DESIGN CONSIDERATIONS FOR 


THE NEXT GENERATION OF 


LONGWALL SHIELDS
 

Thomas M. Barczak 

National Institute for Occupational 

Safety and Health 

Pittsburgh Research Laboratory
 
 

ABSTRACT 

Advancements in support technology have occurred to a large extent because of the demands of 

the mine operators. Shield supports are no exception. The most notable changes in shield 

technology in the past decade have been increases in shield capacity and life expectancy, which 

have increased by a factor of 4 and 7 respectively since the inception of the shield support 25 

years ago. The purpose of this paper is to set goals for the future generations of longwall shields. 

These goals include: (1) 100,000 cycle life expectancy, (2) improved hydraulic diagnostic 

capability, (3) smart load control with adjustable setting pressure algorithms, (4) composite 

material design to reduce weight and increase shield width, (5) constant set leg cylinder design, 

(6) improved joint design to reduce wear and stress, (7) advancements in leg cylinder socket 

design, and (8) an instrumented shield to measure in-service structural component loading. These 

design issues will continue to push the envelope in longwall shield technology that has provided 

the safest and most productive mining system employed in the history of mining. In preparation 

for this, a historical examination of shield failures is made and recent trends in shield design are 

evaluated. Design practices to improve structural margins of current designs are also addressed. 

INTRODUCTION 

Reviewing the 25 year history of the use of the shield support for longwall mining in the United 

States, it is clear that the shield had a major role in making longwall mining the most productive 

and safest mining method available to U.S. operators. The basic shield structure has remained 

unchanged for the past 25 years, although the structures have grown dramatically in size and 

capacity. Early generations of shields experienced several structurally related failures and had to 

be strengthened to prevent premature failures and provide a reasonable working life. Through 

this evolution of improvements, the life expectancy has been increased by a factor of 7 from 

10,000 loading cycles in the late 1970's to 70,000 loading cycles today. Apart from technological 

improvements in electro-hydraulic control systems, which have impacted the operational 

capability of the support, there have been relatively few advances in shield design in the past 10 

years. 

The purpose of this paper is to revisit shield technology and to propose some considerations for 

further enhancements in shield design. The scope of these ideas include both short terms 

solutions that can be accomplished with current technology as well as bigger goals that may 

require some additional research. Nonetheless, these concepts can provide for a new generation 

of shields, which will last longer, provide more intelligent use of the available support capacity, 

and improve structural margins of safety to ensure that premature failures do not occur. 

Reviewing the design history of longwall shields, it is apparent that the U.S. operators have had a 

major impact on the design improvements that have occurred. Dating back to the initial longwall 

installations, where the integrity and capacity of the European shields were not up to the 

standards, the U.S. operators insisted on higher capacity and more reliable structural performance, 
and then continued to demand greater working life as the shields evolved, the U.S. mine operators 

are clearly the motivating force behind the improvements that have been made in the past. It is 

just as likely that any future improvements in shield design, such as those proposed in this paper, 

are more likely to occur if the U.S. operators push for them. 

RECENT TRENDS IN SHIELD DESIGN 

Material Specifications – Shields in recent years have been fabricated from high strength steels 

(100,000+ psi yield) to minimize component cross sectional dimensions. High strength steel 

applications were first used in canopy structures to minimize the cross sectional thickness of the 

canopy in low to moderate seam applications. Today, high strength steel in all shield designs is 

necessary to provide the required strength for various shield components in high capacity designs. 

While the added strength has helped to extend shield life, the high strength fabrications are more 

susceptible to brittle failure and catastrophic fatigue failures. Another 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 for the past 15 years is shown in figure 1 (Coal Age, 1999). Average support 

capacities in the United States have increased by nearly 50 pct since 1985 to an average support 

capacity at yield of 768 tons for the 62 operating longwalls in the U.S. in 1999. Twelve 

installations (19 pct) of the current longwalls employ shields with capacities greater than 900 

tons, and 19 installations (50 pct) have capacities between 700 and 900 tons. The current 

distribution of shield capacities is shown in figure 2. The highest capacity shield used in the 

United States is 1,170 tons (Coal Age, 1999). As shown in figure 1, maximum shield capacities 

have evolved from 800-ton shields which were common from 1985 to 1990 to the current 1,200-

ton shields 1999. 

1,200 

AVG SUPPORT CAPACITY 1,100 

MAX SUPPORT CAPACITY 

1,000 

900 

800 

700 

600 

500 

Figure 1 Trend showing increase in support capacity SHIELD CAPACITY, tons

1985 1987 1989 1991 1993 1995 1997 1999
 

YEAR
 
DISTRIBUTION, pct 

60 

50 

40 

30 

20 

10 

0 

<500 500 - 699 700 - 899 >900 

SHIELD CAPACITY, tons 

Figure 2 Distribution of capacity for shield operating in 1999 

The increase in capacity has significantly impacted shield design and ground control capability. 

In order to provide the increased support capacity, larger diameter leg cylinders had to be utilized. 

One consequence of the larger diameter leg cylinders has been a proportional increase in shield 

stiffness. The increase in shield stiffness means that the load developed in the support structure 

will occur at less displacement or face convergence. While the increased stiffness may provide 

superior control of the immediate roof, main roof weighting that causes irresistible (in terms of 

the shield capacity) convergence on the longwall face will result in greater shield loading with the 

stiffer shield design. As a result, it is not uncommon for the high capacity shields to be fully 

loaded to yield capacity just as often as lower capacity designs that they replaced. An example is 

illustrated in figure 3 where a 500, 800, and 1,000-ton shield are all fully loaded when 0.5 inches 

of convergence occurs. 

0 

200 

400 

600 

800 

1,000 

1,200 

500 Ton 800 Ton 1000 Ton 

Support load, tons 

Setting load Convergence load 

Figure 3 Setting load and load developed due to roof movement for shields of various capacity 
 

Another consequence of the larger diameter leg cylinders is that the leg socket must be designed 

to accept greater loads. Historically the leg socket has been a source for premature shield failures 

and the increased loading makes the design even more demanding. In addition, the larger 

diameter leg cylinders require a wider and longer socket, which places further demands on the 

design of the socket and load transfer to the base and canopy structures. In general, the larger 

diameter cylinders have resulted in wider support components which are more susceptible to 

torsion loading than before. 

Several models have been developed over the years to determine support capacity requirements. 

While these concepts attempt to capture the support and strata interaction principles, state-of-theart shield capacities cannot be justified based on these concepts. A “bigger the better” attitude has 

prevailed, being promoted by the manufacturers largely due to the demands of mine operators to 

improve the life expectancy of the shields. Hence, it is this issue (life expectancy) more than the 

ground control requirements that have controlled recent developments in longwall shield design 

and capacity determinations. 

Type of Shield – There has been a steady increase in the use of two-leg shields in favor of fourleg shields during the past decade, and two-leg shields are becoming the favored support 

worldwide. In 1997, 63 of the 65 longwall faces in the U.S. were two-leg shield systems, 

compared to only 53 pct in 1985. Larger size hydraulic cylinders have been developed in recent 

years to accommodate the increased demands for higher shield capacities, and have allowed these 

capacities to be realized with two-leg designs that were not possible 10 years ago. 

There are several consequences of the two-leg design. From a support strata interaction 

perspective, the two-leg shield provides an active horizontal force toward the coal face due to 

inclination of the leg cylinders. This active horizontal force improves overall strata stability by 

arresting slippage along fracture planes or by prevention the expansion of highly jointed or friable 

immediate roof geologies which may be further damaged by the front abutment loading. In terms 

of the shield loading, this increase in active horizontal loading also translates into proportionally 

higher lemniscate link loading. Historically, lemniscate link pins and wear in bores has been a 

primary cause of premature shield retirement and/or rebuild. Hence, pin diameters and bore areas 

need to be increased in higher capacity shields to prevent this problem from becoming worse. 

Another issue related to the two-leg concept is higher contact pressure on the canopy and base. 

High toe loading, caused by the moment created by the line of action of the resultant vertical 

forces acting on the canopy and base, can be a problem in high capacity two-leg shields and 

should be a consideration in the support design. Base toe pressures of 800 psi or greater can be 

expected on high-capacity two-leg shields. Base toe lifting devices are now standard on most 

two-leg shields to assist in the advancement of the shields particularly in soft floor conditions. 

There has also been a trend toward solid base designs to reduce floor-bearing pressures in two-leg 

shields. 

Size – In addition to wider shields required by larger diameter leg cylinders, there is a trend 

toward wider shields to minimize hydraulic cylinder maintenance and reduce the total cost by 

employing less leg cylinders on a longwall face. Again the issue of torsion is important in 
designing wider shields. It is not as easy as simply extending the width of the canopy and base. 

These components also need to be strengthened for torsional loading. There is also the issue of 

weight. The 2.0 m wide designs may represent an upper limit with current shield construction 

materials. If high capacity designs are to prevail, then lighter weight materials such as composites 

are likely to be needed to develop widths much beyond 2.0 meters. 

In addition to increases in width, shields have increased in length to accommodate one-web-back 

operations and larger face conveyors and deeper shearer webs. Longer canopies and bases create 

much larger bending moments that require stiffer and stronger components than in previous 

generation supports. The increase in length is largely responsible for the need for greater shield 

capacity as the area of roof loading carried by the shield increases or the greater convergence is 

seen by the cantilevering of the immediate roof beam as the resultant shield force moves further 

from the coal face. 

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 4,000 to 4,200 psi 

range. This design practice coupled with the increased stiffness of the higher capacity supports 

means that the higher capacity shields are fully loaded as often as their lower capacity 

predecessors (figure 3). 

Hydraulic Components and Control Systems – Both manufacturers continue to make 

improvements in the electro-hydraulic control systems, making them more reliable, more userfriendly, and easier to diagnose when problems do occur. A description of these systems can be 

found in a paper entitled “Shield Design, Construction, and Operation” published in the 

proceedings of Longwall USA (Barczak et al., 1998). Solenoid-operated valving systems are now 

becoming standard. Spool valves have been shown to be superior to ball and seat designs, which 

are prone to contamination problems. In addition, these systems allow the solenoid to be 

activated upon demand, unlike previous systems, which required the hydraulic feed to be 

interrupted by a control solenoid. This leads to both quicker and smoother control of support 

functions. 

Hydraulic Emulsions – For many years, longwall shields have utilized a water/mineral oil 

emulsion as a hydraulic medium for the leg cylinders. The standard system has been a 5 pct 

oil/water emulsion. There is trend towards the use of synthetic fluids. Most western mines have 

now switched to “low treatment” systems with synthetic oils in concentrations of only 1 to 2 pct. 

The motivation for this has largely been due to environmental issues imposed by the Utah 

Department of Natural Resources. Only one eastern mine is currently using the “low treatment” 

emulsion system. Fazos (Australia) has experimented with an all-water system. 

Although the synthetic oils are environmentally preferred, they also cost significantly more. The 

synthetic concentrate is about 3 to 5 times the cost of mineral oils. Hence, despite the lower 

concentration used in the “low treatment” systems where less than 2 pct oil is utilized, the overall 

cost is typically about 50 pct greater than “high treatment” systems using 4 to 5 pct mineral oil 
concentrations. The major disadvantages of the synthetic “low treatment” system is that there is 

little room for error. A small drop in the oil concentration can lead to lubrication and acidity 

problems. Therefore, maintenance of the oil/emulsion is much more critical than in the “high 

treatment” systems, where the oil content can be reduced from 5 to 4 pct with little if any 

detrimental effects. Bacteria growth can also be accelerated in very low concentrations of oil 

emulsions, which can cause more severe corrosion problems than if there were no oil at all. 

SUMMARY OF FAILURES EXPERIENCED 
 BY MINE OPERATORS 


There have been numerous shield failures throughout the history of the shield support. A 

summary of these is provided below. While failures have declined, particularly premature 

failures that occur early in the shield life, there are still isolated cases of premature failures. 

Shields have a finite life and fatigue failures will eventually occur on all shields if left in service 

long enough. 

Historical Overview of Shield Failures 

Base Failures - Base failures seem to be the most prevalent and usually occur from fatigue after 

the support has been in operation for several panels of extraction. A common failure mechanism 

is for the leg socket casting to break away from the base structure. Formation of this failure is 

difficult to detect while the support is in-service, as the leg socket is housed deep inside the base 

structure and this area usually is full of debris. Once the leg socket breaks loose, the support 

quickly becomes inoperable. The bottom plates of the base have insufficient strength to withstand 

the leg forces and the leg cylinder will literally rip the base apart by tearing off the bottom plate. 

Failure of the base structure (plates) can also occur without the leg sockets failing. The probable 

failure mechanism is bending of the base. This is more likely to occur in mine sites which have 

very strong immediate floor strata. In these hard floor conditions, steps in the floor may be left by 

the shearer as it is difficult to maintain a constant height of extraction from cut to cut. The base 

structure is then simply supported in two locations and is flexed as loading is applied. Repeated 

flexure causes the base to deform (plastically) or promotes fracture from fatigue which eventually 

results in failure of the base structure. In softer floor conditions, the strata deforms to provide a 

fuller contact to the base. This alleviates much of the bending and reduces the risk of failure. 

Standing the support on the toe of the base can also result in damage of the base structure. This 

configuration causes maximum stresses in the toe region and the base will deform or fail usually 

where the cross section is a minimum in the section of the base forward of the leg connection. 

Internally, the base structure is constructed with stiffeners which hold the top and bottom plates 

apart to form a beam arrangement which gives the base its bending strength. Cases have been 

reported where these stiffeners have not been properly welded in place or where the dimensional 

tolerances were not within specifications. In these cases, the stiffeners broke loose and the base 

structures literally collapsed. This problem appears to be largely a matter of quality control, but it 
is critical to support safety. Since the stiffeners are hidden inside the base structure, it is virtually 

impossible (excluding X-ray inspection) to see these deficiencies prior to the failure. 

Canopy Failures - Canopy structures are constructed of stiffened (top and bottom) plates similar 

to base structures, and hence they are susceptible to bending-induced failures as well. 

Structurally, canopies are less stiff than bases making them more susceptible to failure from 

bending than base structures. However, while permanent deformation of the canopy is a fairly 

common occurrence, destructions of canopies appear to be less frequent than observed 

destructions of bases. This suggests that canopies are less often subjected to critical bending. 

Three reasons why canopies might avoid critical loading are: (1) immediate roof strata is usually 

partially fractured and full contact with the canopy is more easily obtained which minimizes 

bending moment; (2) tip loading on the canopy is usually smaller than toe loading on the base, 

since the resultant force is more likely to be located near the toe of the base than the near the tip 

of the canopy; and (3) the canopy surface area is larger than the base area which allows the 

canopy to distribute load more efficiently. 

Another common deformation of the canopy is "wrinkling" of the top plate between the internal 

stiffeners. This is due probably to concentrated loading at locations between the stiffeners, but 

might also be an indication of failure of the weldments that hold the stiffeners in place. If the 

stiffeners are not secure, the plate may buckle from excessive stress that results in bending of the 

plate between the more rigid stiffeners. 

Caving Shield And Links - Link members have become considerably more robust in shield 

designs during the past ten years, and failures have been substantially reduced. Since the caving 

shield-link assembly has very little vertical load capacity (stiffness), links are not highly stressed 

for most load conditions. Almost all link failures can be attributed to conditions or operating 

practices which promote standing the support on the toe of the base or conditions which cause 

large horizontal displacement of the canopy relative to the base. 

Failure of the structure is most likely to occur in the region near the (pin) hole located on each end 

of the member. The failure mechanism is most likely crack formation somewhere on the 

circumference of the hole from localized high stress development. The holes elongate from 

continued wear and contact with the higher strength link pins. This results in point loading of the 

pins and high stress development at the contact areas. These failures are difficult to detect since 

this area is obscured from view by the caving shield clevises or base structures. Although link 

failures are rare, they can be catastrophic as the links provide horizontal stability to the support 

structure. 

Likewise, caving shield failures are fairly rare, but are more likely to occur than link failures. 

While links are designed primarily for axial loading only, shield mechanics indicate the primary 

loading mechanism for the caving shield is bending and torsion. Maximum stresses and failure 

are most likely to occur in the clevis areas where pins connect the link members to the caving 

shield. Some general yielding by bending deformation of the caving shield structure may also 

occur. 
 

Leg Cylinders - Assuming the face area is sufficiently stable to prevent violent outbursts of 

energy (bumps), it is unlikely that leg cylinders will experience structural failure since they are 

designed to control loading by hydraulically yielding at specified pressures. The most common 

failure associated with leg cylinders is seal leakage. 

Another potential failure mechanism for hydraulic leg cylinders is for the yield valves to 

malfunction, allowing leg pressure to increase beyond design levels. Usually, the excessive 

pressure will cause seal leakage, so that it is unlikely that sufficient pressure will build up to 

rupture the cylinder casing. The more common problem is leakage through the yield valve which 

causes unplanned pressure losses during normal loading. This leakage can be caused by dirt or 

contamination not allowing the valve to seat properly or worn seats, defective springs that 

maintain the valve in the closed position, or the fitting itself may leak due to bad O-rings or seals. 

FAILURE ASSESSMENT OF SHIELDS 
 CURRENTLY IN SERVICE
 

The age distribution of the shields operating in 1999 is shown in figure 4. The range of operating 

life extends from 0 to 14 years of service. An average shield operates about 4,000 cycles per 

year, translating into a face production of approximately 2.5 million tons per year. Hence, an 

average 10-year-old shield would have 40,000 loading cycles, while a shield employed on a high 

production longwall (4 million tons per year) would have about 65,000 loading cycles. 

Approximately one third of the shields have been in operation from 4 to 6 years (16,000 - 24,000 

cycles), and slightly more that than one-third (37%) have been operating for more than 6 years 

(24,000 loading cycles). 

0-3 4-6 7-9 9+ 

AGE, yrs 

Figure 4 Age distribution of shields operating in 1999 

0 

5 

10 

15 

20 

25 

30 

35 

PERCENT 
Figure 5 shows a near-linear increase in structural failures with age. Failure is defined as 

structural damage to any support component. It does not necessarily mean catastrophic failure 

that renders the support inoperable. The linear relationship suggests that there are more than 

just fatigue failures occurring, since the frequency of fatigue failures should increase at a much 

greater rate as the shield ages. There was a wide range of structural failures reported. Most of the 

major failures involved either the leg sockets or pin joint problems (i.e., lemniscate link pins and 

clevises). These two failures are most likely to put a shield out of service and/or require structural 

modification to keep the support in service. Other structural failures occurred at areas of high 

stress concentration including the side shield on the canopy structure, the canopy capsule tilt 

cylinder bracket, the base bridge, the link cut out areas in the caving shield, various mounting 

brackets including the pilot-operated/yield valve manifold, and dishing of canopy skin covering 

plate. There were a few isolated cases of leg cylinder casing problems, but no catastrophic 

failures. 

There were more hydraulic problems reported than structural problems (figure 5), 

particularly in shields that have been in operation less than 6 years. There were nearly 4 times the 

number of hydraulic failures in shields operating 3 years or less and nearly twice as many for 

shields in the 4 - 6 year range than there were structural failures for the same age group. 

Obviously, the weak link in shield design is the hydraulic system. Numerous hydraulic failures or 

problems were reported. The high frequency of failures in the first 3 years of operation is 

distorted by failure of a component called a hydrafuse. The hydrafuse is a safety device 

incorporated by Joy as part of the leg cylinder design on their recent generation of shield supports 

in response to a fatality that occurred recently in Australia. An Australian mineworker was killed 

when the hydraulic fluid in the retract annulus of the leg cylinders was inadvertently pressurized 

while setting the support. Normally, the retract port is open through the return line back to the 

hydraulic reservoir whenever the leg cylinder is being pressurized. This is necessary since the 

pressure would be highly intensified (by a factor 30 or more) if the port was blocked, as in the 

case of the Australian fatality. The resulting pressure intensification caused the cylinder end cap 

to rupture, striking and causing a fatal injury to the mineworker. 

0-3 4-6 7-9 9+ 

AGE, yrs 

Figure 5 Distribution of structural and hydraulic failures on shields as a function of age 

0 

10 

20 

30 

40 

50 

60 

70 

80 

PERCENT OF INSTALLATIONS 

STRUCTURAL 

HYDRAULIC 
The hydrafuse is incorporated into the retract circuit and acts as a yield valve to prevent the 

inadvertent build-up of pressure in the retract annulus of the leg cylinder. The design utilized by 

Joy features a brass clip that shears off at a predetermined load (hydraulic pressure). Once the 

clip is broken the “fuse” must be replaced. With a failure pressure of nearly twice the pump 

operating pressure, it was expected that the fuse would provide the desired safety on the rare 

occasions when there was abnormal pressure, and that replacement of the hydrafuse after the 

failure would not be much of an inconvenience. To the surprise of most everyone, hydrafuse 

failures began occurring on several shields at several mines. The failures seemed to occur 

whenever the support had set idle for a period of time, and whenever the leg pressure was at or 

near yield pressure and almost always on the top stage retract circuit. However, it is believed that 

a water-hammer effect is being created by the valve operation. Joy has since installed a restricter 

valve in the circuit to dampen out any potential pressure spike. Early indications from the field 

installations indicate that this may be working to prevent the undesirable failure of the hydrafuses. 

Unfortunately, the restrictor also increases the time required to lower the top stage. Fortunately, 

in two-stage leg cylinder designs, the bottom stage is lowered first and lowering of the top stage is 

only required when the bottom stage is fully collapsed (Barczak, 1998). Hence, the increased 

time to lower the top stage will not be a problem during most production mining. 

Other cylinder problems caused pressure losses that limited support capacity. One example was 

due to poor fabrication where the internal bore of the top stage was off center resulting in a 

weakened casing that was unable to sustain the pressure intensification that occurred in the top 

stage. Other problems were reported with defective yield valves that would not reset resulting in 

the inability of the shield to adequately hold load after yielding. Similar problems were reported 

with the staging valve which also caused loss of pressure in the leg cylinders at a few mines. One 

mine reported clearance problems with the leg cylinder and the base rib plates where the cylinder 

leaned into the rib plates causing internal damage to the cylinder and seal leakage. Failures also 

occurred with the advance ram cylinders including failure of shuttle valve springs and structural 

failure of a hollow tube relay bar design. 

Most hydraulic problems are related to internal leakage due to seal wear and/or corrosion of the 

cylinders. These problems typically begin when shields have been in service 4-6 yrs. As figure 5 

shows, 60 pct of the shields had some sort of hydraulic problems during this time frame. Many 

problems go undetected for extended periods of time, resulting in degraded support capacity that 

can contribute to ground control problems in heavy loading conditions. Methods to detect the 

onset of internal hydraulic leakages are discussed later in the paper. As previously described, 

there is a trend towards the use of low treatment synthetic fluids in western mines. One western 

mine reported severe leg cylinder corrosion due to problems with a low treatment synthetic 

emulsion on shields that have been in operation less than 3 years. 

Problems with the electro-hydraulic control systems were also reported. These included sticking 

solenoid valves (problem with soaping in emulsion formulation) and chattering valves due to fluid 

dynamics at high flows. Most mines reported that there have been significant improvements 
 

made in the latest generation of electro-hydraulic controls. In particular, the DBT (MTA) PM-4 

system seems to be a significant improvement over the previous generation PM-3 design. 

Design Practices That Improve Structural Margins of Safety

and Extend Shield Life 

Since most structural shield failures can be attributed to some form of fracture, the basic elements 

of fracture control can significantly improve shield life. These are: (1) use a lower design stress, 

(2) minimize stress concentrations, (3) reduce flaw size or control crack growth, (4) minimize 

corrosion, and (5) use materials of improved toughness (Barczak, 2000). 

Lower Design Stress - Some margin of safety should be employed in the design stress relative to 

the yield strength of the steel. Civil engineers typically use a factor of 1.66, meaning that the 

allowable stress is about 60 pct of the yield stress for non-fatigue loading and further reduced by 

50 pct or more when fatigue loading applies (AISC, 1980). While these levels of safety are not 

practical in shield design due primarily to cost and weight limitations, it is important to recognize 

that a small reduction in (tensile) stress developments will significantly reduce crack growth since 

the two are related by an exponential function. Past practices of designing to or near yield 

strength should be avoided in modern shield design where the life expectancy exceeds 50,000 

loading cycles. 

Link pins and clevises are a prime example of historically poor design practices in shield supports 

which continue even today. Deformation and/or excessive wear in the pin clevises are 

undoubtedly the primary cause of premature shield retirement and/or structural rebuild. There are 

clear indications that these areas are subjected to stress beyond the yield strength of the steel. 

This poor design is caused partly by manufacturers not giving sufficient credence to the load 

conditions which cause high loading in the caving shield - lemniscate assembly, namely loss of 

frictional contact at the roof and floor interface and standing the support on the toe of the base. 

With the possible exception of shields designed for low seam heights, there is adequate space 

available to increase the bearing area of these clevises and pin diameters to reduce the stress and 

substantially improve the life expectancy of these components. The joint design problem also 

needs to recognize the importance of pin tolerance. First, the pin contacts only a portion of the 

clevis. Typically, arcs of 45 to 60E are used in the design analysis. Obviously, the 45E arc 

assumption will lead to more conservative designs. Conservative assumptions should be made to 

allow for reduced areas as wear occurs (figure 6). Corrosion effects are also often ignored or not 

sufficiently accounted for in the design of pins, despite the fact that corrosion is a leading cause of 

premature pin failures or abnormal wear. Corrosion causes pits to occur in both the pins and the 

clevises (figure 7) which can reduce the bearing area by 25 to 50 pct and cause a proportional 

increase in stress. 
Figure 6 Reduction in lemniscate pin 

contact area for worn joint Figure 7 Corrosion of a lemniscate link 

clevis 

Several failures have occurred where the leg socket casting weldments break due to fatigue or 

stress-corrosion, and thereby cause subsequent failure of the bottom base plate as the loading is 

transferred fully to the bottom plate instead of being distributed to the side rib plates. Most of the 

time the leg cylinder and casting punch through the bottom plate and into the mine floor rendering 

the support inoperable. A common modification to alleviate base leg socket failures of this nature 

is to add another plate to the under side of the base pontoon. Typically, this plate is about 1 inch 

thick and usually covers most of the length of the base pontoon. Reinforcement is also typically 

added to the top area of the side base plate. While this plate stiffens the side plate, its primary 

purpose is to restore the location of the centroidal axis, which was changed by the addition of the 

bottom plate, to its location in the original base design. The stiffness of the base should be fully 

examined during the initial design phase of the shield in order to prevent these failures from 

occurring. 

Some mines have successfully reduced operating stress levels by derating the shield support 

prior to underground installation. The derating is accomplished by installing yield valves with a 

lower operating pressure than specified in the design. Since the leg capacity controls the 

maximum load developments within the support, lower leg loads translates into reduced 

component loading. For example, if a 1,000-ton shield support is derated by 10 pct to a 900-ton 

capacity, the margin of safety relative to the tolerable crack size that will prevent fatigue fracture 

may increase by 20 to 30 pct. 

Minimizing Stress Concentrations - There are numerous sources of stress concentrations in 

longwall shield supports. The most common is a change in geometry. These stress 

concentrations should be identified and their magnitudes quantified during the design and 

performance-testing phase. One way to do this is to use photo-elastic plastics. The photo-elastic 

plastic can be applied to almost any area of the shield structure. Colored fringes will appear on 

the plastic when observed through a polarized lens that correlates to the stress profiles. Stress 

intensity factors of 2-3 are not uncommon for sharp changes in geometry such as holes or sharp 

bends in structural members. 
One example of a sharp change in geometry is a lemniscate link design with offset pinholes as 

shown in figure 8. This link design is typically employed on a shield with a low profile designed 

for operation in low coal seams. The bend is necessary to provide clearance with the caving 

shield in the collapsed or low operating height. Lemniscate links are primarily axially loaded 

members, but the offset pinhole geometry induces additional stresses due to bending, and thereby 

significantly reduces the margin of safety for this component. Figure 8 depicts failure of a bottom 

lemniscate link on a 620-ton Westfalia shield that occurred during performance testing at the 

NIOSH Safety Structures Testing Laboratory. Although the shield had 45,000 load cycles from 

underground service prior to the testing, a new link was installed on the shield that was 

performance tested in the laboratory. This was a new link design that was fabricated for the mine 

by an outside vendor (not the support manufacturer). Failure occurred after only 14,000 loading 

cycles. This failure illustrates two problems. First, the design (allowable stress) was too high. 

Test results revealed that the nominal stress in the link exceeded 90 pct of the material yield 

strength. In addition, it appeared that the link side plates had been torch cut, adding an additional 

stress riser to the already sharp change in geometry at the bend in the link in the area where the 

failure occurred. These two factors resulted in an unacceptable time to failure. Figure 9 

illustrates another problem where the fabrication process left large flaws that led to premature 

fatigue failure of the base rib. This problem may have been alleviated if the surface was 

smoothed to remove most of the surface flaws. 

Figure 8 Failure of bottom lemniscate link Figure 9 Failure of base rib side plate where 

fabrication process left large flaws 

Any hole in a structural plate is another area where stress is concentrated. The structural 

components of a longwall shield (canopy, caving shield, lemniscate links, and base) are connected 

by a pin and clevis arrangement. These areas are also sources of stress concentration and fatigue 

failures on aging longwall shields (figure 10). Another example of a stress concentration caused 

by too sharp of a change in geometry is shown in figure 11. Holes are sometimes cut into the 

canopy or caving shield structure to accommodate the placement of hydraulic hoses. In the 

example shown in figure 11, failure occurred at the stress riser caused by the sharp corner in a 

cut-out on the caving shield made to accommodate hosing for the side shield. The best solution to 

this problem is to avoid the hole altogether, and if the hole is necessary, to make sure that the 

corner radius is as large as possible. Another example of failure due to stress risers created by 

sharp geometries is shown in figure 12, where a crack developed in the caving shield near the 

canopy hinge. A clean-out hole is often placed in the side of the base structure to facilitate 
removal of debris from the leg socket area. This too can be a source for concentration of stress. 

Since this is a critically loaded area of the base, care should be taken in the design to minimize the 

stress concentration by a incorporating a favorable geometry and orientation with respect to the 

stress field in this member. 

Figure 10. Failure in base lemniscate link clevis Figure 11 Failure due to stress concentrations 

in cut out sections of caving shield 

Figure 12 Failures of caving shield due to stress concentration in 

sharp corners where lemniscate links connect 

Figure 13 shows failure of a canopy leg socket casting due to a stress riser created by a sharp 

change in geometry. The failure in this case occurred in the casting itself. This failure occurred 

on several shield supports in the late 1980's, all of which utilized this same basic socket design. 

Again, this failure probably could have been prevented by a smoother geometry. Leg sockets are 

often a source of fatigue failures in aging longwall shields. The leg socket is a critical area since 

the full load developed within the hydraulic leg cylinder must be transferred into the canopy and 

base structures to be distributed to the mine roof and floor. An examination of the structural 

mechanics associated with these base socket failures has led to some other design changes that are 

intended to reduce the stress concentration in the leg socket casting/base structure connection. 
The rectangular geometry of the socket casting and its placement between the side rib plates of 

the base structure position it at right angles to the principal tensile stress caused by the bending of 

the base structure. This orientation creates a stress intensity factor which helps to promote 

fatigue-induced fracturing of the weldment. Some shields are now being designed with more of 

an elliptical or zipper-shaped casting (figure 14) so that the front and/or rear edge is not 

perpendicular to the principal tensile stress, thereby resulting in a reduction of the stress intensity 

factor. 

Figure 13 Failure due to stress riser 

created by sharp corner in 

canopy leg socket casting 

Fig. 14 “Zipper-shaped” leg cylinder casting to 

avoid stress concentration caused by 

orientation of casting to principal 

stress field 

Reducing Flaw Size or Controlling Crack Growth - Weldments are an essential part of the shield 

fabrication process. However, since weldments are a primary source of structural flaws, the 

quality of the welds is critical. As the principles of fracture mechanics illustrate, the initial flaw 

size created by the welding process is critical to the crack propagation and the margin of safety 

achieved in this structure. Once a crack develops, it is desirable to keep the crack contained 

within the weld and not have it progress to the base material adjacent to the weld. This action 

may depend on the nature of the heat-affected zone in the immediate vicinity of the weld. In 

general, the heat affected zone results in anisotropic material properties and residual stresses that 

tend to reduce the toughness of the steel in this area and increase the likelihood of fracture into the 

base metal. For the high strengths steels used in modern shield supports, proper heating and 

cooling of the steel during the weldment process is crucial to preventing crack initiation and 

growth in the heat affected zone. This is why it is very difficult to conduct repairs to damaged 

shields underground, since it is virtually impossible to be able to preheat and properly cool the 

steel when welding at the longwall face. Another approach that can be used to keep crack growth 

contained to the weldments is to create breaks in the weldments. This practice is sometimes used 

in leg socket castings. The break in the weld at the corners of the casting acts like as a crack 

arrester to stop the growth of the crack. This same technique is used in the airplane industry by 

drilling holes in the metal at the end of an observed crack before it reaches a critical crack length. 

Perhaps the best approach to avoid problems associated with weldments is to eliminate them 

when possible. A good example of this pertains to the leg socket design. A typical construction 
for the leg cylinder base socket is shown in figure 15. A casting 3-4 inches thick and 18-24 

inches long with a spherical seat to accommodate the bottom of the hydraulic cylinder is placed 

on top of the bottom cover plate on the base pontoon, and is welded in place along the four sides 

of the top of the casting to the side rib plates and cross plates that connect the two side base plates 

together. This design is highly dependent on the welds to transfer load into the side rib plates and 

maintain the structural integrity of the socket connection. An alternative design that is now being 

employed in some canopy leg sockets to alleviate the weld fatigue problem is to cut rectangular 

holes into the side base rib plates and extend the width of the casting so that it bridges across the 

cylinder opening but is supported by the side rib plates (figure 15). 

In this configuration, leg cylinder loading is transferred directly to the side rib plates of the base 

structure entirely through base metal contact, and is not dependent on the weldments to achieve 

this load transfer. 

 

Side view 

A 

A 

Socket casting

Front View

 A-A 

 Current design 

Side rib plates

 Front View

 A-A 

Socket casting 

 Alternative design

Figure 15 Conventional and alternative leg cylinder casting design 
Corrosion Control - Some steps can be taken to circumvent the problems caused by corrosion. 

The most convenient approach is to try to protect the shield from the environment. For shield 

supports, this mostly applies to painting with an industrial paint that is resistant to the wet mine 

environment. However, for some components such as the pins and clevises where there is 

considerable wear due to the kinematics of the shield during load application, painting or even 

plating with more resistant material is not an option. Impregnated pins with zinc phosphate 

through sheridizing has proven beneficial, but the effectiveness is limited due to the wear in the 

joint. 

Tests have shown that compressive stresses on the surface of materials that are exposed to the 

environment will not necessarily minimize corrosion-fatigue crack initiation, but they can reduce 

the possibility of crack growth and thereby prevent failure from occurring. This could be done by 

induction hardening the joint pins. The concern with this approach would be that the induction 

hardening might make the pins more brittle. 

Material Toughness - The strength of materials approach essentially ignores the toughness of the 

material, which as described is the most significant material property for fracture control. There 

is a tradeoff in the use of high strength steels if they provide superior strength but reduced 

toughness with greater chance for failure. Since the author is not familiar with the German and 

British steels used in the shield construction, no specific recommendations are made, but the issue 

should be investigated with the shield manufacturer when new shields are purchased. 

NEW DESIGN FEATURES AND GOALS FOR THE NEXT 

GENERATION OF SUPPORTS 

1. 100,000 Cycle Life Expectancy – The increases in life expectancy realized during the past 10-

20 years should continue in the near future. It is not inconceivable to have a shield survive 

100,000 loading cycles. In order to reach this goal, some improvements in design may be 

necessary or at least closer attention will need to be paid to fundamental design practices. In 

particular, components will need to be sized such that stress levels are kept at moderate levels 

with respect to the yield strength of the steel. Stress concentrations due to sharp changes in 

geometry must be avoided and weldments must be of high quality. More corrosion resistant 

materials may be used in certain areas to reduce stress corrosion. Make sure that the link bores 

have sufficient bearing areas to avoid high levels of stress and allow for the additional stresses 

caused by corrosion in the design of the link bores. 

2. Hydraulic Diagnostic Capability – All shields will experience leakage in the leg cylinders or 

other hydraulic components that degrade the support capability. An algorithm should be 

specifically written into the control computer that monitors leg pressure histories that will 

provide the longwall coordinator with a record of bad leg cylinders that need to be repaired. 

Another option is to have the computer reset leg cylinders that are leaking below a designated 

setting pressure. While this capability exists on some modern faces, it should become a 

standard part of the operating system, and a report should be written to notify the operator 

when these events are occurring. 
Observations of the relative positions of the cylinder staging can be used to identify cylinder 

problems and the cause of hydraulic leakages. One indication of internal leakages is when 

the bottom stage is consistently fully extended. The bottom stage should be fully extended 

only on operating cycles which establish a new maximum operating height. Hence, on the 

majority of operating cycles, the bottom stage should not be at full extension. Full extension 

of the bottom stage can be determined by analysis of the leg pressure data following the 

setting operation. When the bottom stage is fully extended, there will be a period of nearly 

zero increase in leg pressure until the top stage is sufficiently pressurized to offset the stress 

developed in the cylinder casing by the setting pressure in the bottom stage (figure 16) 

(AISC, 1980). 

0 50 100 150 200 250 300 350 

ADDITIONAL SHIELD LOADING AFTER SETTING, tons 

Figure 16. Difference in hydraulic pressure measurements when bottom is fully or partially extended 

2,500 

3,000 

3,500 

4,000 

4,500 

5,000 

5,500 

6,000 

6,500 

LEG PRESSURE, psi

Full extension of 

bottom stage 

Partial extension of 

bottom stage 

This information can also be used to track the percentage of time that the shield is set with 

reduced setting pressure due to full extension of the bottom stage. When this occurs, load 

development following the setting load should be monitored and analyzed to see if the 

reduced setting load results in higher levels of loading following the setting action. 

Relative changes in staging position can also be used to detect leakages. Some possibilities 

are shown in table 1. The capability to track leg staging does not exist on current shield 

designs, but it is something that could improve the hydraulic diagnostic capability. 

Component Failure Bottom Stage Movement 

Staging check valve ............. Up 

First stage seals ................... Down 

Second stage seals .............. Up 

Pilot operated check valve .. Down 

Yield Valve ............................ Down 

Table 1. Stage movements associated with internal component failures 
3. Smart Load Control – There is evidence to suggest that shield life can be improved by 

minimizing the load development on the shield, particularly since the design stress levels provide 

only a small margin of safety relative to fatigue effects and long term life expectancy. 

The benefits of derating a shield by installation yield valves which operate at lower pressures to 

reduce the total loading on the shield were discussed earlier in the paper. Another option would 

be to have programmable valves that the user can set without the need to physically change them 

out. This would allow the user to have some control over the rated capacity of the support, and 

to derate it if the full capacity is not needed. 

A better approach would be to optimize the use of the capacity that is available. With 1,200 tons 

of support capacity available in modern shields, it is doubtful that all this capacity is needed all 

the time. If the capacity could be used only when it is needed, then adequate ground control 

could be achieved and optimum utilization of the shield would be realized while minimizing the 

duty cycle of the support. To a large extent the shield loading is dictated by the setting pressures, 

which typically represent about 70 pct of the available shield capacity. A challenge for future 

shield designs would be for the control computer to optimize the set pressure by monitoring the 

load development history, and adjusting the set pressure to minimize total shield loading once a 

ground reaction curve for the longwall face is established. The idea of the ground reaction curve 

is to measure the ground response in terms of the support load density. Essentially, the goal is to 

not over support the ground, which may in fact cause more harm that good. The ground 

behavior can be measured in terms of convergence, which will manifest itself in shield loading. 

Generally speaking, the setting pressure is beneficial only if it controls the subsequent ground 

behavior or face convergence. Hence, there is probably some nominal setting pressure that is 

required, as well as some upper limit in which there is no benefit to further increases in setting 

pressure, and a region in between these two extremes where the benefit is measured in terms of 

the amount of convergence that is occurring. Hence, by monitoring the load development history 

of the support for a few cycles, and then lowering the set pressure until subsequent loading 

increases beyond the baseline, an optimum set pressure could be derived. By incorporating an 

algorithm into the control module, the set pressure could be adjusted as the conditions change to 

continually optimize the support capacity utilization. 

Another possibility is to have reduced setting pressures on sets of shields. For example, every 

other shield could be set at a reduced setting pressure. The concern here would be if that causes 

increased loading on the neighboring shields, then this practice would not that helpful since it 

would be a tradeoff of reducing the loading in some shields at the expense of others. However, 

this information could be obtained if the loading histories were monitored. 

4. Lubricated Link Joints – Pin joints are necessary to accommodate the kinematics of a shield, 

but these joints are by far the leading cause of structural rebuild. Design recommendations have 

previously been made regarding improving the structural margin of safety by adding more 

contact area to offset the problems associated with distortion of the clevis due to wear, eccentric 

loading of the joints due to lateral canopy movements, and corrosion effects which reduce the 
 

effective contact area and cause stress concentrations. Efforts should continue to enhance the 

use of wear resistant materials such as impregnating the pins with zinc phosphate, but 

consideration should also be given in future shield designs to a lubricated joint to reduce the 

wear and damage that occurs in these critical areas of the shield structure. 

5. Composite Material Design – Even with the use of high strength steels, modern shields weigh 

as much as 30 tons. Weight is a major barrier in increasing the shield widths beyond 2.0 meters 

with the current steel constructions. Composite materials of equivalent strength weigh only 1/4 

that of steel fabrications. While there are several engineering and economic issues that need to 

be explored before the feasibility of using high technology composite materials for shield 

construction can be determined, this could lead to a new generation of shield supports in the 

future. Axially loaded components such as lemniscate links might be one to examine first, and 

perhaps only sections of the other components could be fabricated from composites, but it 

definitely an area worth investigating. 

6. Forecasting Periodic Weighting and Heavy Roof Loading– Most modern shield systems are 

capable of capturing shield loading histories to the point where we are overloaded with data. 

This data, if properly managed can be useful in evaluating leg cylinder leakage and optimizing 

setting pressures as suggested in above. Another area is of value is the prediction of periodic 

weighting intervals. There has been some research in this area already. For example, NSA 

Engineering Inc developed a shield monitoring program that couples with their GeoGuardTM 

software [Sanford et al. 1999]. The system was recently tested in an Australian mine where 

severe face weighting was observed resulting in hazardous face conditions and weeks of lost 

production. 

7. Advanced Component Load Measurement – Currently, only the leg cylinder pressure is 

monitored. No loading information is obtained on any other components. It would not be 

difficult in the design process to include strain gages on selected components which could be 

monitored by the computerized data acquisition system. This need not be done on every 

production shield, in fact 2 or 3 shields could serve as instrumented shields and provide valuable 

information on the actual load conditions observed underground. This information could then be 

used to refine performance testing as well as help to identify load conditions that actually 

contribute or cause structural failures. Even measurement of the link loading alone would 

provide valuable information that is currently not available. 

8. Constant Set Leg Cylinder Design – Current U.S. shields employ a two-stage cylinder design 

where the first stage extends and retracts first. This operation typically causes the first stage to 

be near full extension most of the time. In addition, seal leakage which occurs on almost all 

aging shields causes the bottom stage to extend outward and further promotes full extension of 

the bottom stage. Since the setting force and the subsequent capacity of the shield to resist roof 

loading is reduced in proportion to the area differential between the top and bottom stage when 

the bottom stage is fully extended, this is not a good design to ensure that the full capacity of the 

shield is utilized. In other words, why buy a 1,000 ton shield when its performance is degraded 

to a 500 ton shield a significant portion of the time. There are alternative designs available, most 
 

 

notably the concept employed by Fazos Inc., which do not use the conventional two-stage design 

utilized in most modern shields. The Fazos design utilizes a central core to eliminate the use of 

independently acting stages (figure 17). While this design works, it has not be adopted by the 

major shield manufacturers for whatever reason. An alternative approach would be to configure 

the standard two-stage design with a system that would cause the top stage to extend first. This 

could be done by using a nominal hydraulic pressure in the retract annulus of the bottom stage 

during the extend operation (figure 18). For current shields, a mechanical device (strap or chain 

holster) could be built that would attach to cylinder casing and the base that would prevent the 

bottom stage from extending. This is not something that would be used on every mining cycle, 

but would be done periodically as part of a preventive maintenance program to restore full 

setting capability to the leg cylinders which are routinely at full bottom stage extension. 

BOTTOM 

TOP STAGE 

STAGE 

Figure 17 Constant set leg cylinder design 

developed by Fazos, Inc. 

 

 

TOP STAGE 

BOTTOM STAGE 

INTERNAL 

CHECK 

VALVE 

Fig. 18 Maintaining nominal pressure on bottom 

stage to force top stage out first 

REFERENCES 

American Institute of Steel Construction, Inc. (1980). Manual of Steel Construction. Eighth 

Edition, pp. 5-89. 

Barczak, T. M. (2000). Examining Longwall Shield Failures from an Engineering Design and 

Operational Perspective. Published in NIOSH Information Circular 9453, pp. 223-244. 

Barczak, T. M. (1998). Performance and Safety Considerations of Hydraulic Support Systems. 

Proceeding of the 17

th

 Conference on Ground Control in Mining, Morgantown, WV, August 4-6, 

pp. 176-186. 

Barczak, T. M., Babich, D. R. and Chekan, G. J. (1998). Shield Design, Construction, and 

Operation. Longwall USA, June 9-11, pp 5-48. 

Coal Age (1999). Longwall Census, February, pp. 22-27. 

Sanford, J., Mahoney, S., Conover, D. P. and DeMarco, M. J. (1999). Shield Monitoring to 

Forecast Severe Face Weightings at the South Bulga Colliery, NSW, Australia. Proceeding of 

the 18th Conference on Ground Control in Mining, Morgantown, WV, August 4-6, pp. 164-176.