot in the mine 

entry. Recently, pumpable support technologies have taken a 

different approach by pumping the support materials from the 

surface into an empty, lightweight, containment bag hung from the 

mine roof. While it can be argued that the material handling 

function was simply transferred from underground to the mine 

surface, the work environment above ground is far less restrictive 

and more conducive to safe working practices than is the 

underground environment. Material handling advantages are also 

included in crib type supports such as the Link-N-Lock or Hercules 

crib, where the components have been reduced in size and weight 

while providing several times the support capacity with less support 

material. One of the first alternative products was the Propsetter 
 

 

 

 

 

 

 

 

 

support. This support provided capability equivalent to that of the 

wood crib, but was fashioned as a timber post that could be much 

more easily installed than the piece-meal construction required by 

wood cribbing. Installation rates improved by as much as 200 pct 

with the Propsetter support. Other than the Propsetter, most proptype supports were non-yielding, and therefore would not function 

well in many longwall applications. Within the past 5 years, there 

has been a dramatic increase in prop-type support development, 

primarily because of their ease of installation. These include: 

(1) Little and Big John Extreme Plus Props, (2) Rocprop, (3) Pencil 

Prop, (4) Spider Prop, (5) Ball Buster, and the (6) MX Prop. 

Overall, their performance has been satisfactory, with most 

premature failures attributed to eccentric load conditions in systems 

where the yield load is too close to the buckling capacity of the 

props. 

NIOSH Full-Scale Support Testing 

and the STOP Software 

All of the support products now on the market have been fullscale tested in the NIOSH Mine Roof Simulator active load frame 

using a rigorous protocol that subjects the supports to a variety of 

loading conditions that are the most realistic simulation of the inservice loading. The design requirements, performance 

characteristics, material handling specifications, and cost 

information for 144 designs of 56 different support systems are 

currently implemented into the Support Technology Optimization 

Program (STOP) (17). New support systems are added annually, as 

they are evaluated and commercialized. The implementation of 

this information into the STOP design software has helped to 

eliminate the trial-and-error approach to new support technologies 

that was common in the past. This provides insurance toward 

successful support applications without exposing miners to 

uncertain, potentially hazardous support performance. STOP can 

provide an engineering foundation to ensure that inadequate 

support designs, as well as ultraconservative support applications, 

are avoided. Safety will be improved by matching support 

performance to mine conditions. This reduces the likelihood of 

roof falls and blocked travel and escapeways. Figure 12 shows that 

tailgate blockages during the past 15 years have decreased by 

64 pct based on MHSA reported data on roof falls. 

80 

70 

60 

50 

40 

30 

20 

10 

0 

Frequency 

1990 1992 1994 1996 1998 2000 2002 2004 

Year 

Figure 12 – Graph illustrating decreasing trend in longwall tailgate 

blockages. 

The Australian Experience 

An Australian Coal Association Research Program (ACARP) 

project was commissioned in 2005 to evaluate longwall tailgate 

standing support design (18). The major outcome of this research 

was the promotion of a “skewed roof deformation mechanism”. 

“Skewed roof” refers to the propensity of the gateroad roof to skew 

towards the adjacent or approaching gob. This activity creates 

shear stress damage to the immediate roof and can often destroy 

intrinsic support, including cable bolts. The major finding of the 

project was that existing standing support products are often too 

soft and of insufficient capacity to effectively interact with the 

strata deformation mechanisms, particularly the “skewed roof” 

condition where 3D numerical modeling indicated that capacities of 

at least 400 metric tons would be required to limit the lateral roof 

movement. 

The “skewed roof” roof movement concept obviously has 

serious implications for standing support design. Equivalent 

measurements of lateral movement have not been obtained for U.S. 

longwall conditions and while lateral movements and associated 

support loading have been observed, my initial thought is that it is 

not as prevalent in the U.S. This may be related to fact that U.S. 

mines typically employ a three-entry gateroad design in lower seam 

heights, which if wider would reduce the effect of the skewed roof 

loading in U.S. gateroads. Some western longwall mines in the 

U.S., which have used two entry designs in deep cover conditions, 

have experienced lateral movements of the roof relative to the floor. 

Preloading of Standing Supports 

Until recently, with the introduction of hydraulic pressure cells, 

preloading of standing supports were limited to a few products, 

primarily the Propsetter support system, which utilized a 

pressurized grout bag on top of a yielding timber post (19). Mettiki 

was one of the first U.S. mines to routinely employ the prestressed 

Propsetter support and reported improved ground control in 

difficult tailgate conditions. The hydraulic prestressing devices are 

two thin sheets of metal that are welded along the perimeter to form 

a capsule that can be filled through an inlet port with plain water 

and pressurized to create a preload on a passive standing roof 

support product. The units can be sized to accommodate various 

support geometries. Yield valves can also be incorporated to allow 

the units to act as load limiting devices and provide some 

controlled yielding to otherwise non-yielding support structures. 

I don’t think any rock mechanics engineer can effectively argue 

that preloading is not beneficial to ground control. The argument is 

really about how much benefit is actually achieved, given the zone 

of influence of the prestress and the practical limitations of 

sustaining it on a standing roof support system. I will reserve the 

right to debate that subject in another paper, but I think an 

important point that needs to be recognized is that the secondary 

benefit of preloading standing support products is that it helps to 

offset many of the installation practices that routinely degrade the 

support response. The examples are too numerous to discuss in 

detail here, but suffice it to say that most mistakes are made in 

“topping off” a support to establish roof contact. If this is not done 

properly, a support such as the Can, which on its own has an 

outstanding performance response in terms of stiffness and stability, 

can be made to act like a poorly built wood crib. 

The Missing Piece of Support Design 

The performance characteristics of standing roof support 

systems have been well defined through rigorous full-scale testing 

and documentation in the NIOSH Support Technology 

Optimization Program (STOP). The ground response and the 

interaction with the support, however, have been poorly understood. 

How much control do the standing supports have on roof behavior? 

Do they influence the stress changes and ground movement around 

the opening? Until this understanding is developed, proper support 
 

 

 

 

 

 

 

design cannot be accomplished, and premature failures of supports 

or excessively conservative applications of support will continue. 

The goal is to match the support to the load imposed by the 

surrounding rock mass. The lack of knowledge in this area is 

responsible for the misconceptions and poor premises that still 

plague proper support design. Recently, the concept of support 

design based on ground reaction has been gaining strength here and 

abroad (20, 21, 22). 

UNDERSTANDING THE GROUND REACTION CURVE – 

THE KEY TO ALL SUPPORT DESIGN 

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 and was used in the 

tunnel industry (23). So why the big deal about it now? The 

problem with the ground reaction concept is not so much that is not 

useful, the problem is how do you get 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. The problem is 

that the support system is not varied (most mines will find a 

support that works and stick with it or make small changes in 

support capability not large ones), and the mine loading conditions 

are not controlled enough to develop a curve. The result is that 

only bits and pieces of the ground reaction curve is ever developed 

from empirical data. 

However, the 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 (24) can be used to develop 

meaningful ground response curves. The software is able 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 tailgate, the ground response 

curve can be determined (25). Once this is done, more appropriate 

design criteria for roof support can be developed. 

The concept has been described in detail in other papers and 

will only be briefly summarized here to facilitate a further 

discussion of its significance in developing support design criteria 

(25). The ground response curve plots the support pressure against 

the convergence, as shown conceptually in figure 13. 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 self-supporting 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 

the chart. Equilibrium is achieved when the support curve 

intersects the ground reaction curve (point B). 

Support Pressure

Convergence

B

A

D

C

Q

P

Initial stress level

Ground response

curve

Support line

Support Pressure 

Convergence 

B 

A 

D 

C 

Q 

P 

Initial stress level Initial stress level 

Ground response

curve

Ground response 

curve 

Support line Support line 

d 

Figure 13 – Conceptual illustration of ground reaction curve. 

One example for a longwall tailgate condition is shown in 

figure 14. Here, a typical three-entry tailgate representative of a 

Pittsburgh seam mine geology is modeled (25). The ground 

response curve was developed by simulating a uniform support 

pressure on the roof and floor of the tailgate entry, while 

sequentially modeling the four external loading stages. Internal 

pressures of 0.01 tons/ft2

 up to 26 tons/ft2 were applied to provide a 

range of results that would bracket the typical range of standing 

support capacities. Ground response curves were developed by 

plotting the support pressure against the convergence for each 

loading stage. The model therefore produced four ground response 

curves, one for each loading stage. The results show that, up to the 

face abutment loading stage, the ground response curves are steep 

and convergence is limited to less than about 3 inches. The curves 

also show that the convergence is not dramatically affected by the 

amount of support. This indicates that this convergence is largely 

uncontrollable from the perspective of the standing support. The 

inby situation results in larger convergence values, more than 

4 inches for 26 tons/ft2 support resistance and apparent collapse is 

indicated when the support resistance is less than 0.26 tons/ft2

. 

Also plotted on this curve are the support responses for two rows of 

conventional 4-pt wood cribs and a single row of 30-in-dia 

pumpable supports on an 8-ft spacing. As seen from the 

intersections of the support response with the various ground 

reactions curves, it is concluded that the there is little impact of the 

stiffer, higher capacity pumpable crib outby the longwall face since 

most of this convergence is uncontrollable. Inby the face, the 

pumpable crib yields and provides about the same degree of control 

98 ft inby the face as the wood crib in this particular model. From 

this perspective, there is no clear advantage of the pumpable crib 

over the wood crib. It should be noted that this model is still 

considered a research effort. There are issues pertaining to the twodimensional aspects of the stress conditions and localized failures 

around the support itself that are fully addressed in this example. 

Additional research is planned in this area as part of the NIOSH 

ground control program. 
 . SUPPORT PRESSURE, ton/ft 2 

100 

10 

1 

0.1 

0.01 

0 2 4 6 8 1 0 1 2 1 4 1 6

CONVERGENCE, in 

Development 

Side abutment 

Face abutment 

33 ft Inby 

98 ft Inby 

197 ft Inby 

Full extraction 

Pumpable crib 

Wood crib 

Figure 14 – Example of a ground reaction curve for longwall 

tailgate developed using 2D Flac numerical model. 

Can the ground reaction concept be applied to shield supports? 

The information shown in figure 3 that was previously presented to 

discuss the impact of setting pressures is actually a ground reaction 

curve. The curve not only shows that an optimum setting load 

exists due to the rapidly diminishing impact of setting loads as they 

increase beyond the inflection point in the curve in figure 3; but 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. The loading produced by this activity 

must be considered in defining the yield load capacity of the shield. 

Terry Medhurst has used the ground reaction concept to 

evaluate design issues for shield support in Australian longwall 

mines (22). Medhurst used the ground reaction concept to evaluate 

the impact of support settting pressure, hydraulic leakage of leg 

pressures, support capacity, extraction height, and extraction depth. 

Conclusions and comments drawn from his work are summarized 

as follows: 

 Longwall mining and the resulting support response is 

largely dictated by convergence - controlled loading. – I 

completely agree and would take this one step further. As 

discussed before, what this is saying is that the roof behavior 

(particularly the overburden response) is governed by the 

stiffness of the supporting elements that include the coal seam, 

and the gob. Since the shield does not control the stiffness or 

properties of these elements, the shield, regardless of its 

capacity, is being loaded in proportion to its stiffness. In other 

words, this convergence is uncontrollable from the perspective 

of the shield design. 

 The successful application of longwall support technology in 

recent years might be partly due to close matching of the 

support stiffness with the coal seam stiffness. – While I 

would like to believe that this was by design, both Medhurst 

and I believe it is far more coincidental than it is by design. 

His point is that 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 example, a 10-ft seam would produce 

about 0.5 inches of convergence from the abutment loading 

before it fails, indicating 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. This reasoning would also imply that the 

shield stiffness by design should be adjusted as a function the 

extraction thickness and abutment loading (i.e. depth of cover), 

neither of which are done. 

Figure 15 – Impact of depth of cover and extraction height on 

longwall face ground reaction curve (from Medurst). 

 Higher extraction heights require higher capacity shields. – 

As shown in figure 15, the ground reaction curve shifts to the 

right and slightly upward as the extraction height increases. At 

higher extraction heights, the coal seam stiffness is reduced; 

additional face spalling is likely which increases the roof beam 

deflection and the roof span supported by the shield. One other 

point, which is not shown in Medhurst’s curves, is that the 

shield stiffness is also reduced with increasing extraction 

height. 

 Setting pressures are critical in weak roof conditions to 

preserve confinement of the roof beam. – It is here where we 

hope that the shield can provide meaningful control. The goal 

is that the shield pressure will reduce the convergence of the 

immediate roof beam, thereby helping avoid loss of 

confinement that may result in the formation of roof cavities. 

As Medhurst correctly points out, the impact of the support 

pressure decreases fairly rapidly as the distance from the 

canopy into the roof increases (he claims the effective zone is 

about 3 ft). Hereafter, the stability is dependent on the selfconfining effects of the roof strata. From a ground reaction 

perspective, lower shield capacity and/or stiffness, can allow 

increased convergence, thereby increasing the likelihood of 

failure of the immediate roof beam. 
 Leaking cylinders result in more convergence and poorer 

roof conditions. – 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. Get to the bottom of the ground reaction 

curve or miss the curve, and guess what --- trouble. 

While Medhurst has done well to incorporate the concept of the 

ground reaction curve into shield design analysis, I think the 

concept can be still be improved by separating the global 

conditions from the more localized conditions that occur on a 

shield by shield basis. As Medhurst correctly points out, the 

ground reaction curve moves to the right for increasing depth of 

cover and increasing extraction heights, two global factors that 

influences ground response and shield loading. 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 16. 

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 the 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 the shield stiffness (point C). The 

conditions remain nearly static once the shearer leaves the vicinity 

of the local area. Then, as the shearer approaches 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 17, 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. 

C

+ 

Support 

loading E 

E 

C 

D Support 

B Support 

lowering 

B 

F 

setting 

A 

Support 

advance 

F 

A A

adva

Shearer cu

GR

djacent shield 

ance GRC Curve 

C cu

cut 

rve 

Condition a

setting 

at 

ad

Adjacent sh

dvance GRC C

ield 

Curve 

SHIELD LOAD -

- CONVERGENCE + 

Figure 16 – Cycle loading pattern on ground reaction curve for 

individual or small group of shields. 

Equilibrium not established –

ground conditions unstable

-

+ SHIELD LOAD

E 

D 

B C 

Equilibrium not established – 

ground conditions unstable A

adva

djacent shield 

ance GRC Curve 

A F Shearer cu

GRC cu

cut 

rve 

Condition a

setting 

at 

ad

Adjacent sh

dvance GRC C

ield 

Curve 

- CONVERGENCE + 

Figure 17 – Equilibrium is not established in this shield loading 

cycle due to inadequate shield setting force. 

This type of analysis may also be 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 though 

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 may will be shifting to the right, and thereby cause a 

dramatic increase in convergence required to reestablish 

equilibrium and substantially reduce the number of yielding events 

that will create poor ground conditions. 

SUMMARY AND CONCLUSIONS 

The full potential of longwall mining was not realized until 

powered roof supports, particularly shield support systems, were 

developed. The stability of the shield permitted the application of 

longwall mining in conditions where caving was difficult to control 

and where chock and frame supports were inadequate. 

Shield capacities continue to increase as a “bigger-the-better” 

philosophy prevails in longwall mining throughout the world. 

Several 1,280-ton shield systems are now in operation in U.S. coal 
 

 

 

 

 

 

 

 

mines. Several models have been developed based on theoretical 

and empirical support and strata interactions to determine shield 

capacity requirements. However, it is difficult to justify the current 

capacities strictly in terms of ground control requirements. 

Increasing the life expectancy of the shield has become a driving 

factor in current shield design and capacity considerations. 

The increase in shield capacity has resulted in higher setting 

forces. Historically, mine operators have welcomed the higher 

setting force capability under the premise that it will help to 

strengthen the immediate roof beam by increasing the frictional 

restraint between bedded layers, and provide additional active 

horizontal loading to provide further confinement to the face-towaste strata movements that lead to roof falls in front of the shields. 

However, rock mechanics theory and limited underground shield 

loading data suggest that there is an optimum setting pressure that 

state-of-the-art high capacity shields may have exceeded. More 

research needs to be conducted to determine optimum setting 

pressures in order to provide the most effective ground control in 

various mine geologies while maximizing the life of the shield. 

An often overlooked consequence of the increase in shield 

capacity is an increase in shield stiffness. If there is a strata control 

benefit in using higher capacity shields, it may well be due to their 

increased stiffness more so than their increased capacity. By 

developing the supporting forces more quickly in relation to the 

face convergence, the stiffer shields may reduce separation of 

immediate strata layers and minimize movement of the rock along 

fracture planes that lead to premature roof failures. Conversely, 

when the face convergence is irresistible, as is the case with main 

roof and overburden loading, the increased shield stiffness will 

result in higher shield loading. Since the stiffer shields will use up 

their available capacity as a passive support in irresistible strata 

movements more quickly, they will be fully loaded to their yield 

capacity as frequently as the less stiff, lower capacity shields they 

have replaced. In this scenario, one should not expect an 

improvement in the life expectancy of the high capacity shield over 

that of the previous generation of supports. This provides further 

justification for optimizing the setting forces, since the final 

support load will largely be dependent on the setting loading. 

Significant advancements have been made in standing roof 

support technology for longwall mining. Over 50 new support 

concepts are now available for longwall tailgate support. These 

new support products have shown to provide superior roof control 

as well as material handling advantages that ease installation 

requirements. Full-scale testing in the NIOSH Mine Roof 

Simulator has provided accurate information on the performance 

characteristics and limitations of these support products. The 

NIOSH STOP software has helped to eliminate trial and error 

approaches to application of these support products by providing 

performance and design information in an easy to use design tool 

for standing support. 

Yet, with all these improvements in roof support technology, 

there is still a poor understanding of exactly how the support 

interacts with the strata to provide ground control and how much 

control it actually provides in this process. The ground reaction 

curve is a means to help answer these questions and provide a 

better engineering design approach to ground control than can lead 

to optimization of support designs and application strategies. 

Numerical modeling has matured to the point where it can provide 

the means to more easily evaluate ground reaction implications 

than are possible through empirical in mine studies, but the models 

need to be calibrated with field data and more of these studies need 

to be done. The procedure can be applied to all support, including 

longwall shields to help answer the question about how much 

support is really needed. The “bigger-the-better” design 

philosophy cannot last forever and is already based on false 

pretenses. Actually, obtaining a ground reaction curve for longwall 

faces is as easy as it gets, simply lower the setting pressures in a 

controlled manner and monitor the change in loading which can be 

back calculated to convergence through the shield stiffness and you 

have a ground reaction curve. Hey, lets go try it! 

REFERENCES 

1. Peele, R. Mining Engineering Handbook, Third Edition, John 

Wiley and Sons Inc. Copyright 1941, Vol. 1, pp. 10-5-5 to 

10-511. 

2. Mason, R.H. The First Longwalls Were Primitive Compared 

to the Ones Now in Use. Coal Mining and Processing. 

December 1976, pp. 59-64. 

3. U.S. Longwall Census – Population Pushes Past 50, Coal Age, 

February 2005, pp. 26 - 30. 

4.	 Adam, R.A., Barczak, T.M. and Irresberger. Strata Dynamics 

in Underground Coal Mines, Longwall USA, Pittsburgh PA, 

June 1991. 

5.	 Wilson, A.H. Support Load Requirements on Longwall Faces. 

The Mining Engineer, 1975, pp. 479 - 491. 

6.	 Peng, S.S. and Chiang, H.S. Longwall Mining. John Wiley 

and Sons, Copyright 1984, p. 185. 

7.	 Gupta, R.N. The Influence of Setting Pressure on Support 

Performance and Strata Behavior on Longwall Faces. Ph.D. 

Dissertation. University of New Castle upon Tyne, U.K., 

March 1982, 576 pp. 

8.	 Barczak, T.M. and Oyler, D.C. A Model of Shield - Strata 

Interaction and Its Impact on Active Shield Setting 

Requirements. Tenth International Conference on Ground 

Control in Mining, Morgantown WV, June 1991. 

9.	 Smart, B.G.D. The Evaluation of Powered Support 

Performance from Geological and Mining Practice 

Information@. Proceedings: 27th U.S. Rock Mechanics 

Symposium, Alabama, 1986, pp. 367 - 377. 

10.	 Barczak, T.M. Design and Operation of Powered Roof 

Support Systems for Longwall Mining. U.S. Department of 

Interior, Bureau of Mines, IC 9320, 1992, 14 p. 

11.	 Barczak, T.M. and Schwemmer, D.E. Stiffness 

Characteristics of Longwall Shields. U.S. Department of 

Interior, Bureau of Mines, RI 9154, 1988, 14 pp. 

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