Comparison of methods: dynamic versus
 

hydrostatic testing of mine ventilation seals
 

M.J. SAPKO, S.P. HARTEIS AND E.S. WEISS 

M.J. Sapko, member SME, S.P. Harteis and E.S. Weiss are 

senior research physical scientist (retired), research engineer 

and manager (mining engineer) of the Lake Lynn Laboratory 

Section, respectively, with the U.S. Department of Health and 

Human Services, Centers for Disease Control and Prevention, 

National Institute for Occupational Safety and Health, Pittsburgh 

Research Laboratory, Pittsburgh, 

Abstract 

From 2001 to 2007, the Pittsburgh Research Laboratory 

(PRL) of the National Institute for Occupational Safety and 

Health (NIOSH) conducted studies to develop alternative 

methodologies to full-scale explosion testing for determining the ultimate strength of mine seals. As a result, the 

PRL developed and proposes an alternative seal-strength 

evaluation method based on a hydrostatic pressure-loading 

concept. The researchers suggest pressure loading a seal 

using water to twice the expected dynamic design load.The 

hydrostatic chamber test offers a means of validating seal 

designs, establishing appropriate resistance functions and 

determining the ultimate strength of seals through testing 

to failure.This article contrasts the full-scale explosion and 

hydrostatic testing of mine seals using a simple dynamic 

system model and principles. 

Introduction 

Tragic mining disasters in West 

Virginia (Sago Mine) and Kentucky 

(Darby Mine) in 2006 greatly heightened interest in the development of 

quality seals to protect miners from 

blast effects and toxic gases produced 

by uncontained gob explosions. Following the enactment of the 2006 

Miner Act and the U.S. Mine Safety 

and Health Administration (MSHA) 

issuance of the Emergency Temporary Standard (ETS) on the Sealing 

of Abandoned Areas in 2007, MSHA is conducting detailed technical evaluations of all proposed seal designs 

for each underground sealing location. These technical 

evaluations are based on sound structural engineering 

design approaches followed by certification of as-built 

construction. MSHA will also accept the results from 

full-scale testing of mine seals. If full-scale performance 

testing is required to develop seal stress-strain response 

data, the authors propose a hydrostatic test method as 

an alternative to full-scale explosion testing.This method 

uses water to load the seal to pressures at least twice the 

expected dynamic design load to determine stress-strain 

data. This article contrasts the full-scale explosion and 

hydrostatic testing of mine seals using a simple dynamic 

system model and principles. 

There are many three-dimensional, finite-element 

codes for designing various load-bearing structures that 

can simulate not only the seal itself, but also the interaction of the seal with the surrounding support strata. However, when studying a single mode of seal response, the 

basic analytical model used in most blast-design applications is the single-degree-of-freedom (SDOF) system. In 

the current study, the U.S. Army’s Wall Analysis Code 

(WAC) (Slawson, 1995) was used to 

identify test conditions where hydrostatic pressure loading can serve 

as an alternative yet equivalent test 

method to full-scale methane-air explosions for evaluating the strength 

performance of mine ventilation 

seals. Such hydrostatic tests would 

build confidence in design methodology through comparisons between pretest predictions and actual 

measurements, as well as provide a 

timely, cost-effective means of seal 

testing. 

Single-degree-of-freedom system (SDOF) 

All structures possess more than one degree of freedom regardless of how simple their construction. However, many structures can be adequately represented as a 

SDOF system for purposes of analysis.The accuracy of an 

SDOF approximation depends on how well the deformed 

shape of the structure and its resistance can be represented with respect to time. The procedure for obtaining the 

equivalent SDOF approximation for a structural component is based on its deformed shape under the applied 

loading and the strain energy equivalence between the 

actual structure and the SDOF approximation. 

Equivalent mass, stiffness and loading are obtained 

through the use of transformation factors. Chapter 3 of 

the U.S. Department of Army, the Navy and the Air Force 

design manual TM 5-1300 (1990) contains tabulated 

transformation factors for typical structural elements, 

including slabs. The derivations of the equations for the 

transformation factors are also given in this reference 

and used in the WAC. 

The WAC is an SDOF model developed for the U.S. 

Army Engineer Waterways Experiment Station (WES), 

Structures Laboratory (Slawson, 1995). WAC was developed to provide a tool for the easy calculation of the 

response of typical walls subjected to blast loads. The 

WAC can calculate the resistance function (R(y), pressure-deflection) of a wall given its construction details, 

including dimensions, material properties and support 

conditions or accepts user defined resistance functions 

based on experimental data. The WAC transforms the 

wall model to an equivalent SDOF model, calculates the 

actual and SDOF equivalent loads and solves the equation of motion to determine the response time history of 

a central point on the wall. 

The equation of motion for an SDOF system is 

M·yʺ(t) + Cd

·yʹ(t) + R(y(t)) = F(t) (1) 

where 

M is the equivalent or “lumped” mass of the system; 
 

Cd is the damping coefficient taken as 5 percent of the 

critical value, i.e. very lightly damped; 

y(t) is the placement of the mass as a function of time 

t; 

yʹ(t) is the velocity of the mass or first derivative displacement; 

yʺ(t) is the velocity of the mass or second derivative 

displacement; 

R is the structural resistance as a function of displacement y; and 

F(t) is the structural load as a function of time, i.e. 

expected blast pressure history. 

FIGURE 1 

One-way arching failure mechanism in WAC. 

Mine seals and one-way arching failure 

The U.S. Bureau of Mines (Rice et al., 1930) conduct

ed a series of tests and found that restraining the edges o

a seal caused a dramatic increase in the seal strength t

a level much higher than predicted by plate theory. Full

scale explosion experiments also showed concrete wall

that were recessed into the roof, ribs and floor and had 

thickness to width ratio of at least 0.1 resisted much high

er pressures than the theoretical design pressure. Thes

results showed that recessing the ends of the concret

wall into the surrounding strata allows the wall to ac

as a “flat arch.” This arching behavior transmits a latera

thrust to the strata, which then act as a buttress to preven

seal movement. According to Anderson (1984), seal fail

ure data from tests conducted in the NIOSH chamber

correlated well with the following simplified formula fo

one-way-arching action in transverse laterally loade

masonry wall panels 

Pmax = Fc*(T/H)2 

where 

Pmax is the predicted ultimate failure pressure; 

Fc is the effective compressive strength of the bloc

and mortar; and 

­

f 

o 

-

s 

a 

­

e 

e 

t 

l 

t 

­

s 

r 

d 

(2)

k 

T and H are the thickness and height of the seal, respectively. 

In the current study, the SDOF model WAC was used 

to contrast the differences in seal deflection as a function 

of seal loading rate. The analysis option chosen in the 

WAC was an “unreinforced wall with one-way arching.” 

In this option, the supports are rigid at the roof and fl oor, 

while the sides of the walls are unrestrained.This arching 

model for wall behavior applies best when the wall height 

to wall thickness ratio is greater than 4 and less than 15. 

Note that in the actual mine environment, this mode of 

failure would require stiff abutments at roof and fl oor to 

mobilize the arching effect. Otherwise, the resistance of 

the seal to the blast would be reduced signifi cantly. 

In the arching failure mechanism, the wall is assumed 

to crack horizontally at mid-height and at the roof and 

floor upon application of the blast load. As shown in Fig. 

1, the two wall segments remain rigid, rotate through an 

angle θ and develop arching forces to resist the blast loading. The wall will begin to crush at the points indicated 

by the vertical solid arrows in Fig. 1, and the magnitude 

of the resisting forces will depend on the compressive 

strength of the wall material and the contact area at the 

crush points. 

FIGURE 2 

WAC-calculated resistance function for the example 

seal assuming one-way roof-to-floor arching. 

Figure 2 shows the WAC-calculated resistance function (R(y)) for a 1.8-m- (6-ft-) high, 6.1-m-(20-ft-) wide, 

406-mm- (16-in.-) thick unreinforced masonry solid block 

wall (herein referred to as the “example seal”). This example seal was used for comparing the dynamic and static 

responses calculated with the WAC for assumed and actual experimental mine pressure histories. As shown in 

Fig. 2, the displacement is linear to about 4.3 mm (0.17 

in.), ~1.48 MPa (~214 psig), i.e., in the elastic range. In the 

elastic range, the deflection of the seal will return to its 

starting defl ection (zero) when the pressure is removed. 

Above 4.3 mm (0.17 in.), the response of the seal enters 

the elastic-plastic range, and above approximately 8.6 

mm (0.34 in.), 1.63 MPa (237 psig), the seal response is 

in the fully plastic range, where the seal will either suffer 
 

 

permanent displacement and/or failure. All calculated 

dynamic responses were in the elastic or slightly elasticplastic range implying that the example seal sustained 

little or no permanent damage. 

FIGURE 3 

Simulated pressure loading histories F(t). 

Response of the example seal using simulated

pressure histories 

The WAC is used in the following analysis of the example seal to illustrate the ranges for dynamic and static 

responses. The seal loading F(t) is assumed in this case 

to be a uniform pressure history. Shown in Fig. 3 are fi ve 

simulated pressure histories used in the analyses. Each 

pressure history peaks at 830 kPa (120 psig) with corresponding rise times of 0, 2, 4, 8 and 16 ms. 

WAC calculates the undamped natural period of the 

wall (example seal) by using the stiffness Ke and mass M

as shown in 

Tn = 2π (M/Keg)1⁄2 (3) 

where 

M is the equivalent or “lumped” mass of the system 

(lb), 

Ke

 is the elastic stiffness for slabs (psi/in.) and 

g is the acceleration of gravity (in./sec2). 

The WAC-calculated natural period is 8 ms for the 

example seal (solid-concrete-block with a compressive 

strength of 17.2 MPa (2,500 psi) and a density of 2,080 kg/ 

m3 (130 pcf)).The WAC-calculated, one-way arch natural 

periods for a 6.1-m- (20-ft-) wide, 406-mm- (16-in.-) thick 

example seal as a function of height and compressive 

strength, [10.3 to 17.2 MPa (1,500 to 2,500 psi)] varied 

from 6 to 36 ms for seal heights ranging from 1.5 to 3.7 

m (5 to 12 ft). 

FIGURE 4 

WAC-calculated centerline deflection for the example seal as a function of 

the simulated pressure loadings shown in Fig. 3. NP is the natural period of 

the example seal (8 ms). 

Figure 4 shows the WAC-calculated centerline displacement for the example seal when subjected to the fi ve 

pressure loading histories shown in Fig. 3.The largest dynamic deflection of 4.8 mm (0.19 in.) occurred when the 

example seal was instantaneously (0 

ms rise time) loaded to 830 kPa (120 

psig). At rise times of 2 and 4 ms, the 

peak displacement of the example seal 

decreased as the pressure rise times 

increased. When the pressure rise 

time from a dynamic load was greater 

than or equal to the natural period of 

the seal (8 ms for the example seal), a 

maximum deflection of 2.4 mm (0.095 

in.) occurred. This is the same defl ection that would be obtained from a 

830-kPa (120-psig) static load applied 

slowly. When a seal sustains a blast, 

i.e., a rapidly applied load, the protection that the seal can provide is based 

on its dynamic resistance.The applied 

dynamic load for a particular situation can be regarded as having some 

equivalent static load. An equivalent 

static load is defined as a load that is 

distributed in the same manner as the 

dynamic load and produces the same 

shear and bending moment at any 

given cross-section of the seal as the 

dynamic load. 

To best compare the dynamic and static response of 

an elastic system, the Department of Army, the Navy and 

the Air Force (1990) Design Manual TM5-1300 refers to 

the concept of dynamic load factor (DLF) sometimes 

referred to as the Dynamic Magnification Factor (DMF). 

DLF used here is defined as the ratio of the maximum dynamic deflection produced by rapid loading of pressure P

to the deflection that would have resulted from the static 

or slow application of pressure P. 

DLF = Xm/Xs (4) 

where 

Xm is the maximum dynamic deflection produced when 

the peak pressure is applied rapidly and 

Xs is the static deflection or the displacement produced 

in the system when the peak pressure is applied 

slowly. 

Because deflections, spring forces and stresses in an 

elastic system are all proportional, the DLF may be ap­
 

plied to any of these to determine the ratio of dynamic 

to static effects. 

FIGURE 5 

Seal test area in the LLEM. 

In many structural problems, only the maximum value 

of the DLF is of interest. For the most prevalent load case, 

namely, the rectangular and step load with instantaneous 

or 0 ms pressure rise time, the maximum value of the 

DLF is 2.This indicates that all maximum displacements, 

forces and stresses due to the dynamic load are twice the 

value that would be obtained from static analysis for the 

same maximum pressure. It should be noted that under 

particular circumstances, the DLF could exceed 2, for 

example, if the seal is subjected to numerous shocks at or 

very near the natural period of the seal. The WAC calculations help demonstrate that the response of any given 

candidate seal design for explosion isolation is dependent 

on the peak explosion pressure exerted on the seal and, 

equally important, the rise time to peak pressure relative 

to the natural period of that seal. 

Response of example seal using actual LLEM

pressure histories 

LLEM drift explosions. Recently, NIOSH conducted 

larger-scale experiments within the LLEM to simulate, in 

part, the dynamic response of seals relevant to the direction of blast loading and the influence of refl ected waves 

on seal response. The measured pressure histories from 

these experiments were used in the WAC to calculate the 

dynamic displacement of the example seal. 

Figure 5 is a close-up plan view of the seal test area in 

the multiple-entry area of the LLEM (Weiss et al., 2002, 

2008). Methane-air explosions were conducted in C-drift, 

which was blocked with a test seal constructed perpendicular to the axis at ~98 m (~320 ft) from the closed 

end ignition zone, as shown in Fig. 5. The gas-ignition 

zone was confined by a plastic diaphragm across C drift, 

14 m (47 ft) from the face. About 18.7 m3 (661 cu ft) of 

natural gas was injected into the ignition zone to produce 

a mixture of ~10 percent CH4 in air.The 187 m3 (6,610 cu 

ft) flammable methane-air mixture filled only ~15 percent of the total sealed volume. The fl ammable gas was 

ignited at the face, and the pressure pulse propagated 

out C-drift past the seals in cross-cuts X-1, X-2 and X-3, 

as shown in Fig. 5. In this system, a refl ected pressure 

wave results when the head-on pressure wave impacts 

the seal across C-drift.This reflected pressure wave then 

propagates back toward the ignition source, impacts the 

closed end, reflects and propagates back outby. Pressure 

histories from an LLEM explosion test sampled at 5,000 

Hz and recorded at pressure transducers indicated by 

“�” in Fig. 5 were then used in WAC to calculate the 

dynamic displacement of the example seal had that seal 

been constructed in those locations. 

FIGURE 6 

Pressure history from the confined LLEM test and WAC-calculated center 	

displacement from the example seal had it been located 98 m (320 ft) from the 	

face of C-drift (see Fig. 5). 

Shown in Fig. 6 is the explosion history recorded in 

front of an actual test seal constructed across C-drift 98 m (320 ft) from 

the closed end. This history shows 

the pressure of the refl ected wave to 

be approximately 360 kPa (52 psig), 

followed by several oscillations with 

a peak pressure of about 390 kPa 

(57 psig). Also shown in Fig. 6 is the 

WAC-calculated dynamic displacement of the example seal if subjected 

to this same pressure loading. The 

WAC-calculated DLF for the example seal was ~1.54 times the expected 

static deflection of the 390 kPa (57 

psig) load. Therefore, a static test of 

about 610 kPa (88 psig) (~1.54 x 57 

psig) is needed to produce the same 

response as the explosion test of 390 

kPa (57 psig). 

Using the pressure history of the 

LLEM test recorded at a test seal in 

Crosscut 3, the WAC-calculated dynamic displacements for the example 

seal are shown in Fig. 7. The initial 

outgoing peak pressure wave impacting the Crosscut 3 test seal was approximately 240 kPa (35 psig), while 
 

the reflected pressure wave returning from the C-drift seal produced a 

second peak of approximately 250 

kPa (36 psig). The calculated DLF 

for the example seal for this same 

pressure loading was 1.05.This DLF 

would suggest that a static test of 

only ~260 kPa (~38 psig) (~1.05 x 

~36 psig) is needed to produce the 

same response as the 250 kPa (36 

psig) dynamic test. 

FIGURE 7 

Pressure history from the confined LLEM test and the WAC-calculated center 

displacement for the example seal had it been located in Crosscut 3.

Shown in Fig. 8 is the actual 

recorded pressure history from a 

sensor located near the closed end 

ignition zone during the confined 

LLEM test, as well as the WACcalculated response of the example 

seal had it been placed across the 

closed end of C-drift during this test. 

The initial pressure built slowly as 

the fireball expanded and the fl ame 

and pressure propagated away from 

the ignition.The calculated DLF for 

this first peak pressure loading of 

150 kPa (22 psig) was approximately 1.05. When the reflected wave 

returned from the test seal located 

across C-drift 98 m (320 ft) away and 

impacted the simulated example seal located at the face, 

WAC calculated a much higher DLF of 1.79. Although 

the calculated displacement frequency before and after 

the pressure wave reflection is the same, the amplitude 

increases signifi cantly. This increase in amplitude is due 

in large part to the more rapid rate of loading produced 

by the reflected wave, similar to the seal response shown 

in Fig. 4. 

FIGURE 8 

Pressure history from the confined LLEM test and the WAC-calculated displacement of the example seal had it been located at the closed end ignition 

zone of C-drift. 

As these studies indicate, several 

factors influence the dynamic load and, 

subsequently, the DLF. These factors 

include the location of the seal in the 

entry and the occurrence of multiple 

reflecting waves. In the current studies, these factors produced DLFs that 

ranged from 1.05 to 1.79, with peak 

pressure loadings ranging from 150 to 

390 kPa (22 to 57 psig). Because there 

are many unknown factors that may 

occur in an actual underground coal 

mine explosion, it would be prudent to 

design the seal based on a static load of 

twice the dynamic performance pressure (DLF of 2), which is appropriate 

for most designs responding within the 

elastic range. If design engineers know 

of specific circumstances in which pressure loading might match the natural 

period of the seal, then a higher DLF 

may be required for design and testing. When designing a seal using a dynamic loading model, such as WAC or 

three-dimensional dynamic numerical 

model, then one should use a pressure 

loading history with an instantaneous 

pressure rise to the peak design pressure lasting for ≥2 times the natural period of the seal. For 

new seal designs, MSHA (2007) uses an instantaneous 

pressure rise lasting for 1.5 seconds. 

Chamber tests using methane explosions. A largescale underground chamber was constructed within 

LLEM to conduct hydrostatic and confi ned explosion 

pressure-loading evaluations of seals (Sapko et al., 2005). 
 

 

 

The 6.1-m- (20-ft-) wide, 2.7-m- (9-ft-) high, 3.0-m- (10-ft-) 

deep chamber can accommodate a seal design with a 

cross-sectional area up to 16.7 m2 (180 sq ft). 

To determine the response and ultimate failure pressure of various seal designs, methane was injected into 

the ~42.5 m3 (~1,500 cu ft) closed volume behind the seal, 

mixed and then ignited in the geometric center. Figure 9 

shows three of these explosion pressure-loading histories. 

The combustion of 6 percent methane-air mixture, the 

slowest burning rate tested, peaked at about 8,000 ms at 

390 kPa (56.5 psig), while the most rapid burning rate, 13 

percent methane with added oxygen, peaked at 340 ms at 

690 kPa (100 psig). Note the maximum theoretical pressure from the constant volume (no leakage or heat loss) 

of 9 percent methane-air in a sealed chamber would be 

about 770 kPa (112 psig). However, due to gas leakages 

through and around the perimeter of the seal and heat 

loss, the empirical peak pressure was about 630 kPa (91 

psig) for the 9 percent mixture. As the combustion rate 

increased for the oxygen enriched mixture, there was less 

time for significant pressure loss due to gas leakage from 

the chamber. 

When the actual pressure history from the most rapid 

burning rate test was used to then calculate the defl ection 

response of the example seal, the time to peak pressure 

of 690 kPa (100 psig) was 340 ms. This 340 ms rise time 

is very slow compared to the 8 ms natural period of the 

example seal and therefore is essentially equivalent to a 

static load despite being an explosion test.The calculated 

dynamic deflection of 2 mm (0.078 in.) is the same defl ection one would obtain with an equivalent static load of 

690 kPa (100 psig) (DLF = 1). 

FIGURE 9	 

Chamber explosions pressure histories recorded during testing of an actual 

standard-type, 2.6-m- (8.6-ft-) high, 5.5-m- (18-ft-) wide, 406-mm- (16-in.-) thick 

solid concrete block seal. 

Chamber tests using water. After conducting the 

chamber tests using methane explosions, new seals were 

constructed and water pressure was then used to hydrostatically load seals of the same design. To achieve this 

hydrostatic loading, the chamber was filled with water 

while air was displaced through a top vent pipe. When 

water was observed venting through the pipe, the air vent 

valve was closed, allowing the water pressure behind the 

seal to increase. The water pressure 

history, as recorded from a transducer 

located on the test seal at 1.5 m (5 ft)

above the floor peaks at 220 kPa (32 

psig) in 68 min. 

Figure 10 shows the corresponding centerline deflection for three 

confi ned explosions tests for a 2.5-m- 

(8.3-ft-) high, 406-mm- (16-in.-) thick 

solid concrete-block test seal. The 

deflections increased linearly with 

increasing pressure at a nearly constant rate for all three explosions.Also 

shown is the response of another seal 

of the same design subjected to water loading. The centerline response 

of the seal was nearly the same as 

the explosion tests despite different 

rates of pressure loadings. All rates of 

loading generated from explosion and 

water loading tests conducted in the 

chamber were much longer than the 

natural period of the seal, and their 

responses were consistent with a static 

load (DLF of 1). Figure 10 also shows the WAC calculated 

resistance-displacements for the test seal when subjected to the pressure history obtained from the 13 percent 

methane-air plus added oxygen. The WAC calculated 

resistance-displacements are in reasonable agreement 

with water and explosion loading experiments indicating that one way roof-to-floor arching is an appropriate 

mode of resistance for these tests. Taken together, these 

results indicate that the hydrostatic approach coupled 

with an SDOF model shows promise as a tool for mine 

seal design and testing. Studies are underway to further 

refine and quantify the response of various seal designs 

under different loading conditions. 

FIGURE 10 

Comparison of experimental deflection responses of an 

actual solid-concrete-block seal design in the chamber 

exposed to three methane air explosions and water 

loading test. 

Summary 

These studies demonstrate that hydrostatic testing 

of mine seals is a reasonable alternative or adjunct to 
 

 

full-scale explosion testing. Hydrostatic chamber testing 

allows for multiple water loadings for stress-strain measurements and for determining the ultimate strength of 

the seal through testing to failure. Furthermore, empirical resistance data derived from the hydrostatic testing 

of seals can then be used to improve or validate the 

WAC and other design codes. Because it is not possible 

a priori to predict the dynamic load at each seal location surrounding a coal mine gob, the authors suggest 

statically designing or testing seals to at least twice the 

expected dynamic performance pressure, i.e., a DLF 

of 2. 

The current studies show reasonable agreement in 

seal performance between explosion and hydrostatic 

tests and the WAC. The arching mechanism of resistance described in this report requires the seal to have 

stiff abutments at roof and floor to mobilize the arching 

effect.The study details results obtained with limestone 

strata that provided near ideal rigid supports. Structural 

engineering designs specifying mine strata interface 

and supports must be addressed to ensure the in situ 

performance of the seal design. 

Sound structural engineering design followed by 

certifi cation of as-built seal construction, coupled with 

hydrostatic performance testing as needed, will help 

facilitate the development of reliable seals and enhance 

the level of protection for underground personnel. 

Disclaimer 

The findings and conclusions in this report are those 

of the authors and do not necessarily represent the 

views of the National Institute for Occupational Safety 

and Health. 

References 

Anderson, C., 1984, “Arching action in transverse laterally loaded 

masonry walls,” The Structural Engineer, Vol. 62B, No. 1, March, 22 pp. 

CFR Code of Federal Regulations, U.S. Government Printing Offi ce, Office of the Federal Register, Washington DC. 

Congress of the United States, 2006, “New Emergency Response 

Act of 2006” (The MINER Act). 

Department of Army, the Navy and the Air Force, 1990,“Structures 

to resist the effects of accidental explosions,” TM 5-1300, Chapter 3, 

Principles of Dynamic Analysis, Washington, D.C. 

MSHA, 2007, “120 psi reinforced concrete seal,” MSHA seal approval number 120–75.336.1.07.01.0, http://www.msha.gov/Seals/Approved120PSIseal.pdf, accessed October 2007. 

Rice, G.S., Greenwald, H.P., and Howarth, H.C., 1930, “Tests of 

the Strength of Concrete Stoppings Designed to Resist the Pressure of 

Explosions in Coal Mines,” USBM RI 3036, September, 11 pp. 

Sapko, M.J.,Weiss, E.S., and Harteis, S.P., 2005,“Methods for evaluating explosion resistant structures,” 8th International Mine Ventilation 

Congress, Brisbane, Australia. July 6-8, 2005. 

Slawson, T.R., 1995, “Wall Response to Airblast Loads: The Wall 

Analysis Code (WAC),” Structures Laboratory, U.S. Army Engineers 

Waterways Experiment Station, Vicksburg, MS. ATTN CEWES-SS. 

Weiss, E.S., and Harteis, S.P., 2008, “Strengthening existing 20-psi 

mine ventilation seals with carbon fi ber-reinforced polymer reinforcement,” U.S. Department of Health and Human Services, Public Health 

Service, Centers for Disease Control and Prevention, National Institute 

for Occupational Safety and Health, Pittsburgh, PA, DHHS (NIOSH) 

Publication No. 2008-106, RI 9673. 

Weiss, E.S., Cashdollar, K.L., and Sapko, M.J., 2002, “Evaluation 

of Explosion-Resistant Seals, Stoppings, and Overcast for Ventilation 

Control in Underground Coal Mining,” U.S. Department of Health and 

Human Services, National Institute for Occupational Safety and Health, 

Pittsburgh, PA, RI 9659.