II^^^V" United States Department of Commerce

I ^1 1 National Institute of Standards and Tectinology

NAT L INST OF STAND & TECH R I.C NIST

PUBLICATIONS

A111D3 mVMlS

NIST BUILDING SCIENCE SERIES 169

Strength and Creep-Rupture Properties of

Adhesive-Bonded EPDM Joints

Stressed in Peel

Jonathan W. Martin

Edward Embree

Paul E. Stutzman

James A. Lechner
NATIONAL INSTITUTE OF STANDARDS &

TECHNOLOGY

Research Information Center

Gakhersburg, MD 20899

DATE DUE

Demco. Inc. 38-293
NIST BUILDING SCIENCE SERIES 169

Strength and Creep-Rupture Properties of

Adhesive-Bonded EPDM Joints

Stressed in Peel

Jonathan W. Martin

Edward Embree

Paul E. Stutzman

James A. Lechner

National Engineering Laboratory

National Institute of Standards and Technology

Gaithersburg, MD 20899

Issued May 1990

U.S. DEPARTMENT OF COMMERCE, Robert A. Mosbacher, Secretary

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY

John W. Lyons, Director
National Institute of Standards and Technology Building Science Series 169

Natl. Inst. Stand. Technol. Bldg. Sci. Ser. 169, 59 pages (May 1990)

CODEN: NBSSES

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1990

For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, DC 20402-9325
ABSTRACT

The most frequently reported defect in ethylene-propylenediene terpolymer (EPDM) single-ply roofing systems is in fieldformed joints. The causes of these joint defects are largely

unknown, but they tend to occur a short time after a roof is

installed.

The objectives of this research were 1) to determine the

importance of the following material and fabrication variables in

causing joint failures: adhesive thickness, cure time,

mechanical load, adhesive type and surface cleanness, 2) to rank

these variables as they affect the creep-rupture life and

strength of butyl-adhered EPDM joints, and 3) to determine the

maximum peel stress which EPDM joints can sustain over their

design life.

From our results, cure time and level of cleanness of the

EPDM membrane have the greatest effect on joint strength; while

the thickness of the adhesive and the magnitude of the mechanical

load have the greatest effect on a joint's creep-rupture time-tofailure. Thus, the rankings of variables in creep-rupture and

short-time strength experiments were different and this

difference should have important implications in standards

writing and in establishing performance requirements for EPDM

roofing joints.

Creep-rupture results were used in determining the maximum

design stress that butyl and neoprene-adhered EPDM joints can

sustain when loaded in peel. The maximum sustainable stress is

only a small fraction (less than 5%) of the short-time y .eld

strength of a joint. Thus, efforts should be made to insure that

the field-formed joints are only subjected to small peel loads.

KEYWORDS: butyl adhesive; creep-rupture life; EPDM; neoprene;

peel performance criteria; seam; single-ply roofing;

strain rate; stress-ratio; time-to-failure

CONTENTS

Abstract iii

1. Introduction 1

2 . Background 2

2.1 Stress Distribution Along the Peel Line 3

2.2 Peel Force 3

2 . 3 Rate of Peel 4

3. Peel Force of Butyl-Adhered EPDM Joints 6

3 . 1 Experimental 6

3.1.1 Materials 6

3.1.2 Specimen Contamination 7

3.1.3 Contamination Level Quantification 8

3.1.4 Specimen Preparation and Test Procedure . . 8

3.2 Model of Strength Versus Cure Time 10

3.3 Results 11

3.3.1 Failure Modes 12

3.3.2 Effect of Adhesive Thickness on Peel. ... 12

Strength

3.3.3 Strength Versus Cure Time at Different. . . 13

Contamination Levels

3.4 Discussion 16

3.5 Peel Strength Conclusions 17

4. Creep-Rupture Times-to-Failure of Butyl-Adhered EPDM . . 17

Joints

4 . 1 Experimental 18

4.1.1 Materials 18

4.1.2 Experimental Procedure 18

4.2 Results 19

4.2.1 Peel Strength at Long Cure Times 19

4.2.2 Failure Mode 2 0

4.2.3 Creep-Rupture Times-to-Failure 22

4.3 Discussion 25

4.4 Creep-Rupture Time-to-Failure Conclusions 25

5. Comparison of Times-to-Failure for Butyl- and 26

Neoprene-Adhered EPDM Joints as a Function of Load Ratio

6. Summary 28

7. Acknowledgments 2 9

8. References 30

9. List of Symbols 34

V

1. INTRODUCTION

EPDM (ethylene-propylene-diene terpolymer) single-ply

roofing membranes are widely used in the United States for lowsloped roof applications. The popularity of this system stems in

part from the chemical inertness of EPDM, which provides the

membrane with its excellent weathering properties. This

inertness also makes it difficult to achieve the strong adhesive

bond between two EPDM sheets which is needed to assure the

quality of a field-formed seam. Evidence for this difficulty is

available from Project Pinpoint's data (Cullen, 1984, 1988,

1989) , from case studies conducted by NIST staff (Rossiter and

Seller 1988, 1989), and from general survey papers on the

performance of single-ply roofing systems (Dupuis, 1983; Dregger,

1988 ; Russo 1988)

The adhesives most commonly used in fabricating field-formed

EPDM seams are neoprene and butyl adhesives. Prior to 198 6, the

neoprene adhesives were dominant, but since that time, the

neoprene adhesives have been largely supplanted by the butyl

adhesives. This switch resulted from improvements in the

formulation of butyl adhesives and from problems with the

durability of neoprene adhesives. The change in adhesives has

not, however, produced any noticeable improvements in Project

Pinpoint's seam defect statistics, which have remained relatively

constant, at approximately 2 5% of the defects reported for

single-ply roofing membranes.

Since no significant change has occurred in the defect

statistics for seams of single-ply membranes, it seems reasonable

to conclude that the causes of these defects have not, as yet,

been identified. The source of these defects, however, can be

garnered from Project Pinpoint's 1984 survey. In that survey

(Cullen, 1984) , it was reported that 62% of all reported seam

defects occur within the first year, 74% within the first 2

years, and 86% within the first 3 years after installation.

Thus, we can conclude that most seam defects occur within a short

time after a roof is installed, implicating weaknesses in the

materials, the design, or the fabrication of a seam, as opposed

to the chemical or physical aging of one or more of the seam

components. The presumption that seam defects are not caused by

the chemical degradation of the seam components is further

supported by the chemical inertness of butyl adhesives (Stucker,

1977) and EPDM rubber, and by the observation that switching from

neoprene to butyl adhesives did not affect Project Pinpoint's

seam defect statistics.

The fabrication of a quality field-formed seam depends on

many material, environmental, and application factors, including

the shelf life of the adhesive, the thickness of the adhesive,

the membrane cleaning procedure, and the membrane temperature and

1
wetness. The effect of each of these factors should be known to

assure the fabrication of a quality field-formed seam. These

interactions can be unraveled through the application of welldesigned multivariate experiments (Box et al., 1978).

The objectives of this study were to determine the relative

importance of several fabrication variables (adhesive thickness,

cure time, level of surface contamination of the EPDM sheet,

mechanical load, and adhesive type) on the short-time strength

(sec. 3) and creep-rupture lives (sec. 4) of butyl-adhered EPDM

joints. These results were then compared with previously

published results for neoprene-adhered EPDM joints (sec. 5)

Only four factors were varied in each experiment, two of which

(adhesive thickness and level of surface contamination) were

common to both the short-time strength and creep-rupture

experiments. This redundancy was built into the experimental

design to determine if the rankings of these factors remained the

same when the strain rate was varied.

2 . BACKGROUND

The seams in this study were stressed in peel, as opposed to

lap-shear (see discussion in Rossiter, 1987) . Failure in peel is

implicated as the cause of many different types of seam defects,

some of which are shown in figure 1. Peel tests are also easier

to interpret on a scientific basis, are more sensitive to small

changes in the magnitude of different application variables

(Rossiter, 1988) , require lower stresses (Rossiter, 1987) , and

have shorter creep-rupture lives than do similarly fabricated

joints under comparable lap-shear loads.

The factors to be included in the experimental design were

selected based on a review of the literature and on extensive

conversations with EPDM membrane manufacturers, roofing

contractor representatives, and consultants. From these sources,

the removal of contaminants from the membrane prior to the

application of the adhesive was consistently identified as the

single most important factor for insuring the quality of a fieldformed seam. The contaminants of interest are the talc-like

particulates (talc, clay, or mica; hereinafter termed

contaminants) which are commonly applied over the surface of most

commercial EPDM membranes to prevent interply fusion during

vulcanization. These contaminants also act to deter interply

sticking during the storage and shipment of the cured rubber

(Wake 1969; Westley 1982) and to facilitate the movement of the

rubber sheets during roof construction. These positive

attributes are somewhat counterbalanced, however, by the negative

impact that they have on the formation of a strong adhesive bond

(Wake 1969; Rossiter 1988; Martin, Embree 1989).

In the following subsections, factors affecting the

mechanics of peel are reviewed. This review emphasizes systems

2
similar to those investigated in our experiments, namely,

uncrosslinked, elastomeric adhesives, having thicknesses less

than their critical adhesive thickness, which are stressed at a low peel rate and at a temperature far above the glass transition

temperatures of both the adhesive and EPDM.

2.1 Stress Distribution Along the Peel Line

Equations for cleavage stress and peel force have been

developed from small beam theory by Kaelble (1965) . In deriving

these equations, the adhesive and adherend were assumed to be

linear elastic and the stresses developed across the width and

the thickness of the specimen were assumed to be uniform and

constant. The resulting equations for cleavage stress, o^, at a distance, -x (see fig. 2) , from the peel front (where x=0) are

given by

= o (cos 6x + K^sin 6x)exp(6x) (1)

where

'adh

adh

= (Eadh^/^EIT^,,)^/^ = 6m/ (6m + sin 0) = boundary stress at x = 0 = tensile modulus of elasticity of the adhesive

= tensile modulus of elasticity of the adherend

= joint width

= moment of inertia

= moment arm

= peel angle

= adhesive thickness.

(2)

(3)

Equation 1 predicts that the cleavage stress is a highly damped

harmonic function consisting of alternating zones of tension and

compression. Equation 1 assumes that the joint does not carry

any stresses behind the peel front; that is, for values of

X ^ 0 (see fig. 2) . This latter prediction is contrary to

experimental observations, since joints stressed in peel often

form filaments between the separating peel strips. These

filaments are known to carry substantial tensile stress (Kaelble,

1965; Niesilowski, Aubrey 1981), as can be seen by comparing the

theoretical and actual experimental stress distributions in

figure 2.

2.2 Peel Force

The peel force, P, can be derived from the work equation of

peel (Kaelble, 1960, 1971) and is given by

3
P = WT3,,k/o^V(2E3,,(1-cos 0)). (4)

As is the case in our experiment, Kg = 1.0 when 0 = 9 0°, which

simplifies eq (4) to

From eq (5) , it can be seen that peel force per unit width is

equal to the strain energy per unit area, Tg^h^o^/'^-^adh ' ^hich in

turn is proportional to adhesive thickness T^^^.

Gent and Hamed (1977, 1978, 1982) have derived an

alternative peel force equation based on the energy balance of

detachment during steady-state peel. From their analysis they

concluded that the dependence of peel force on adhesive thickness

results from energy dissipation losses within the bulk adhesive.

The magnitude of these dissipation losses and, hence, the

magnitude of the peel force, increases linearly with adhesive

thickness up to a critical thickness; beyond which, the peel

force remains constant. The existence of a critical thickness

implies that energy can only be dissipated within a finite radial

distance from the peel line, beyond which energy dissipation

losses are not transmitted.

Gent and Petrich (1969) have earlier shown that these

dissipation losses account for virtually all of the energy

required to fail a joint. They deduced from their analysis that

the tensile strain energy of the adhesive provides an independent

approximation of peel strength. (Hereinafter, this approximation

will be referred to as Gent and Petrich 's approximation model .

Graphically, the tensile strain energy is given by the area under

the stress/strain curve up to the point of maximum elongation;

mathematically, it is computed from

where o„^^ is the maximum tensile stress and e is the tensile max strain of the adhesive.

2.3 Rate of Peel

Strain rate and temperature have long been known to have

equivalent effects on the viscoelastic response of a polymer.

These effects can be modeled via the Williams, Landell, and Ferry

(WLF) superposition principle, from which data collected at

different strain rates and temperatures can be reduced to a

single master curve by multiplying the data by a suitable shift

factor (see for example Ferry, 1961) . When the WLF-superposition

principle was first discovered, it was verified against

adh. (5)

(6)

4
homogeneous materials. Later Bright (1954) and Kaelble (1964)

applied the WLF-superposition principle to composite systems when

they established that the shift factors for cohesive and

interfacial joint failures were almost identical.

A typical peel force versus peel rate diagram (Gent and

Petrich 1969) is presented in figure 3. At low peel rates, the

macromolecules have time to disentangle and slip past one another

(that is, the adhesive displays liquid-like flow behavior)

resulting in a small peel force. As the peel rate increases,

however, the internal friction within the polymer increases,

since the macromolecules no longer have time to disentangle,

resulting in an increase in peel force. This increase continues

up to a critical peel rate, at which time an abrupt and

precipitous drop in the peel force occurs (see fig. 3) and, with

this drop, the failure mode changes from cohesive to interfacial.

This critical peel rate has been interpreted by Gent and Petrich

as a transition from liquid-like to rubber-like macromolecular

behavior. As the peel rate increases from the liquid-like to

rubber-like transition, the peel force again begins to increase

until the peel rate reaches another critical value, at which time

the peel force again drops. Gent and Petrich have interpreted

this second critical peel rate as a transition from rubber-like

to glass-like macromolecular behavior and attributed the loss in

peel strength to the decreased flexibility of the solid adhesive.

In the present experiment, however, the second transition is of

little interest, since the peel rates associated with the rubberlike to glass-like transition are much higher than a roof seam

would ever experience in service.

Gent and Petrich explained the complicated shape of the peel

rate versus peel force diagram (fig. 4) using their approximation

model. They assumed that the interfacial bond strength of a joint remains constant (the horizontal dashed line in fig. 4)

regardless of the peel rate or temperature (a reasonable

assumption according to Gent and Hamed (1978)). The observed

peel force, therefore, depends only on the strain rate of the

adhesive.

In figure 4, stress-strain curves are shown for an adhesive

stressed at three different strain rates. Failure of a joint

occurs whenever the maximum stress in the peel joint exceeds

either the tensile strength of the adhesive (in which case,

failure is by cohesion) or the interfacial bond strength (in

which case, failure is interfacial) . From eq (6) , peel force is

the area under the stress strain curve up to 1) the point of

maximum extension of the adhesive, in the case of cohesive

failure (e.g., e^^^^ and e^^^^ in fig. 4) or 2) the point that the

tensile strength of the adhesive exceeds the interfacial bond

strength of the adhesive to the substrate, in the case of

interfacial failure (e.g. , e^^g^j in fig. 4) . Curves I and II

correspond to an adhesive exhibiting liquid-like behavior, while

5
Curve III corresponds to an adhesive exhibiting rubber-like

elastic behavior. At the lowest peel rate (Curve I) , the tensile

strength of the adhesive is much less than the interfacial bond

strength of the joint, thus the joint fails cohesively. At a

slightly higher peel rate (Curve II) , the tensile strength of the

adhesive is still below the interfacial bond strength of the

joint, so the joint still fails cohesively, but its tensile

strain energy, and thus the observed peel force, is much greater

than that of specimen I. The peel force for specimen II is the

maximum force achievable for this adhesive system; that is, it

corresponds to the liquid-like to rubber-like transition in

figure 3. At a higher peel rate, (Curve III), the strength of

the adhesive surpasses the interfacial strength of the joint and

thus, the joint fails interfacially . The observed peel force for

specimen III, however, is much less than that of specimen II,

since specimen Ill's tensile strain energy is much lower than the

tensile strain energy for specimen II.

3. PEEL FORCE OF BUTYL-ADHERED EPDM JOINTS

Short-time strength tests are the most common method for

assessing the quality of an adhesive joint (for example, ASTM D

1876) and are an easy and convenient way to obtain results which

may be useful in assuring seam quality. The objectives of the

short-time strength experiments in this study were

* to determine whether the strength properties of joints

adhered with two commercial butyl-adhesives are

significantly different,

* to assess the sensitivity of joint strength to varying

amounts of talc-like contaminants at different cure times,

* to determine the locus of failure for uncontaminated and

contaminated joints,

* to determine the effect of adhesive thickness on peel

strength at different cure times and contamination levels,

and

* to determine if the peel strength of an EPDM butyl-adhered

joint is more sensitive to adhesive thickness or to the

presence of talc-like contaminants.

Each of these objectives is important in understanding the

failure mechanisms and providing the technical basis for

guidelines for the formation of reliable joints.

3 . 1 Experimental

3.1.1 Materials

Two of the most widely-used commercial butyl-splicing

adhesives were selected. The exact formulations of these

adhesives were unknown, except that they 1) contained a complex

blend of several different rubbers including an isocyanate

6
component (Chmiel and Cotsakis, 1985) , 2) had a low solidsfraction, and 3) had a low glass transition temperature.

Isocyanates cure in the presence of moisture and, thus, steps had

to be taken to minimize the effects of adhesive pre-cure on the

experimental results. To this end, the adhesive was removed from

its bulk container at the beginning of the experiment and poured

into a number of small containers, each of which was immediately

sealed and only reopened when needed. The solids-fractions of

the two adhesives were determined by heating three replicate

samples of each in an oven at 100 °C for 4 days. The average

solids-fraction of Adhesive 1 was 29.6% with a standard deviation

of the three replicates of 0.1%; while that of Adhesive 2 was

29.0 % with a standard deviation of the three replicates of 2.1%.

The glass transition temperatures of the adhesives and EPDM sheet

were measured using a thermal mechanical analyzer equipped with

an unloaded, 3.5 mm diameter probe and heated at a rate of 10

°C/min. The glass transition temperature of the cured adhesive

was -2 0 °C, while that of the EPDM sheet was -4 0 °C.

The 1.5 mm thick EPDM membrane material used in this

experiment was from one sheet, which was cut into approximately

800 joint strips, each having dimensions of 25 x 127 mm. Each

strip was cleaned by vigorously brushing both sides with soap and

water, after which it was thoroughly rinsed with water. The

cleanness of a strip was checked using the computer image

processing procedure described below (see sec. 3.1.3). If the

cleanness of the strip did not meet the predefined mean and

standard deviation criteria, it was either recleaned or

discarded.

After the strips were cleaned, they were matched into pairs,

given an identification number, and randomly assigned to one of

10 treatments (that is, 2 adhesives x 5 levels of contamination)

Each treatment contained 3 6 joint specimens.

3.1.2 Specimen Contamination

Within each pair of joint strips, one (designated Side A)

always remained clean, while the other (designated Side B) was

contaminated. Contamination was achieved by moving a cleaned,

Side B strip at constant velocity under an atomizer which

uniformly sprayed the strip with a well-mixed, low concentration

of commercial grade talc-like particles suspended in heptane (see

Martin and Embree (1989) for a detailed discussion of this

procedure) . The talc-like particles were obtained from the

manufacturer of the EPDM membrane material. Heptane was selected

mainly because it is a hydrocarbon solvent, akin to the splice

wash solvents commonly used in preparing field seams for

adhesion. It also had other desirable properties including high

volatility and the ability to wet the surfaces of both the talc

and the EPDM. The strips were contaminated to one of five levels

corresponding to zero, two, four, six, and eight passes under the

7
atomizer. The contamination levels were determined as described

in the next section.

3.1.3 Contamination Level Quantification

The level of contamination of each strip was quantified

using a computer image processing system, consisting of an 8-bit

mini-computer-based image processor, a computer-controlled X-Y

positioning table, a camera, and a voltage-regulated fiber-optic

tungsten light source (fig. 5) . The principle underlying the

measurement of cleanness or level of contamination is that the

reflection of tungsten light from the black EPDM surface

increases with increasing contamination of the surface by the

whitish talc-like particles. Reflectance was quantified in terms

of greyscale level, where a perfectly absorbing surface has a

greyscale value of zero and a perfectly reflecting surface has a

greyscale value of 255. The intensity of the radiation was

controlled by maintaining the electrical input into the fiberoptic light source at 120 ± 0.3 V via a voltage regulator, and by

fixing the positions of the optical fiber cables at approximately

15° from the vertical axis of the camera. Calibration curves for

relating greyscale value to the amount of talc-like particles

deposited (iig/cm^) and procedures for insuring that the greyscale

measurements remained in calibration are described in Martin and

Embree (1989)

Greyscale values were measured for 2 0 areas, each having

dimensions 5x5 mm, within the adhesion area of each T-peel

strip. From these 2 0 measurements, the average and standard

deviation of the greyscale values were computed. If the standard

deviation within a strip exceeded eight greyscale values, the

strip was replaced with one which was more uniform. For the 360

Side A strips (the clean side of a T-peel joint) , the average and

standard deviation of the greyscale values were 33.0 and 2.8,

respectively. For the 360 Side B strips (the contaminated side

of a T-peel joint), the average greyscale values were 33, 60, 90,

130, and 160 with a standard deviation of approximately 4.0. For

comparison purposes, uncleaned commercial sheets of EPDM received

directly from the manufacturer typically have greyscale values

ranging from 170 to 250. (Hereinafter, strips cut directly from

uncleaned, commercial sheet will be referred to as "as-received"

strips.) Using analysis of variance techniques, no statistical

difference at the 0.05 level of significance was found for the

mean greyscale values for strips assigned to the two adhesives

and at each contamination level.

3.1.4 Specimen Preparation and Test Procedure

Prior to applying the adhesive, the mass of each strip was

determined. As subsequently discussed (see fig. 6) , a highly

linear relationship exists between adhesive mass and thickness.

The mass of the adhesive, however, was experimentally easier to

8
determine. In the text which follows, therefore, "adhesive mass"

and "adhesive thickness" will be used interchangeably.

The butyl adhesive was applied to 7 5 mm of the strips

comprising a joint beginning from one end of each strip. The

adhesive was uniformly applied with a brush first to the clean

strip (Side A) and then the contaminated strip (Side B) , after

which the brush was discarded. This procedure was adopted for

every pair of strips to eliminate the transfer of talc. Once the

adhesive was spread, the solvent was allowed to evaporate at

laboratory conditions (approximately 24 °C and 45% relative

humidity) for approximately 3 0 min. The strips were then joined

and pressed in a hydraulic press at 700 kPa (approximately 100

psi) for 10 s to form a T-peel joint. When the two adhesive

layers were brought into contact with each other, the adhesive

macromolecules interdif fused or coalesced resulting in the

disappearance of the interfacial boundary between the two

adhesive layers (Voyutskii 1963; Wake 1969). For these

experiments, the start of the cure time was defined as the time

at which a joint was removed from the press; 12 cure times (0.5,

1.0, 1.5, 2, 4, 6, 24, 48, 96, 192, 768, and 1320 h) were used in

this experiment. The joints were cured at 24 °C and 45% relative

humidity.

Three replicates were tested at each cure time, at each

contamination level, and for each adhesive. The following

information was collected for each replicate: T-peel strength,

mass of the adhesive on the clean (Side A) and contaminated (Side

B) strips, and the mode of adhesive failure. Peel strength was

measured in a tensile testing machine at a crosshead speed of 50

mm/min.

The failure mode of a joint was assessed in two ways. The

first was a visual assessment in which the failure mode was

categorized as cohesive, interfacial (often termed adhesive

failure) , or a mixture of cohesive and interfacial failure. The

second was a quantitative assessment based on the premise that

the mass percent of adhesive remaining on the clean (Side A)

strip depended on the contamination level of the Side B strip.

Thus, if both the Side A and B strips were clean, 50% of the

adhesive, on the average, should remain on the Side A strip;

whereas, if the Side B strip was heavily contaminated, then a disproportionately large amount of adhesive should remain on the

Side A strip. The mass percent of adhesive on the Side A strip

was determined by weighing the separated strips (Mg^^.^+M^^^.^ and

^str-B"''^adh-B) subtracting the initial mass of the strips (M^^^.^

and Mg^^.g) , from which both the mass of adhesive remaining on each

strip (Mg^,^.^ and M^^^.g) and the total mass of the adhesive

remaining on both strips, M^^,^, could be determined. The mass

percent of adhesive on the Side A strip, M'^^,^.^, was computed from

9
M'adh-A = J^adh-A/ (Madh-A +^3,,.,) X 100. (7)

A comparison was made between the visual and mass percent

assessments from which it was ascertained that, on the average, a

mass percent of 60% corresponded to cohesive failure, 75%

corresponded to a mixture of cohesive and interfacial failure,

and 90% corresponded to interfacial failure.

A separate set of joints was prepared for establishing the

relationship between the mass (g) of the adhesive, M^^^^, and

adhesive thickness, T^^^. The mass of the adhesive was measured

according to the procedure discussed above. The adhesive

thickness was determined by subtracting the sum of the

thicknesses of the individual joint strips from the thickness of

the fully adhered T-peel joint measured after allowing it to cure

for 2 weeks. All thickness measurements were made with a

micrometer. The relationship between adhesive thickness and the

mass of the adhesive applied to each strip over an area of 25 X

7 5 mm was linear (see fig. 6) , having the form

Tadh = + ^adh (8)

where a„ and b are coefficients. Estimates of the coefficients m , m , . for the fit of eq (8) were (1) a^^^ = -29.83 with the estimated

standard error of the estimator of 7.29; (2) b^^ = 554.85 with the

estimated standard error of the estimator of 21.66. The standard

error of the estimate, or residual standard deviation, was 10.35

and the squared correlation coefficient, r^, for the fit of eq

(8) to the adhesive mass data was 99.8%.

Since peel strength is known to be proportional to adhesive

thickness (see eq (5) )

, it was important to assure that

variations in the adhesive thicknesses were not significantly

different among the five contamination levels. An analysis of

variance was conducted, from which it was concluded that the

adhesive thicknesses were not significantly different between

contamination levels at the 0.05 level of significance.

3 . 2 Model of Strength Versus Cure Time

An empirical model was presented in Martin and Embree (1989)

for modeling changes in the peel strength with time; this model,

called the Gompertz growth model, has the form

t is the cure time,

Sp, a, and b are coefficients; S^ represents the maximum

achievable peel strength of a joint; and 0 < a < 1 and

0 < b < 1.

10
Equation 9 can be transformed into its more familiar

distributional form (Ahuja & Nash 1967) , scaled by parameter S^,

by substituting

-a = In a (10)

and

-p = In b (11)

into eq (9) , and obtaining

S = S^e-«e"P^ (12)

where a and P > 0

Although empirical, the coefficients of the model can be

related to many of our experimental observations. For example,

at t = 0, eq (9) becomes

S = S a (13)

which is the minimum peel strength of the joint or equivalently

the cohesive strength of the butyl adhesive when the joint is

first removed from the press. As t - 0°, eq (9) becomes

S = (14)

the maximum strength of the adhesive joint. The logarithm of the

Gompertz parameter b is the rate at which the logarithm of the

logarithm of the scaled strength increases with cure time. This

can be shown by taking the logarithm of both sides of eq (9)

In S = In + (In a)b* (15)

substituting eq. (10) into eq (15)

(In - In S) = ab* (16)

and again taking the logarithm of both sides of eq (16)

ln(ln - In S) = In a + t In b. (17)

Thus, the change in the logarithm of the difference in the

logarithms of the maximum strength and the observed strength,

ln(ln - In S) , is linear with time.

3.3 Results

In a previous paper (Martin, Embree 1989) , it was concluded

that the strengths of joints adhered with the two commerciallyavailable adhesives were not significantly different at cure

times greater than 6 hours. In this paper, the data for the 2

11
adhesives are combined and reanalyzed and the effects of adhesive

thickness on peel strength assessed.

3.3.1 Failure Mode

The mass percent of the adhesive adhering to the clean (Side

A) strip is plotted against cure time at each of the five

contamination levels in figure 7. For joints com.prised of two

clean strips (that is, greyscale equal to 33) , failure was always

cohesive; whereas for joints comprised of one contaminated and

one clean strip, failure was cohesive for cure times of 2 hours

or less, a mixture of cohesive and interfacial failure for cure

times of 4 and 6 hours, and interfacial for cure times greater

than 6 hours.

Scanning electron microscopy (SEM) was used to determine the

locus of failure within a joint. In this analysis, the strips of

the failed joint were opened like a book so that their common

edges abutted and the failed surfaces of both strips could be

imaged simultaneously by the SEM at low magnification (4 OX) (see

fig. 9a) . In addition, lOOX magnification SEM micrographs of the

Side A and B strips were made in the same area as the 4 OX

micrographs (see figs. 9b and 9c)

For uncontaminated joints and for joints cured for less than

2 hours, the joints failed cohesively and thus the locus of

failure was within the adhesive layer (Location 1 in fig. 8)

For joints containing a contaminated strip cured for longer than

6 hours (see fig. 9) , the locus of failure was most often between

a talc-like particle and the EPDM substrate (Location 3 in fig.

8) or between two stacked talc-like particles (Location 4 in fig.

8) . Less often, the locus of failure was between a talc-like

particle and the adhesive (Location 2 in fig. 8)

3.3.2 Effect of Adhesive Thickness on Peel Strength

From eq (5) , peel strength S should change linearly with

adhesive thickness, and therefore, with adhesive mass M^^^^, since

the thickness of the adhesive and the mass of the adhesive are

linearly related (see fig. 6) ; that is.

where a^,^ and b^,^ are coefficients which are tabulated in table 1.

From figure 10, the slope, b^,^, of the regression line was

observed to be negative at short times and increase with cure

time, eventually becoming positive. A negative slope was

observed for joints cured for less than 2 hours (see fig. 10)

This was believed to be due to the presence of solvents (Gent,

Schultz 1972) or to the incomplete interdif fusion of the adhesive

macromolecules between the two adhesive layers (Voyutskii 1963,

12
1977) . At intermediate cure times (cure times between 4 and 6

hours) , the regression line is close to zero because of the

transition from a negative to a positive slope; for this reason,

the peel strength appears to be independent of adhesive

thickness. At long cure times (cure times greater than 6 hours)

peel strength increases with adhesive thickness as predicted by

eq (5) , except for the cure time at 192 h which was negative

because of one outlying value for a specimen having a thick

adhesive layer.

For contaminated joints, the trends in the a^,^ and b^,^

coefficients with cure time mimicked those for clean joints,

except that the scatter in the coefficient estimates was greater.

This scatter is believed to result from the confounding effects

that surface contaminants have on peel strength. These

confounding effects are discussed below.

Table 1. Peel strength versus adhesive mass coefficients for

joints containing two cleaned strips

Cure Time a^,^ se(a^|^)* b^^ se(b^i^) r^ * (h)

0.48 0. 54 0. 12 -0.30 0. 94 0. 16

0.96 0. 66 0. 07 -0.32 0. 38 0. 13

1.44 0.72 0. 07 -0.26 0. 09 0. 12

1.92 0. 89 0. 06 -0.36 0. 07 0. 87

4 . 08 0.99 0. 11 -0.27 0. 14 0.48

6. 00 0.72 0. 16 0.22 0. 22 0. 19

24 . 00 1. 11 0. 06 0.23 0. 09 0.62

48.00 1. 14 0. 17 0. 55 0. 27 0.52

96.00 0.90 0. 22 1.32 0. 45 0. 69

192 . 00 2 . 05 0. 11 -0.93 0. 25 0.78

768. 00 0. 58 0. 26 2.43 0. 64 0.78

1320. 00 0.76 0. 18 2 . 14 0. 43 0.86

se = estimated standard error of the estimator

* r^ = squared correlation coefficient

3.3.3 Strength versus Cure Time at Different Contamination

Levels

Strength data at the five levels of contamination are

plotted against cure time in figure 11. Superimposed on each

data set is the predicted Gompertz growth curve (eq (9) )

, the

parameters of which are shown in Table 2 . In computing these

13
statistics, the strength data were not corrected for adhesive

thickness and thus, part of the variation observed in the peel

strength data in figure 11 can be attributed to adhesive

thickness variations.

The Gompertz parameters, a, b, and S^, along with the

predicted minimum peel strength, S^a, are plotted against

greyscale level in figure 12 and were fitted to the equations

described below.

The estimated values of the Gompertz parameter a at the five

contamination levels were modeled by

^ = ^.axd - ^"') (19)

where

G is the observed greyscale value,

^max asymptotic value of the Gompertz parameter a as

greyscale value, G, approaches the maximum

value of 255, and

c is a coefficient which was chosen for best fit.

The estimated coefficients were a„,„ = 0.437 with an estimated max . standard error of the estimator of 0.004, and c = 0.0319 with an

estimated standard error of the estimator of 0.00115. The

squared correlation coefficient, r^, was 99.5%.

Table 2 . Gompertz growth curve parameter estimates at each of

the five contamination levels

Gompertz growth curve parameters

Nominal

greyscale So se(SJ* a se (a) b se(b) r2 #

33 1 . 573 0. 021 0. 286 0. 014 0. 911 0.008 0. 96

60 1 .242 0. 021 0. 367 0. 018 0. 925 0. 010 0. 91

90 1 . 021 0. 018 0. 419 0.023 0. 898 0. 018 0. 87

130 0 .836 0. 015 0. 430 0. 032 0. 839 0. 029 0. 82

160 0 .703 0. 019 0. 432 0. 052 0. 767 0. 054 0. 64

se = estimated standard error of the estimator

* r^ = squared correlation coefficients

14
The estimated values of the Gompertz parameter b, the rate

of change in strength, remained relatively constant over the

range of contamination levels studied, and were modeled by

b = f(e-hG (20)

where

G is the observed greyscale level and

f, h, and k are coefficients which were selected for best

fit.

The estimated coefficients were f = 1.069 with an estimated

standard error of the estimator of 0.0377, h = 0.002 with an

estimated standard error of the estimator of 0.00029, and k = 0.0705 with an estimated standard error of the estimator of

0.0105. The squared correlation coefficient, r^, was 97.4%.

Equation 20 is different from that originally proposed in Martin

and Embree (1989) . When the individual adhesives were analyzed,

it appeared that the value of b remained constant at 0.94.

The estimate values of the Gompertz parameter, S^, the

maximum strength of a joint, at the five contamination levels

were modeled by

G is the observed greyscale level, and

n and p are coefficients which were selected for best fit.

Estimates of the coefficients were n = 1.857 with an estimated

standard error of the estimator of 0.0396 and p = 0.00615 with an

estimated standard error of the estimator of 0.00037. The

squared correlation coefficient, r^, for the fit of eq (21) to

the estimated values was 99.4%. From figure 12, the maximum

strength decreases by greater than 50% over the range of

experimental contamination levels.

The minimum peel strength of a joint, S^a, is the product of

eqs (19) and (21) and has the form

ne-pG (21)

where

S^a = q(e"'"'^ - e'^'^) (22)

where

G is the observed greyscale level and

q, r, and s are coefficients.

15
The estimated coefficients were q = 0.681 with an estimated

standard error of the estimator of 0.058, r = 0.00497 with an

estimated standard error of the estimator of 0.000628, and s = 0.04 09 with an estimated standard error of the estimator of

0.00863. The squared correlation coefficient, r^, for the fit of

eq (15) to the estimated S^a data was 99.3%.

3.4 Discussion

From figure 11, it can be seen that the Gompertz growth

curves fit the strength data well and that the strengths of the

two adhesives are not significantly different.

A comparison of the sensitivities of peel strength to

contamination and adhesive thickness can be made from figures 10

and 12, respectively. From figure 10, at long cure times, the

difference in the strength of joints at long cure times and over

the range of adhesive thickness is 0.2 kN/m (approximately 1

Ibf/in) . From figure 12, the maximum strength, S^, of the peel

joints decreased by more than 50% over the range of contamination

levels; that is, strength decreased by approximately 0.9 kN/m

(5.1 Ibf/in) over the experimental range of contamination levels.

Thus, contaminants affect peel strength to a much greater extent

than adhesive thickness.

Superimposed on figure 12 are the peel strengths for joints

fabricated from an "as-received" strip (that is, from a strip cut

from a commercial EPDM membrane) , which had greyscale values

greater than 2 00 and which were cured for longer than 48 h.

Extrapolation of the values of via eq (21) to these higher

"as-received" greyscale levels resulted in a predicted strength

which was greater than the observed strengths. This discrepancy

may be due to the non-uniform distribution of contaminants on the

surface of the "as-received" strips or from cleaning the surface

of the laboratory prepared strips prior to applying the

contaminants

From the visual inspection and SEM micrographs, contaminants

appear to act by blocking the formation of adhesive bonds with

the EPDM substrate. The presence of contaminants reduces the

effective load-carrying width of a joint (W in eq (5)). Based on

this assumption, our results can be graphically explained using

Gent and Petrich's approximation model. In figure 13, only one

stress-strain curve for the adhesive is displayed, since the peel

rate was kept constant. The interfacial strength of the joints

(the horizontal lines in fig. 13) decreases with increasing

contamination, since the contaminants effectively reduce the

joint width and thereby, the load-carrying ability of the joints.

For clean joints, the tensile strength of the adhesive was less

than the interfacial bond strength of the joint, so the joints

failed cohesively. For contaminated joints, the tensile strength

16
of the adhesive exceeded the interfacial strength, so the joints

failed interfacially . As the level of contamination increased,

the interfacial strength of the joint decreased and, with this

decrease, peel strength decreased.

3.5 Peel Strength Conclusions

The sensitivity of the strength in peel of butyl-adhered

EPDM joints to surface contaminants, adhesive thickness, cure

time, and adhesive type were experimentally determined. From the

experimental results and for the range of conditions covered by

them, it is concluded that:

1. The gain in strength with cure time for clean and

contaminated joints can be predicted using the Gompertz

growth model

2. For cure times less than 2 hours, the strengths of

clean and contaminated joints depend on the cohesive

strength of the butyl adhesive and not on the level of

contamination

3. For cure times greater than 2 hours, the interfacial

strength of butyl-adhered EPDM joints was highly

dependent on contamination level and the contaminated

joints failed interfacially . We believe that the

contaminants act by preventing the adhesive from

bonding with the substrate, thus effectively reducing

the load-carrying width of a joint.

4. The maximum achievable strength, S^^, for butyl-adhered

EPDM joints occurs when both strips are clean. The

maximum peel strength of joints containing one

contaminated and one clean strip decreased to half of

the strength of our clean joints over the range of

experimental contamination levels.

5. Surface contaminants decrease the peel strength of our

peel joints more than adhesive thickness increases

strength.

6. No significant difference was observed in the strengths

of the two commercial butyl adhesives at long cure

times

4. CREEP-RUPTURE TIMES-TO-FAILURE OF BUTYL-ADHERED EPDM JOINTS

The major difference between short-time strength and creeprupture experiments is one of strain rate. We believe that the

low strain rates inherent in creep-rupture experiments are more

representative of the strain rates experienced by an in-service,

field-formed seam. To date, few creep-rupture experiments have

been performed (Martin et al. 1987); instead, short-time strength

experiments are preferred (Westley 1982; Dupuis 1983; Rossiter et

al. 1988; and Watanabe and Rossiter 1990). Before the results of

short-time peel strength experiments are accepted, it is

important to insure that the effects of each experimental

17
variable on the short-time strength and on the creep-rupture life

are the same. Otherwise, performance requirements may be imposed

which have little to do with the in-service performance of

roofing seams.

The objectives of the creep-rupture time-to-failure

experiments were

* to determine the effect of contamination level, adhesive

thickness, cure time, and mechanical load on the creeprupture life of butyl-adhered EPDM joints stressed in peel,

* to determine the failure mode of the contaminated and

uncontaminated joints,

* to derive a mechanistically-based model for predicting the

time-to-failure of a joint stressed in peel, and

* to compare the rankings of the principal experimental

variables as they affect short-time peel strength with the

corresponding rankings for creep-rupture times-to-failure.

4 . 1 Experimental

The experimental procedure for the creep-rupture experiments

was basically the same as that used in the short-time strength

experiments. Therefore, only differences are highlighted.

4.1.1 Materials

Six hundred pairs of EPDM peel strips were randomly assigned

to one of 2 0 treatment combinations (5 contamination levels x 4

mechanical loads) . Each treatment combination was applied to 30

replicate joints. The five levels of contamination were the same

as those used in the short-time strength experiments; that is,

the average greyscale levels were 33, 60, 90, 130, and 160. The

four creep-rupture loads were 0.154, 0.231, 0.308, and 0.386

kN/m. Only one adhesive, the one designated Adhesive 1 in the

short-time strength experiment, was used, since the performance

of the two commercial adhesives used in the short-time strength

experiment was not significantly different, especially at long

cure times. The other experimental materials, the methods for

cleaning and contaminating the surface of the peel strips, the

procedure for quantifying the level of contamination, and the

methods for preparing the joints were the same as those described

in section 3.1, except that the experimental materials came from

a different production lot.

4.1.2 Experimental Procedure

The creep-rupture apparatus was designed to load 12 0 joints

at once; thus, the 600 joints were separated into five batches.

In creating a batch, one contamination level was randomly

selected for each of the four mechanical load levels. The joints

within each batch were randomly assigned to the 120 positions in

18
the apparatus. The starting and ending times of the joints were

monitored electronically. The starting time was the same for all

of the joints assigned to a batch; the ending time, however, was

unique to each joint and occurred when the strips completely

separated. Separation caused the deactivation of the electronic

clock assigned to that joint and the recording of its time-tofailure. The accuracy of the recorded time-to-failure was within

1 second. All of the creep-rupture experiments were conducted at

24 °C ± 1 °C and 45% ± 2% relative humidity.

Prior to the start of a new batch, all of the joints in the

batch currently under load were allowed to fail; hence, the cure

times differed from batch-to-batch. (The cure times for the five

batches in this experiment were 1224, 1562, 1825, 2161, and 2449

h, respectively) . It was assumed that differences in cure time

did not adversely affect the chemical or physical stability of a joint. To test this hypothesis, three joints were randomly

selected from each of the four treatments assigned to a batch and

their short-time strengths determined using the procedure

described in section 3.1.4. The remaining 27 joints within each

treatment or, equivalently , 108 joints within each batch (27

joints per treatment x 4 treatments per batch) were loaded to

their assigned creep-rupture load and the time-to-failure of each

joint observed. Besides time-to-failure, the following data were

collected for each creep-rupture joint: 1) a visual assessment

of the failure mode, 2) the mass of the adhesive remaining on

both the clean (Side A) and the contaminated (Side B) strips, and

3) the mass percent of the adhesive remaining on the clean (Side

A) peel strip.

4.2 Results

4.2.1 Peel Strength at Long Cure Times

The peel strengths, S, of the joints removed for short-time

strength determination from each contamination-mechanical load

treatment were determined at the end of the creep-rupture

experiment. The strengths of the joints within each

contamination level were then plotted against cure time (see fig.

14) and fitted to a linear regression model of the form

S = a^ + b^t (23)

where a^ and b^ are coefficients, the values of which are shown

in Table 3 along with an estimated standard error of the

estimators and the mean strength for all of the joints having the

same level of contamination.

19
Table 3. Creep-Rupture Results—Linear regression coefficients

for changes in strength with cure time. The cure times ranged

from 1224 h to 2449 h. The coefficients are computed in units of

kN/m.

nominal

greyscale

it se(aj se(b,)

41

ff ± s^

(kN/m)

33 1. 61 0.072 -2. 10X10'•b 3.75X10 T- accept Ho 1. 57±0. 27

60 1. 56 0.15 -2. 98X10'5

8 . 10X10 5 accept Ho 1. 50±0. 26

90 1. 27 0.12 4. 60X10 •5 6.37X10 •5 accept Ho 1. 36±0. 24

130 1. 08 0.10 5. 23X10 •5 5.88X10 •5 accept Ho 1. 16±0. 20

160 1. 43 0.21 2. 10X10 5

1. 00X10 4 accept Ho 0.99±0. 17

* se = estimated standard error of the estimator

* H^: = 0. Null hypothesis, H^, that the slope of the

regression line, b^, is different from zero at the 0.05

level of significance.

* x^±s = the mean and standard deviation in the strengths of all

joints assigned to a contamination treatment level.

The stability of the joints was assessed by determining if

the slope of the regression line, b^, was significantly different

from zero. Using a standard statistical analysis procedure (the

Student t test), at the 0.05 level of significance, the strengths

of the joints assigned to each contamination level did not change

significantly between 1224 to 2449 h of cure, an outcome

consistent with that predicted by the proposed strength versus

cure time model (eq (9)).

The mean strengths for the creep-rupture joints are plotted

against contamination level in figure 15 and the values compared

with the maximum strength estimates, S^, for the short-time

strength experiments. From figure 15, the mean strengths of the

clean joints (greyscale = 33) were the same in both experiments,

but the mean strength of the contaminated creep-rupture joints

were higher than the corresponding mean strengths from our shorttime strength experiments. The reason for this discrepancy is

not understood.

4.2.2 Failure Mode

The rate of peel of all of the creep rupture joints was

highly s, that is, no stick-slip behavior was observed. The mass

percent of the adhesive adhering to the clean (Side A) strips is

plotted against time-to-failure in figure 16. From visual

20
Table 4. Creep Rupture Results—Linear regression coefficients

and estimated standard error of the estimators for changes in the

mass percent of the adhesive remaining on Side A versus cure time

at each contamination and mechanical load level

contam.

lev

load

lev

se(a^) se(b,) H : b =0*

0 A

33 1 47. 67 1. 38 0. 03 0

02 accept 49 . 7±2 . 6

33 2 48. 33 1.72 0. 05 0. 10 accept 49 . 1±2 . 8

33 3 50

65 1 . 87 -004 0. 30 accept 50 . 4±2 . 3

33 4 48 . 73 1.20 0

26 023 accept 49 . 6±2 . 3

60 1 50. 36 0.79 0. 17 0

01 rej act 51. 6±2 .4

60 2 54 . 15 2.27 -0. 15 0. 17 accept 52 . 2±3 . 1

60 3 50

14 2 . 04 033 0. 32 accept 52 . 8±2 . 4

60 4 53 . 18 2 . 11 -0

1 3 56 accept 52 . 7±2 . 6

90 1 53 . 51 0.92 0. 0 1 02 accept 54 . 1±2 . 6

90 2 53 . 04 1. 37 0. 10 0

12 accept 54.1±2 .5

90 3 54 . 14 1.70 0

15 0. 33 CAa f-r'pi'n't" *—' V—- V—54 . 9±2 . 0

90 A 58 79 2 03 -1 11 0. 7

1 V-^ \^ 55 P + 3 7

130 1 58 . 5 1 . 63 -0

10 0. 04 rej ect 54 . 8±3 . 4

130 2 54 . 6 2 . 16 0. 26 0. 23 accept 57 . 0±2 .8

130 3 55. 5 4.31 0. 46 0. 93 accept 57.6±6 . 1

130 4 77. 8 5. 02 -5. 93 1. 84 rej ect 61. 9±4 . 8

160 1 66. 8 1.87 -0. 12 0. 04 rej ect 60. 9±4 . 1

160 2 70. 4 1.51 -0. 87 0. 14 reject 61. 9±5 . 3

160 3 77. 4 3 .33 -2 . 89 0. 87 rej ect 66.5±4 . 1

160 4 85. 2 3.06 -7. 12 1. 23 rej ect 68.0±5 . 4

* se = estimated standard error of the estimator

* H^: b^ = 0. Result of testing the null hypothesis that the

slope of the regression line, b^, is zero at the 0.05 level

of significance.

* x^±s^ = the mean and standard deviation of the mass percent

adhesive remaining on the clean (side A) strip for the

joints assigned to a contamination/mechanical load level.

21
inspection of the failed peel strips and from the mass percent

measurements (see criteria in sec. 3.1.4), it was concluded that

all the creep-rupture joints failed cohesively regardless of

their level of contamination or applied mechanical load.

To determine if the mass percent of adhesive remaining on

the clean (Side A) strip depended on the time-to-failure of a

joint, the data were fitted to a linear regression equation of

the form

^Adh-A = ^A + V (24)

where a^ and are coefficients, the values of which are

tabulated in Table 4 along with an estimated standard error of

their estimators and the average and standard deviation of the

mass percent of the adhesive remaining on the clean (Side A)

strip within each treatment level. A statistical analysis was

performed (the Student t test) to determine if the slope of the

regression line, b^, was different from zero at the 0.05 level of

significance. From this analysis, the null hypothesis that the

slopes were zero could not be rejected, except at very high

levels of contamination where the slopes were negative. As

discussed in the next section, contaminants only affected highly

contaminated joints having short times-to-failure and thin

adhesive layers.

4.2.3 Creep-Rupture Times-to-Failure

The dependence of the time-to-failure on applied stress has

been studied for a wide variety of materials including metals,

polymers, and polymeric fibers (Bartenov and Zuyev 1968; Martin

1982) . From the analysis of the creep-rupture data for these

materials, two empirical models have repeatedly been found to fit

the time-to-failure data, t , as a function of applied stress, o^.

The general forms of these models are as follows:

T = a^ exp (-bgO^) {the "exponential model"} (25)

and

t = ao^'*^ {the "power law model") (26)

where a^, b^, a, and b are coefficients. For joints stressed in

peel, the applied stress, o^, is given by (see eq. (5))

^0 = (4PE3,,/WT3,J'^^ (27)

where, for the present, it is assumed that Young's modulus, E^^,^,

and the width of a joint, W, are constants, while the creeprupture load, P, and the thickness of the adhesive, T^^^'

measurable variables.

22
From the literature, the exponential model tends to provide

a better fit to the time-to-failure data for polymers stressed at

temperatures below their glass transition temperature; while the

power law model tends to provide a better fit to the time-tofailure data for polymers stressed at temperatures above their

glass transition temperature (Martin 1982) . However, this glass

transition temperature criterion is not inviolate, since, for

example, some researchers have used the exponential model to

obtain satisfactory fits to the time-to-failure data for polymers

stressed above their glass transition temperature (Martin, 1982)

For this reason, the time-to-failure data were fitted to both

models and a selection was made between the two models on the

basis of their fit to the data, even though our experiments were

conducted at a temperature well above the glass transition

temperatures of both the membrane and adhesive.

Substituting eq (27) into eq (25) , the exponential model

becomes

T = a^ exp(-b^(4PE3,,/WT3,j'^^) . (28)

Taking the logarithm of both sides of eq (28)

log^o T = log,, a^ - 2.303 h^{^^^,^^/^T^^^y' (29)

which can be simplified further to

log^o t = log,, a^ - b^' (P/T^,,)'^^ (30)

when Eg^j^ and W are assumed constant and where

b^- = 2.303 bJ4E3,,/W)\

The power law model can be reformulated by substituting eq

(27) into eq (26) yielding

^ = ^ (4PE3,,/WT^,,)-^/2^ (31)

Taking the logarithms of both sides and using the simplifying

assumptions, eq (31) becomes

log,,x = a' - b/2 log^^ P/T^,, (32)

where a' = log^Qa - b/2 log^Q ^^adh/^'

In figure 17, both the exponential and the power law models

are fitted to all of the time-to-failure data (number of joints =

540) . The power law model (eq (31) ) fit the time-to-failure data

much better than did the exponential model. The coefficient

estimates of the power law model (eq (31)) are (1) a' = 1.10

with an estimated standard error of the estimator of 0.007, (2)

b = 5.13 with an estimated standard error of the estimator of

23
0.03 4. The standard error of the estimate, the residual standard

deviation, is 14.2 and the squared correlation coefficient, r^,

for the fit of eq (31) to the time-to-failure data is 91.3%.

The power law model, eq (31), predicts that the time-tofailure of a joint decreases with an increase in the mechanical

load and/or a decrease in the thickness of the adhesive. The

dependence of the times-to-failure on mechanical load is welldocumented (Martin et al. 1982; Bartenev et al. 1969). The

dependence of the times-to-failure on adhesive thickness are

presented in figure 18. From figure 18a, it is evident that the

time-to-failure of a peel joint could be greatly increased by

even incrementally increasing the mass (and, thereby, the

thickness) of the adhesive. To verify this prediction, three

joints, comprised of two cleaned strips and having very thick

adhesive layers, were prepared and subjected to a mechanical load

of 0.308 kN/m. The masses of the adhesive at the time-of-failure

of each joint were 0.69, 0.79, and 0.83 g and the times-tofailure were 1030, 1469, and 788 h, respectively. These timesto-failure are much greater than the times-to-failure predicted

by substituting the adhesive mass into the power law model (the

predicted times-to-failure were 99, 140, and 159 h respectively),

indicating that the power law model underestimates the real

improvement in the creep-rupture life of joints when the adhesive

thickness is increased.

Contaminants may also affect the time-to-failure of a joint

(see fig. 16) . It was assumed that contaminants reduce the

effective load carrying width of a joint in proportion to the

level of contamination. Thus, the joint width term, W, in eq

(31) was modified as follows:

^effective = W {G^,,-G^,,) /G^^^ (33)

where 0 < c^ < 1 and where

G,^^ = observed contamination level and act G^^^ = maximum level of contamination for all the max experimental joints (here, G^^^ = 170) . Substituting eq (33) into eq (31) and taking the logarithms of

both sides yields an equation of the form:

log,, X = h, + l°gio[(G.ax-Gact)/Gn,ax] + V2 log^^ [P/^adh] (^4)

where b^ = log^^ a - b/2 log^g 4E + b/2 log^^ c^ W. Estimates of

the coefficients in eq (34) (number of joints = 540) are (1) b^ = 1.2 8 with an estimated standard error of the estimator of

0.0068, (2) ^2 ~ ~1«89 with an estimated standard error of the

estimator of 0.017, and (3) b^ = -0.02 with an estimated

standard error of the estimator of 0.01. The residual standard

deviation is 0.10 and the squared correlation coefficient, r^, is

24
95.8 %. The effect of contaminants on time-to-failure is

statistically significant, but, relative to the thickness of the

adhesive and the mechanical load, the effect of contaminants on

the creep-rupture life of a joint is minor.

4.3 Discussion

The debonding of all of the creep-rupture peel joints

progressed at a constant velocity; that is, the peeling process

was stable. The times-to-failure of our adhesively bonded joints

are highly dependent on the magnitude of the mechanical load and

adhesive thickness and to a much lesser extent on surface

contamination. This ranking of variables is different from that

for our short-time strength results which were highly dependent

on the level of contamination, and to a much lesser extent, on

the thickness of the adhesive. This transposition in the ranking

of adhesive thickness and contamination results from the

differences in peel rates between short-time strength and the

creep rupture experiments.

At low strain rates, the adhesive exhibits liquid-like flow

behavior. The tensile strength of the adhesive is small and, for

our experimental conditions, the tensile strength of the adhesive

never exceeds the interfacial bond strength of the adhesive to

the substrate; that is, the joints fail by cohesion.

Graphically, these results can be explained using Gent and

Petrich's approximation model. As the level of contamination

increases, the interfacial strength of the joints decreased, as

shown by the horizontal lines in figure 19. The effect of

increasing the creep-rupture load is equivalent to increasing the

strain rate of the adhesive; the effect of which is represented

by the four stress-strain curves in figure 19. Since all of the

creep-rupture joints failed cohesively, the stress-strain curve

corresponding to the highest mechanical load must be below the

lowest interfacial strength line. Presumably, if the joints had

been contaminated to a higher level, the strength of the adhesive

would have eventually exceeded the interfacial strength of the

joint resulting in interfacial failure.

4.4 Creep-Rupture Time-to-Failure Conclusions

The creep-rupture sensitivity of butyl-adhered EPDM joints

to surface contaminants, adhesive thickness, cure time, and

mechanical load were experimentally determined and the

significance of each factor on time-to-failure ranked. From the

creep-rupture data, for the range of experimental conditions, it

is concluded that:

1. The strengths of the creep-rupture joints remained

constant over the range of experimental cure times.

25
2. The failure mode for all the creep-rupture joints was

cohesive regardless of contamination level or applied

mechanical load. The rate of crack-growth for the

creep-rupture peel joints was also highly s; that is,

no stick-slip behavior was observed.

3. A mechanistically based model, depending on mechanical

load, contamination level, and adhesive thickness, was

derived and fitted to the creep-rupture data. The fit

of the model to the data was quite good and from this

analysis, it was concluded that the creep-rupture life

mainly depends on mechanical load and adhesive

thickness and to a lesser extent on contamination

level. Thus, the ranking in importance of the two

principal experimental variables (level of

contamination and adhesive thickness) in the creeprupture experiment were reversed from their ranking in

the short-time strength experiment. This reversal is

attributed to changes in the rheological properties of

the adhesive when stressed at different peel rates.

4. Performance requirements for seams should be based on

creep-rupture results as opposed to short-time strength

results

5. COMPARISON OF TIMES-TO-FAILURE FOR BUTYL-AND NEOPRENEADHERED EPDM JOINTS AS A FUINCTION OF LOAD RATIO

It is common in engineering practice to minimize the

probability of failure of a structural element by assigning a

working or design load, P„ork,rig/ to the element which is only a

fraction of its yield load, Py^g^^j. Practically, this is achieved

by reducing the yield load by a factor of safety, FS; that is,

FS = Pyield/Pworking (35)

where in industrial applications, it is typical to apply factors

of safety whose product is between one and three (Boresi et al.

1978) . (Applied loads are directly proportional to the applied

stresses; thus, the working load and yield load can just as

easily be expressed as a working stress and yield stress.) The

factors of safety account for unexpectedly high in-service loads

and for reductions in the resistance properties of a structural

element resulting from statistical variations in the material

properties and from variations introduced during fabrication (see

for example, Ellingwood et al. 1980). For viscoelastic

structural materials (such as wood, geosynthetic materials, and

seams in EPDM roofing membranes) , the greatest reduction factor

compensates for the creep-rupture effect; that is, the ability of

viscoelastic materials to sustain only a fraction of their yield

loads for sustained periods of time.

26
Table 5. Creep-Rupture Results—Stress ratios and the mean and

standard deviations of the times-to-failure for the butyl- and

neoprene-adhered EPDM creep-rupture joints. Data for the

neoprene-adhered EPDM joints was published in Martin, Embree, and

Bentz (1987)

contain. stress C4. V ^ • O >— J_ O O std

greyscale lev. load strength ratio ttf ttf

level (kN/m) (kN/m) (%) (h) (h)

butyl-adhered EPDM

J J 0. 154 1.57 10 66. 60 25. 60

T 1 J J z 0.231 1.57 15 15.70 5. 30

J J J 0. 308 1. 57 20 6. 14 1. 30

33 4 0. 386 1. 57 25 4 . 80 1. 90

50 ± 0. 154 1. 50 10 77. 10 52 . 50

50 Z 0.231 1. 50 15 13 . 00 3 . 70

60 3 0. 308 1. 50 21 6.41 1. 13

60 4 0. 386 1. 50 26 3 . 65 0. 89

90 1 0. 154 1.36 11 48.40 31. 00

90 2 0.231 1.36 17 10.70 4 . 00

90 3 0. 308 1.36 23 4.94 1. 19

90 4 0.386 1.36 28 2.81 0. 77

130 1 0. 154 1. 16 13 35.40 14 . 10

130 2 0.231 1. 16 20 9 . 06 2 . 35

130 3 0. 308 1. 16 27 4 . 50 1. 44

130 4 0. 386 1. 16 33 2.70 0. 43

160 1 0. 154 0.99 16 49.07 18 . 80

160 2 0.231 0.99 23 9.72 4 . 65

160 3 0.308 0.99 31 3.75 0. 78

160 4 0.386 0.99 39 2 . 42 0. 56

neoprene-adhered EPDM

33 1 0.080

33 2 0.155

33 3 0.235

33 4 0.315

0.75 11 199.5

0.75 22 5.8

0.75 31 1.3

0.75 42 0.6

27
The reduction factor for creep-rupture is found by plotting

the mean time-to-failure of a roofing seam against the applied

load ratio, LR = Pappiied/'^yieid.' equivalently , the stress ratio)

and selecting that load ratio (stress ratio) having a mean timeto-failure which is equal to or greater than the expected design

life of the component. In figure 20, the load ratios of several

viscoelastic building materials are plotted against mean time-tofailure (the data for the butyl- and neoprene-adhered EPDM peel

joints are shown in table 5)

From figure 20, the stress ratio versus mean time-to-failure

plots are greater for 1) semi-crystalline materials, such as

wood stressed in bending (Wood, 1951) and high density

polyethylene (HDPE) stressed in lap-shear (Halse et al. 1989)

than for 2) amorphous materials, such as neoprene and butyladhered EPDM seams loaded in peel at temperatures far above their

glass transition temperatures. For example, wood has been tested

down to 65% and HDPE to 45%. Straight-line extrapolation of the

wood and HDPE curves would indicate that wood is able to sustain

approximately 60% of its short-time stress while HDPE can only

sustain 20% for a design life of 10 years. On the other hand,

curve-line extrapolation of neoprene- and butyl-adhered EPDM peel

joint data indicate that these materials can only sustain 5% of

their yield load for this period of time. Thus, it is concluded

from figure 2 0 that the reduction factors for creep-rupture are

the same for both neoprene-and butyl-adhered EPDM seam joints

when stressed in peel and that only a very small peel load is

required to cause in-service seam failures.

6 . SUMMARY

The most frequently reported defect in EPDM single-ply

roofing systems occurs in the field-formed seams. Seam defects

occur most often within the first 3 years after installation of

the roof and may result from weaknesses in the materials,

deficiencies in the fabrication technique, or as a result of

environmental stresses. The objectives of this study were to

determine the importance of several material and fabrication

variables as they affect the short-time strength and creeprupture life of butyl-adhered EPDM joints stressed in peel and,

where possible, to derive mechanistically based models. The

joints were stressed in peel because this stress mode has been

implicated in several commonly observed field-formed seam

defects. This is not to imply that peel stresses cause all seam

defects, for other stress modes (e.g., lap shear) may also be

important

From the short-time strength experimental results, joints

failed cohesively for cure times less than 2 hours, failed

cohesively and interfacial failure for cure times between 4 and 6

hours, and failed interfacially for cure times greater than 6

28
hours. The gain in strength of the joints as a function of cure

time was modeled with the Gompertz growth curve equation. From

the analysis of our strength data, peel strength was highly

dependent on cure time and the level of contamination and to a lesser extent on adhesive thickness.

From the creep-rupture experimental results, the joints

always failed cohesively regardless of the level of contamination

or the magnitude of the mechanical load. A mechanistically based

model, which was functionally dependent on mechanical load,

adhesive thickness, and contamination level, was fitted to the

time-to-failure data. Times-to-failure were highly dependent on

mechanical load and adhesive thickness and to a much lesser

extent on the level of contamination. One prediction of this

model was that an increase in the thickness of the adhesive layer

above the experimental thicknesses could result in a great

increase in the service life of a joint.

Although the magnitude of the peel loads to which fieldformed seams are subjected are unknown, it is possible to

determine from the creep-rupture results the upper limit of peel

stresses which can be sustained by a joint over time. Our

experimental results indicate that butyl and neoprene-adhered

EPDM joints stressed in peel can only sustain less than 5% of

their yield stresses over the design life of a roof. Thus,

field-formed seams should be designed to minimize peel loads.

Where this is not possible, our results indicate that the service

life of seams can be significantly improved by minimizing the

magnitude of the peel loads and increasing the thickness of the

adhesive and, to a lesser extent, by insuring the cleanness of

the seams.

Our experimental results have identified several fabrication

(surface cleanness) , material (adhesive thickness and cure time)

and stress (peel load magnitude) variables as having a

significant effect on the service life of EPDM roof seams. As

everyone knows, who works with materials, especially building

materials, the variables examined in this experiment are only a

subset of the material, fabrication, design and stress variables

which may affect the service life of EPDM seams. In future

experiments, the other variables can be interjected to determine

which have a significant effect on a roofing seam's service life.

7 . ACKNOWLEDGMENTS

The authors acknowledge the help of their colleagues. Jack Lee

and John Winpigler, for their invaluable assistance in helping to

conduct the experiments and, Willard Roberts, for measuring the

glass transition temperatures of the EPDM membrane and the butyl

adhesive

29
8 . REFERENCES

Ahuja, J.C. and Nash, S.W. (1967) "The generalized GompertzVerhulst family of distributions," Sankhya. Series A 29:141.

ASTM D 1876, (1983) "Standard test method for peel resistance of

adhesives (T peel test)," Annual Book of ASTM Standards, Vol

15.06, ASTM, Philadelphia, PA.

Bartenev, G.M. and Zuyev, Yu. S. (1968) Strength and Failure of

Visco-Elastic Materials . Pergamon, New York.

Boresi, A. P., Sidebottom, O.M. , Seely, F.B., and Smith, J.O.

(1978)

Advanced Mechanics of Materials , Wiley, New York.

Box, G.E.P., Hunter, W.G., and Hunter, J.S. (1978) Statistics for

Experimenters , Wiley, New York.

Bright, W.M. (1954) "The adhesion of elastomeric pressuresensitive adhesives: rate processes," in Adhesives and Adhesion , J. Clark, J.E. Rutzler, and R.L. Savage [Eds.], Wiley, New York.

Chmiel, C.T. and Cotsakis, D.J. (1985) "Adhesive for bonding

cured EPDM rubber," U.S. Patent 4,501,842.

Cullen, W. (1984) "Survey compares practices to problems" Roofing

Spec. 12 (10) , 37

Cullen, W. (1988) "Project pinpoint: NRCA's diagnostic

database." Prof. Roofing 16(7), 17.

Cullen, W. (1989) "Interest runs high in project pinpoint's data

bank." Prof. Roofing 17(7), 29.

Dregger, P. and Ellingson, W. (1988) "What seams to be the

problem." Roofer Magazine 8(2): 35.

Dupuis, R.M. (1983) "Analysis and design of adhesive lap splices

for elastomeric single-ply membranes." Proc. 7th Conf. on Roofing

Technology, Joint National Bureau of Standards/National Roofing

Contractors Association Symposium, held at NBS, Gaithersburg, MD.

Ellingwood, B. , Galambos, T.V. , MacGregor, J.G. , and Cornell,

C.A. (1980) "Development of a probability based load criterion

for American National Standard A58," National Bureau of Standards

NBS Spec. Publ. 577.

Ferry, J.D. (1961) Viscoelastic Properties of Polymers , Wiley,

New York.

30
Gent, A.N. and Hamed, G.R. , (1977) "Peel mechanics of adhesive

joints," Polym. Eng. Sci. 17(7): 462-466.

Gent, A.N. and Hamed, G.R. (1978) "Adhesion of elastomers, with

special reference to triblock copolymers," Rubber Chem. Tech. 51:

354

Gent, A.N. and Hamed, G.R. (1982) , "Some rheological aspects of

the peel strength of rubbery adhesives," Rubber Chem. Tech.

55(2) : 483-493.

Gent, A.N. and Petrich, R.P. (1969) "Adhesion of viscoelastic

materials to rigid substrates," Proc. Roy. Soc. A, 310: 433-448.

Halse, Y.H., Koerner, R.M. , and Lord, A.E., Jr. (1989)

"Laboratory evaluation of stress cracking in HDPE geomembrane

seams," in Durability and Aging of Geosynthetics . R.M. Koerner

[ed.], Elsevier, New York, p. 177.

Johnson, J.E., Dupuis, R.M. , and Lee, J.W. (1977) "Physical

properties of BUR membranes tested at low rates of loading,"

Proc. of the Symposium on Roofing Technology, Joint NBS/NRCA

Symposium held in Washington, D.C., Sept. 21-2 3.

Kaelble, D.H. (1960) "Theory and analysis of peel adhesion: bond

stresses and distributions," Trans. Soc. Rheology 4: 45-73.

Kaelble, D.H. (1964) "Theory and analysis of peel adhesion:

rate-temperature dependence of viscoelastic dependence of

viscoelastic interlayers , " J. Colloid Sci. 19: 413.

Kaelble, D.H. (1965) "Peel adhesion: micro-fracture mechanics of

interfacial unbending of polymers," Trans. Soc. Rheo. 9(2): 135-

163

Kaelble, D.H., (1971), Physical Chemistry of Adhesion . Wiley, New

York, pp 423-432.

Lord, A.E. and Halse, Y.H. (1989) "Polymer durability — the

materials aspects," in Durability and Aging of Geosynthetics , R.M. Koerner [ed. ] , Elsevier, New York, p 293.

Martin, J.W. (1982) "Time transformation functions commonly used

in life testing analysis," Dur. Building Mat. 1: 175-194.

Martin, J.W. and Embree, E. (1989) "Effects of contaminants and

cure time on EPDM single-ply joint strength." J. Mat. Civ. Eng.

1(3): 151.

31
Martin, J.W., Embree, E., and Bentz, D.P. (1987) "Effect of

temperature and stress on the time-to-failure of EPDM T-peel

joints." Proc. 8th Conf. on Roofing Technology, Joint NBS/NRCA

Conference held at the National Institute of Standards and

Technology, Gaithersburg, MD 20899, April 16-17, 1987.

Niesiolowski , F. and Aubrey, D.W. (1981) "Stress distribution

during peeling of adhesive tapes," J. Adhesion 13: 87-98.

Rossiter, W.J. (1987) "Tests of adhesive-bonded seams of singleply rubber membranes," Roofing Research and Standards

Development, ASTM STP 959, R.A. Critchell [ed. ]

, American Society

for Testing and Materials, Philadelphia, pp 53-62.

Rossiter, W.J. (1988) "Further investigation of the effect of

application parameters on adhesive-bonded seams of single-ply

roof membranes." Materials and Structures 21: 24 3.

Rossiter, W.J. and Seller Jr., J.F. (1988) "Report of roof

inspection: partial delamination of adhesive-bonded seams at an

army facility." National Institute of Standards and Technology

NISTIR 88-3893.

Rossiter, W.J. and Seller, Jr., J.F. (1989) "Results of a survey

of the performance of EPDM roofing at army facilities." National

Institute of Standards and Technology NISTIR 89-4085.

Russo, M. (1988) "Single-ply seaming: the state of the art."

Roof ing/Siding/Insulation 65(8): 20.

Stucker, N.E. and Higgens, J.J., (1977) "Butyl Rubber and

Polyisobutylene , " Chap. 16, Handbook of Adhesives , I. Skeist

[ed.]. Van Nostrand, New York.

Voyutskii, S.S. (1963) Autohesion and Adhesion of High Polymers . Interscience, New York.

Voyutskii, S.S. and Ustinova, Z.M., (1977) "Role of autohesion

during film formation from the latex," J. Adhesion 9: 39-50.

Wake, W.C., (1969), "Elastomeric adhesives," in Treatise on

Adhesion and Adhesives , Vol. 2, Materials, R.L. Patrick [ed.].

Marcel Dekker, NY.

Watanabe, H. and Rossiter, W.R. Jr. (1990) "Effects of adhesive

thickness, open time, and surface cleanness on the peel strength

of adhesive-bonded seams of EPDM rubber roofing membranes," to be

published in Roofing Research and Standards Development , 2nd

Volume, T.J. Wallace and W.R. Rossiter [eds.], ASTM STP 1088,

American Society for Testing and Materials, Philadelphia, PA.

32
Westley, S.A. (1982) "Bonding ethylene propylene diene monomer

roofing membranes: the theory and practice of adhering

vulcanized ethylene propylene diene to itself." in Single-ply

Roofing Technolocfv . ASTM STP 790, Gompertz, W.H. [ed. ]

, American

Society for Testing and Materials, pp 90-108.

Wood, L.W. (1951) "Relation of strength of wood to duration of

load," USDA Forest Service Paper No. 1916.

Wu, S. (1982) Polymer Interface and Adhesion . Marcel Dekker, New

York.

33
9. LIST OF SYMBOLS

—In a

FQ -In b

5 (E W/4EIT „ ) ^^adh*TM/ ^^-^-^adh^

h pin c; 1 1 cj-h >-a in r>"F hhiP' arlhipcjiv^

& = peel angle

= cleavage stress

ft yj \J LI 1 iV_4. CI i_ y O L« J_ O

0

max THr^viTniTm 'hp"nc;'i 1 p ci1~T"^^c;c^ "Fot" "Hip ?^ f^hpci "i vp

t

JQj U^lio_LXt^ lUL^vUUlX Ulo vJ X ci: X Cl o X L.* X U> jr v^X L-ilt^ ClVJ.Xit:::X

^adh L. v3 1 1 O J L \^ ili\J \A L4. X L4. O w X \3 X CI »3 X X L» V wX L>i. X dLAllCOXVw

TJ. Tnmnnp"n1" trP "i npyl" "i 111 will ^1 1 L.* w X XI X W X d.

ill IlK^illt^Il U Ctx ILL M^^adh I^Cl X lUCl AO ^m/X wii^ d^ilCoXVC^ Clk.^kJXXCL4. K^KJ jLJXIiL^

adh lUd O O ^t:?X V^C-llL- v^X L-XIC Cllk.4.XXCDXVC? dLAXXCXXXlV^ \^\J L-XXt^ O XUC 2^ D l,^X XL/

M Tn^^cici Xlld OO 0"P \J ±. 't'Vio L-IXC ^^HViOcinwo dLJ.lXCOXV^ ^^rlHoyiTirT dUiXXCXXXX^ "Ho \^\J "HV^o L-XiC^ OXLACf Qirlca A c:^>"1k^ O L.X X

M

^^adh-B ILLdoo KJ X. Uli^ dLlXx^oXVc? dLliiCfX Xll^ \^\J l^ilc? O XLlt^ D o L.X X^

Mstr-A

111 i-X 4^ V-/ J_ L«X -L. \jL X. ^ M^^str-B Xlld k-D w3 V-/ X. L»X X w L_J _L V-4. XJ O X X

p ^^t^x XvJXv-«c:^

X. oL^UdX^d v^tJx X tJxd UXUll L'Uc:^ J- J- XOJ.t:^il U c WX/O^XVCLA ^CCX OL»XC^XX^L-Xx

So Goinpertz parameter for iQaximum peel strength

"^adh adhesive thickness

t time

X location parameter

W joint width

34
Figure 1. Seam defects in which a peel-type failure mode is

implicated: a,b) ripples in the EPDM membrane and c)

T-joint. It is not clear what the failure mode is for

the seam defect pictured in d)

35
Figure 2. Distribution of stresses along upper half of peel

specimen.

36
LOG PEEL RATE (cm/sec)

Figure 3 . Typical peel force versus peel rate diagram for a

flexible adherend adhesively bonded with an

uncrosslinked elastomeric adhesive (after Gent and

Petrich, 1969)

37
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TOTAL MASS OF ADHESIVE (g)

0.6

Figure 6. Calibration curve for adhesive thickness versus

adhesive mass. Joints cured for two weeks prior to

making measurements.

40
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41
Figure 8. Schematic representation of failure locations: 1)

failure within the adhesive layer which is typical in

the case of cohesive failure, 2) failure at the

adhesive/talc-like particle interface, 3) failure at a

talc-like particle/EPDM interface, and 4) failure at a

talc/talc interface.

42
Figure 9. SEM micrographs of a failed peel joint cured for 13 00 h

which was made from a strip which was contaminated to a

greyscale of 90. In Micrograph A, the two strips of

the joint were "bookmatched" and imaged at 4 OX (the

clean strip is on the right-hand side) . In Micrographs

B (the clean strip) and C (the contaminated strip) , the

disjoined surfaces were imaged at lOOX. Note that most

of the talc-like particles are on the clean side.

43
p

c

Q)U

Q)

•H

<D

0)

5-1

4-> +J

asm

m

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(LU/N>I) HiONayiS 1336

U

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« 8 O OD

CO ^ ^ d o

(UJ/N>|) Hi3N3BiS

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•HP

(dC

•H

rO PCou

4-1o>

•H

<4-l +J

(0

<D •H

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0)u

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2.0

c/)

^ 1.6

-J

<

>

DC

h- 1.2

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cr

Hi

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Butyl-adhered EPDM

As received

peel strengths 7^

50 100 150 200

GREYSCALE VALUE

250

Figure 12. Gompertz parameter estimates versus contamination

level. Parameters a and b are dimensionless

46
Greyscale

STRAIN

Figure 13. Schematic representation of the effect of contamination

on peel strength when the peel rate is kept constant.

47
TjoCO

Q

UJ

n 03

CD O

am D

o

LUmO

o

LU

m o

X

lU

UJz>o

o

1- OJ h- d d

(UJ/NM) HiDN3HiS 133d

CO <q e>j

(UJ/NM) HiONSUiS 133d

O

m

0)>

0)

0)>

•H

xi

o

O

<0

0)

-P

<0

0)e

•H

+J

Q)U

O

W

W

>

•P o C +J

0) fO

^1

4-J CH

g

dJ C

(U O

eu 0

0)

•H

48
2.0

E

1, 1.5

1.0

0.5

Butyl-adhered EPDM

J

X ± s, creep-rupture experiment a

A X, short-time strength experiment

30 60 90 120

GREYSCALE

150 180

Figure 15. Peel strength versus greyscale level for the

creep-rupture and short time strength joints.

49
rao

•Hc

(0

3

M

^1

Q)> O

•H C

^1 o

+J -H W +J

(d

<; -H

0) g

T3 O

-HWw (1) >

C -H

(0 4J

Q) fO rH -P O C

<D

QJ W

Xi Q) P ^

U

Q)

c

•Hc (1) •H > c

0) 0

g M •H

CD 4J U -P (0

(t3 C

0) •H >

-H to W 3 -P

(U iH c

45 •H 0

TS (d o

fO M-l

1 -d 0 c

c -p to

Q) 1 0 Q) T! e nJ

(U •H 0

dlblS NV310 NO 3NINIVVM3H HQV iN30a3d

rH

0)

^H

-H

50
Io

I

O O o o o o o O LT) O lO O If) CO CO CM T-

{^) aaniivd-oi-ai^ii

H

Q)U

•H

51
(M) 3amivd-oi-3wii

(M) 3amivd-oi-3i^ii • CO

•H

+J

rO iH C

-H

u e

Q) fd > -p

<u c w 00 O T3c

<d

w

Q) (n C o O

•H rH

n3 -P 0

•H

0)

C > fO W 0

Q) Q) X5

TS

<C <4-l 0

rH

fO

w P c O 0

+J -H

-p

W (dc W •H u Sig > 00 U 0)>

rH •H -H +J

to (d 1 c 0 CD •p w 1 Q)e a

•H a) Eh u

00 (O

(M) 3UnilV3-Oi-3l^ll

00H

•H

52
CO

CO

LU

cr

Ico

0.386 kN/m

Greyscale -- 33

--60

-- 90 -- 130 -- 160

0.231 kN/m

0.308 kN/m ^0.154 kN/m

STRAIN

Figure 19

Schematic representation of the creep-rupture

experimental results

53
2.0

Butyl-adhered EPDM

CO

^ 1.6

_i

<

>

Lli H

LU

1.2

<

<

^

0.8 N

DC

LU

Q.

O 0.4

CD

0-

0

Sna

As received

peel strengths

50 100 150 200

GREYSCALE VALUE

250

igure 20. Time-to-failure versus stress or load ratio for

several viscoelastic materials.

54
NIST-114A U.S. DEPARTMENT OF COMMERCE

(REV. 3-89) NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY

RIRI inriPAPHIP RATA QHPPT

1. PUBUCATION OR REPORT NUMBER

NIST/BSS-169

2. PERFORMING ORQANIZATION REPORT NUMBER

3. PUBUCATION DATE May 1990

4. TITL£ AND SUBTITLE Strength and Creep-Rupture Properties of Adhesive-Bonded EPDM Joints Stressed in Peel

5. AUTHOR(S)

Jonathan W. Martin, Edward Embree, Paul E. Stutzman, and James A. Lechner

6. PERFORMINQ ORQANIZATION (IF JOINT OR OTHER THAN NIST, SEE INSTRUCTIONS)

U.S. DEPARTMENT OF COMMERCE

NATIONAL INSTITUTE OF STANDARDS AND TECHNOLOGY

QAITHERSBURQ, MO 20899

7. CONTRACT/QRANT NUMBER

8. TYPE OF REPORT AND PERIOD COVERED

Final

9. SPONSORINQ ORQANIZATION NAME AND COMPLETE ADDRESS (STREET, CITY, STATE, ZIP)

Same as item #6

10. SUPPLEMENTARY NOTES

DOCUMENT DESCRIBES A COMPUTER PROGRAM; SF-185, FIPS SOFTWARE SUMMARY, IS ATTACHED.

11. ABSTRACT (A 200-WORO OR LESS FACTUAL SUMMARY OF MOST SIGNIFICANT INFORMATION. IF DOCUMENT INCLUDES A SIGNIFICANT BI8UOGRAPHY OR

UTERATURE SURVEY, MENTION IT HERE.)

The most frequently reported defect in ethylene-propylene-diene terpolymer (EPDM)

single-ply roofing systems is in field-formed joints. The causes of these joint defects

are largely unknown, but they tend to occur a short time after a roof is installed.

The objectives of this research were 1) to determine the importance of the following

material and fabrication variables in causing joint failures: adhesive thicJcness, cure

time, mechanical load, adhesive type and surface cleanness, 2) to ranlc these variables as they affect the creep-rupture life and strength of butyl-adhered EPDM joints, and 3) to

determine the maximum peel stress which EPDM joints can sustain over their design life.

From our results, cure time and level of cleanness of the EPDM membrane have the

greatest effect on joint strength; while the thicJcness of the adhesive and the magnitude of

the mefchanical load have the greatest effect on a joint's creep-rupture time-to-failure.

Thus, the rankings of variables in creep-rupture and short-time strength experiments were different and this difference should have important implications in standards writing and in

estciblishing performance requirements for EPDM roofing joints.

Creep-rupture results were used in determining the maximiam design stress that butyl

and neoprene-adhered EPDM joints can sustain when loaded in peel. The maximum sustainable

stress is only a small fraction (less than 5%) of the short-time yield strength of a joint.

Thus, efforts should be made to insure that the field-formed joints are only subjected to

small peel loads.

12. KEY WORDS (S TO 12 ENTRIES; ALPHABETICAL ORDER; CAPITAUZE ONLY PROPER NAMES; AND SEPARATE KEY WORDS BY SEMICOLONS)

butyl adhesive; creep-rupture life; EPDM; neoprene; peel performance criteria; seam;

single-ply roofing; strain rate; stress-ratio; time-to-failure

13. AVAILABIUTY UNUNITED

FOR OFFICIAL DISTRIBUTION. DO NOT RELEASE TO NATIONAL TECHNICAL INFORMATION SERVICE (NTIS). X

ORDER FROM SUPERINTENDENT OF DOCUMENTS, U.S. GOVERNMENT PRINTING OFFICE. WASHINGTON, DC 20402. ORDER FROM NATIONAL TECHNICAL INFORMATION SERVICE (NTIS), SPRINGFIELD. VA 22161.

14. NUMBER OF PRINTED PAGES

59

15. PRICE

ELECTRONIC FORM



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