THE EFFECTS OF MOLECULAR WEIGHT ON THE SINGLE LAP 

SHEAR CREEP AND CONSTANT STRAIN RATE BEHAVIOR OF 

THERMOPLASTIC POLYIMIDESULFONE ADHESIVE 

Stanley K. Dembasky and Erol Sancaktar 

CLARKSON UNIVERSITY 

Potsdam, New York 13676 

Grant NAG- 1-284 

February 1985 

REIGBl (NASP-CB-1812e3) CI THE EPPEC’IS OF IfIOLECULAB N8 7-27 7 S 5 THE SINGLE LAP SEEdF CEEEP All? 

CCISTABT STRAIN BATE EEAAVICF CP 

!IEERMCPLASSIC FCLYIHIDESULFCLI ACHESIVE Unclas 

(Clarkson Uniw.) 75 F Avail: b3XS HC G3/27 0093165 

-* c 
NOMENCLATURE 

- 

Mn number average molecular weight 

MU weight average molecualr weight 

- 

one dimensional shear strain and shear strain rate 

TYf one dirrneional shear stress and 8haar atrears rate 

Y,? 

0 

‘I: elaatic limit (yield) shear strese 

Y alaatic limit (yield) shear strain 

Y 

Y 

GO?G elastic shear nodulus 

G1 viscoelastic ehear modulus 

n viscosity coefficient 

K,n material constants for viscosity equation 

D,m,a,b, 

r,d material constants for nonlinear equations 

T ultimate shear stress ult 

e elastic limit shear strees 

0 elaetic limit ehear strain 

k,n material constants 

0 plane stress 

E; plane atrain 

X9K tine dependent material properties 

A’,B’,C’ material constanta of Crochet’s equation 

A?B?C material constants of Crochet’s equation in shear 

and of creep relations based on Crochet’s equation 

in shear 

T level of constant stress 0 

Y<t> time dependent streas 
LIST OF FIGURES 

Figure 

1 

2 

3 

4 

5 

6 

9 

10 

11 

12 

The Viscoelastic Chase-Goldsmith Model Used 

to Describe Adhesive Behavior. 

Clip-on Gage Attachment to the Single Lap 

Specimen for Adhesive Deformation 

Measurement. 

Constant Strain Rate Stress Strain Behavior 

of Thermoplastic Polyinidesulfone Adhesive 

and Comparison with Theory. 

Variation of Elaetic Shear Modulus with 

Molecular Weight for Thermoplastic 

Polyimidesulfone Adhesive. 

Variation of Viscosity with Molecular 

Weight for Thermoplastic Polyimidesulfone 

Adhesive. 

Constant Strain Rate Stress Strain Behavior 

of Thermoplastic Polyinidesulfone Adhesive 

and Corpariaon with Theory. 

Constant Strain Rate Stress Strain Behavior 

of Thermoplastic Polyinidesulfone Adhesive 

and Comparison with Theory. 

Variation of Ultimate Shear Stress with 

Molecular Weight for Thermoplastic 

Polyimidesulfone Adhesive. 

Variation of Maximum Shear Stress with 

Initial Elastic Strain Rate for Different 

End-Cap Forms of Thermoplastic Polyirnidesulfone Adhesive. 

Variation of Maximum Shear Strain with 

Initial Elastic Strain Rate for Different 

End-Cap Forms of Thermoplastic Polyimidesulfone Adhesive. 

Constant Strain Rate Stress Strain Behavior 

of Thermoplastic Polyinidesulfone Adhesive 

and Comparison with Theory. 

Constant Strain Rate Stress Strain Behavior 

Page 

10 

26 

28 

30 

32 

33 

35 

3a 

39 

40 

42 
13 

14 

15 

16 

17 

18 

19 

20 

21 

22 

23 

24 

of Thermoplastic Polyinidesulfone Adhesive 

and Comparison with Theory. 43 

Constant Strain Rate Stretss Strain Behavior 

of Thermoplastic Polyiaidesulfone Adhesive 

and Conparlaon with Theory. 

Creep Behavlor of Three Percent End-Cap 

Thermoplastic Polyinidesulfone Adhesive 

at 700F (2lOC) Condition. 

Creep Behavior of Three Percent End-Cap 

Thermoplastic Polyimidesulfone Adhesive 

at 2500F (121OC) Condition. 

Creep Behavior of Three Percent End-Cap 

Thermoplastic Polyinidesulfone Adhesive 

at 350oF (177OC) Condition. 

Creep Behavior of One Percent End-Cap 

Thermoplastic Polyinidesulfone Adhesive 

at 700F (210C) Condition. 

Creep Behavior of One Percent End-Cap 

Thermoplastic Polyialdesulfone Adhesive 

at 250oF (1210C) Condition. 

Creep Behavior of One Percent End-Cap 

Thermoplastic Polyimidesulfone Adhesive 

at 350OF (177OC) Condition. 

44 

4a 

49 

50 

51 

52 

53 

Variation of Creep Strains at 50 Minutes 

with Holecular Weight for Thermoplastic 

Polylnidesulfone Adhesive. 54 

Creep Rupture Data and Comparison with 

Theory for Three Percent End-Cap Thermoplastic Polyimidesulfone Adhesive. 58 

Creep Rupture Data and Comparison with 

Theory for One Percent End-Cap Thermoplastic Polyimidesulfone Adhesive. 59 

Creep Rupture Data and Comparison with 

Theory for Zero Percent End-Cap Thermoplastic Polyinidesulfone Adhesive. 60 

Variation of Maximum (Safe) Creep Stress 

with Molecular Weight for Thermoplastic 

Polyimidesulfone Adhesive. 61 
25 Variation of Haxinun (Safe) Creep Stress 

with Environmental Temperature for Different 

End-Cap Forms of Thermoplastic Polyiridesulfone Adhesive. 62 
ABSTRACT 

The bonded shear creep and conetant strain rate 

behaviors of zero, one, and three percent endcapped 

Theraoplaetic Polyinidesulfone adhesive were examined at 

room and elevated temperatures. Endcapping was accomplished 

by the addition of phthalic anhydrides. Single lap 

specimens with titanium adherends were provided by 

NASA-Langley Research Center. Mechanical testing was 

performed at Clarkson University. 

The primary ob3ective was to determine the effects of 

molecular weight on the mechanical properties of the 

adhesive. Viscoelaetic and nonlinear elastic constitutive 

equations were utilized to model the adhesive. Ludwik’e and 

Crochet’e relatione were used to describe the experimental 

failure data. The effect8 of molecular weight changes on 

the above mentioned mechanical behavior were assessed. 

The viscoelastic Chase-Goldsmith and elastic nonlinear 

relations gave a good fit to the experimental stress strain 

behevior . Crochet’s relations based on Maxwell and 

Chase-Goldsmith models were fit to delayed failure data. 

Ludwik’s equations revealed negligible rate dependence. 

Ultimate stress levels and the safe levels for creep 

stresses were found to decrease as molecular weight was 

reduced . 
CHAPTER 1 

INTRODUCTION 

The characterization and optimization of adhesives is 

essential to the production of bonded joints which are 

strong and reliable. Structural adhesives are widely 

employed by aerospace, automotive, naval and other 

industries. Adhesives provide a practical means for joining 

lightweight materials such as composites and metals , while 

reducing stress concentrations and maintaining structural 

integrity. 

The model adhesive for the present invesigation, 

Thermoplastic Ployimidesulfone, is a novel thermoplastic 

adhesive currently under development at NASA-Langley 

Research Center. It is eolvent resistant and can be used in 

thermal environments with temperatures up to 450 OF (232 OC> 

c13. 

In an attempt to maximize mechanical strength and 

bonding properties, while reducing moisture penetration C23, 

NASA has provided specimens of varying molecular weight for 

mechanical testing at Clarkson University. Molecular 

weights were varied by the addition of phthalic anhydride 

end-caps, The degree to which changes in molecular weight 

affects the viscoelaetic properties of Thermoplastic 

Polyimidesulfone will be studied at room and high 

1 
terperaturrr. Single lap samples (ASTM D1002 standard 

geometry) are used because they are common to many 

structural fabrications where adhesives are used. 

The derigner of adherively bonded joints nust take into 

account three posrible nodee of failure: 

1) failure of the adhesive layer when the rupture 

atrers is reached under monotonic loading; 

2) delayed failure caused by creep under a sustained 

load; 

3) adherend failure. 

Adhesive failure in the creep and constant strain rate 

modes, tho effects of molecular weight upon these modes of 

failure, and the adheeive’s stress strain behavior in 

general, will be discussed in this paper. Adherend failure 

is not a prominent consideration for the specimens used. 

Obi ect ive 

The purpose of this investigation is to provide insight 

to the effects of molecular weight on the creep and constant 

strain rate behavior of Thermoplastic Polyimidesulfone. The 

basic objectives are to: 

*obtain experimental data on the delayed failure 

behavior of zero, one, and three percent endcapped 

Thermoplaetic Polyinideaulfone-titanium single lap 

specimens; 

2 
*obtain experimental data on the constant shear strain 

rate behavior of the epecimens; 

odeterrine if Crochet's equation describes the delayed 

failure observed; 

*identify a vi6coelaetic or nonlinear nodel which will 

deecribe the mechanical behavior of the adhesive; 

'determine if Ludwik's equation is applicable; 

*determine the effect8 of molecular weight on the above 

mentioned mechanical properties of Thermoplastic 

Polyimidesulfone baaed on experimental data. 

3 
Literature Review 

The magnitude of their molecular weights is known to 

affect the mechanical strength of adhesives. The effects 

vary from material to material, and are most evident in the 

cohesive (pertaining to the cohesive matrix) and adhesive 

(pertaining to the interface) strengths of the adhesive. 

The two most commonly used expreeaions for molecular 

weight are the number average (fin) and, weight average (ZW) 

molecular weights as defined below: 

C NiMi 

i 

I: N, 

- 

Mn = 

C NiMi 

i 

I: N, 

- 

Mn = 

i 

(1) 

In equations (1) and (2) Ni and Mi refer to the number of 

molecules and the molecular weight of each species, 

respectively. The number average molecular weight can be 

calculated from boiling point, freezing point, vapor 

pressure, osmotic pressure, and end-group analysis of a 

polymer; whereas the weight average molecular weight is 

found by diffusion, light scattering, or ultracentrafugation 

C33. Although many authors fail to state which molecular 

weight relation is being used, Alfrey suggests that the use 

4 
of the number average molecular weight is more uaeful from 

the concept of polymer strengths. The number average gives 

preferential weighting to the effects of low molecular 

weight segments (even a small fraction of low molecular 

weight material represents a large number of molecules) 

C43. 

Very little research has been performed on the effects 

of molecular weight on adhesives in the bonded form, where 

joint failures may occur at the adhesive-adherend 

interface. Tsuji, Masuoka and Nakao studied the effects of 

molecular weight on the peel strength of polyisobutylene 

used to bond polyethylene films 153. Their results show 

increasing peel strength with increaaing molecular weight up 

to a maximum value. Then, with further increases in 

molecular weight, the peel strength decreases abruptly due 

to adhesive/adherend interfacial failure. This behavior is 

attributed to reduced wetting and adhesive behavior of the 

polymer at increased molecular weight valuea. 

When an adhesive joint fails within the adhesive , and 

not at the adhesive-adherend interface, the mode of failure 

is cohesive. Zhurkov presented a theory stating creep 

rupture failures in the cohesive mode to be dependent on the 

rupture of chemical bonds within the polymer C61. 

Increasing creep strength will result from increasing 

molecular weight because a material of higher molecular 

weight yields more chemical bonds. Zhurkov’s experiments on 

5 
Nylon 6 fibres supported his theory. The limitation of 

Zhurkov’s theory was the conditions under which he performed 

the experiments. The Nylon 6 fibres used were highly 

oriented. Orientation increases the intermolecular forces 

between the monomeric segments of neighboring molecules, and 

even though the intermolecular forces are much weaker than 

the primary valence forces within a chain, addition of 

intermolecular forces can become great enough to cauee 

rupture to occur within the chains E43. Thie tends to 

decrease the chances of failure due to stress 

concentrations, chain ends, chain slippage, and other 

factors which may be affected by changes in molecular weight 

and molecular weight distribution. In crosslinked and 

network stuctures with chain entanglements (usually common 

to structural adheaives), the polymer will not have a high 

degree of orientation, permitting fracture to initiate at 

other points besides the chemical bonds. 

Because of the finite length of the polymer chains and 

imperfections in network geometry, initial breaking of 

intermolecular bonds is most likely until the stress can be 

carried by the remaining chain structure which is strong 

enough to carry the applied stress C73. The force applied 

to a bulk adhesive must be percieved first by the 

intermolecular bonds. The disintegration of these bonds 

will largely determine the strength of the adhesive. 

Therefore, an adhesive with lower molecular weight will 

6 
reach a level of catastrophic failure at lower stress levels 

due to the decreased number of intermolecular bonds. 

It is possible that an optimum molecular weight can be 

reached for an adhesive. Margolies C83 suggestr that 

increaring molecular weight beyond an optimum value 

increases chain entanglement to the point where chain 

straightenin-g under an applied stress is greatly inhibited. 

This results in chemical bond failure with the possibility 

that the strength of many of the polymer chains is not being 

utilized. Decreasing the molecular weight below the optimum 

point may lead to failure by chain slippage. In this case a 

chain orientation which is prone to slippage will be 

achieved more rapidly because of decreased viscosity and 

lower intermolecular forces. 

Fellers and Kee C93 report that the number average 

molecular weight of polystyrene must be greater than 2Me for 

the material to be useful. Me is stated to be the chain 

entanglement molecular weight. At molecular weights below 

2Me the network of entangled chains is insufficient to 

sustain plastic flow before failure. This leads to low 

tensile strength and brittle fracture. Above 2Me the 

polystyrene shows increasing rupture elongation and tensile 

stress with increasing molecular weight. 

Although molecular rupture has been interpreted as 

being a process of flow, reorientation, or bond rupture, 

these phenomena are not clearly separated 1103. Without 

7 
specific regard to the microscopic behavior of a material, 

Alfrey t43 offers the opinion that the dependence of 

mechanical strength of a high polymer upon molecular weight 

is qualitatively the same for all polymers. Tensile 

strength is proportional to molecular weight, increasing as 

molecular weight increases, up to a limiting value. Once a 

certain molecular weight is reached, further increase in 

molecular weight does not result in appreciable improvements 

in strength. Alfrey states that low molecular weight 

polymers break mainly by pulling apart entangled chains, 

while high molecular weight specimens fail largely by the 

rupture of primary valence chains. 

Gardner and Martin C113 observed behavior similar to 

Alfrey’s qualitative explanation. Lexan 141 samples were 

subjected to molecular weight reductions by humid aging, and 

then tested in tension. Humid aging reduced the molecular 

weight by partial dieintegration of the polymer chains. The 

results indicate a critical molecular weight value, below 

which, tensile strength dropped off rapidly with decreases 

in molecular weight. Above the critical molecular weight 

the tensile strength showed little increase with molecular 

weight. The results of the study also revealed a transition 

from ductile to brittle failure as the molecular weight was 

decreased. This was attributed to the effects of annealing 

caused by the aging process used. The elongation also 

decreased rapidly as the molecular weight decreased, and no 

8 
indication was given that the elongation will reach a 

plateau at higher molecular weights. 

Vlachopoulos E123 reports an increase in rupture 

elongation with increasing molecular weight. Experiments 

with monodisperse <En= Mw> polystyrene exhibited increasing 

tensile strength and increasing rupture elongation as 

molecular weight was increased. Shinozaki E133 found 

similar results with his studies on PVC. 

- 

Schenck C143 investigated the mechanical behavior of 

zero percent endcapped thermoplastic polyimidesulfone at 

four temperature levels. His results provide a set of data 

to be used in conjunction with the present investigation. 

9 
CHAPTER 2 

ANALYTICAL CONSIDERATIONS 

A convenient way of relating the mechanical properties 

of a viscoelastic material to its molecular weight is 

through viscosity. The viacoeity coefficient in a 

viscoelastic model represents a dashpot within the schematic 

of the model, signifying some degree of fluid-like 

behavior. The Chase-Goldsmith model ( which is a modified 

form of the three paraneter solid model ) incorporates 

viscosity into material characterization in this manner (see 

figure 1). A eliding element prevents the dashpot from 

affecting the stress strain behavior until after an initial 

linear elastic response is seen. 

Figure 1. The Viscoelaetic Chase-Goldsmith Model 

Used to Describe Adhesive Behavior 

LO 
A similar behavior was observed during our experiments with 

Thermoplastic Polyinideeulfone adheaive. This behavior was 

also similar to that of another polyimide adhesive ( LARC-3) 

which was characterized previouely at Clarkson. Sancaktar 

and Padgilwar sucessfully used the Chase-Goldsmith model to 

characterize the constant strain rate behavior of LARC-3 

adhesive C153. 

In shear, the constitutive equations for the 

Chase-Goldsmith model are: 

where 

G 

(T + G1y) 0 T 

s G +G1 y 0 

and 

G 

(4) 

The constants appearing in the Chase-Goldsmith 

model, T, T , y , y , Go, GI. and q , represent the shear 

strees, yield stress, shear strain, shear strain at the 

Y Y 

yield stress, initial elastic shear modulus, the 

11 
viscoelastic shear modulus, and the viscoeity coefficient, 

respectively. 

To obtain the constant strain rate relation, the above 

constitutive equation is solved according to the condition 

f = R = constant (6) 

to result in 

where 

and 

0 Q tG 

;;R 

(9) 

The material constants,including the viscosity 

coefficient, n, appearing in equation (7) can be evaluated 

from constant strain rate data. The coefficient of viscosity 

is a function of a number of environmental parameters such 

as temperature and rate, and process variables such as the 

polymeric molecular weight. Nielsen 1161 suggests the use of 

equation 

Log rl = log K + n log M (10) 

12 
where n is equal to 3.4 

for relating the viscosity coefficient to the molecular 

weight M. This relation was derived empirically by Fox and 

Flory (173 based on viscosity-temperature experiments with 

polyisobutylene fractions. The constant K of equation (10) 

is dependent on the solvent and temperature, and n was found 

to be equal to 3.4 for polyisobutylene samples with number 

average molecular weights above 17,000. Nielsen suggests 

that equation (10) is valid for polymeric materials having 

molecular weights above the critical molecular weight. The 

critical molecular weight is defined as the value above 

which the polymer chains become long enough for chain 

entanglements to form C93. Apparently sufficient 

information on the chemical structure of a material is 

necessary t3 evaluate the constant K. 

Comparison between the constant strain rate data of 

Thermoplastic Polyimidesulfone and the behavior predicted 

through application of the Chase-Goldsmith model are 

presented in Chapter 4. A comparison between the 

viscosities uaed to fit the Chase-Goldsmith model and those 

predicted by equation (10) can also be found in Chapter 4. 

It should be pointed out that equation (10) does not 

account for any variation in adhesion at the 

adhesive-adherend interface. Increasing viscosity tends to 

hinder wetting and diffusion at the interface during bond 

13 
I. 

formation C183. This can result in inferior bonds at higher 

molecular weights. Observations of the fracture surfaces of 

one and three percent endcapped Thermoplastic 

Polyimidesulfone specimens reveal the presence of such 

behavior. 

When the effects of rate on stress strain behavior are 

not strong, the use of a nonlinear elastic-plastic model may 

be adequate. One such model is the Raaberg-Oagood model 

which is based on Hencky’s stress strain relationa 1193. 

This theory assumes the plastic strains to be functions of 

the current state of stress and independent of the loading 

history. In one dimension the Raaberg-Osgood relation has 

the form 

T m 

G 

~=-+DT (11) 

where y is the shear strain,T is the shear stress, G is 

the elastic shear modulus, and D and m are material 

constante. The constant m denotes the plastic flow, end 

consequently the work-hardening characteristics. This model 

can be modified to account for rate effects. For example, 

Brinson and Renieri used a rate dependent version of the 

Ramberg-Osgood model to characterize Metlbond 1113 and 

1113-2 thermosetting (epoxy) adhesives in the bulk form 

C193. For the present investigation, observations revealed 

that separation of the linear and nonlinear terms of 

14 
equation (11) into the form 

T 

Y'ES (T(T 1 

Y 

m y'D T , (T<T 1 

Y 

(12) 

(13) 

would provide a more adequate representation of the observed 

stress strain behavior. 

Schenck used a nonlinear equation of the form 

b 

T = ay 

where 

(14) 

't = the shear stress 

y = the shear strain 

and 

a and b are material constants 

to characterize the constant strain rate behavior of the 

zero percent endcapped form of Thermoplastic 

Polyimidesulfone C143. Apparently equation (14) is not 

separated into linear and nonlinear segments as in equations . 

(12) and (13). However, results of the current investigation 

revealed that a more accurate representation could be 

derived for Thermoplaetic Polyimidesulfone adhesives of 

different molecular weights using equations (12) and (13). 

15 
The rate dependence of the rupture stresses under 

constant strain rate conditione can be expressed with the 

semiempirical equation proposed for metals by Ludwik in the 

form 

0 whereT is the ultimate shear stress, y is the initial ult 

01 elastic shear strain rate, and TI TI', and y are material 

constants C203. Brinson C213 used this equation to 

successfully describe the rate dependence of rupture 

stresses for polymeric materials in the bulk tensile mode. 

The same form of equation (15) can also be used to 

describe the variation of shear rupture stress ( 8 ) and 

strains ( 0 1 with initial elastic shear strain rates in the 

bulk and bonded forms. These relations may be written as 

where additional material constants are defined 

accordingly. 

Delayed failure of an adhesive can be described with 

the use of Crochet's creep-rupture equation (equation 21). 

Crochet's equation is based on Naghdi's definition of a 

yield surface as a function of stress <aij >* plastic strain 

16 
<E p), work hardening due to path history (x ),and the time 

history of loading deformation ( Kij ). The functional 

relationship 

iJ ij 

P ij$ Xijs K = 0 ij f = f(aij9 E (18) 

was simplified 1203 with the use of a Mises yield criterion 

to give 

2 -k =O f=%S ij ’ij 

k = k(X,K) 

(19) 

(20) 

Crochet neglected the work-hardening term, on the basis of 

the theory of elastic perfectly plastic solids, and assumed 

K to be only a function of X . Based on Crozhet’s 

assertions that K ia a monotonically decreasing function of 

X , and that K has a lower bound for larger values of 

X , Brinson (223 mathematically interpreted the functional 

behavior of K as 

Y(t) = A’ + B’ exp (-C’x> (21) 

in equation (21) Y(t> is the time dependent maximum stress 

and A’, B’, and C’ are material parameters. Crochet’s 

specific functional form for X was 

17 
V 

ij ij (22) 

where E VandE E refer to viscoelastic and elastic strains 

respectively. Sancaktar (233 interpreted equation (21) to 

apply to pure shear in the form 

ij ij 

with 

V E = xs 92 - y12 (24) 

E refer to the "engineering" shear 12 and y y12 

where 

strains, and A, B, and C are material constants. Since 

delayed failure occurs only in elements loaded up in to the 

viscoelastic region, an expression for X can be obtained 

by subtracting the elastic shear strain from the proposed 

equation. 

When the constant strain rate strass strain behavior of 

an adhesive is determined to be approximately nonlinear 

elastic, and the creep behavior is not characterized with a 

viscoelastic model ,then the simplest model creep relation 

can be utilized for the evaluation of X . This procedure, 

however, should be applied only as a first approximation for 

practical purposes. Following the procedure for the 

evaluation of X in conjunction with the Maxwell model 

creep relation (equation 26) results in a solution to 

18 
‘. ’ ‘. ’, . 

Crochet’s creep-rupture equation (equation 23) of the form 

(25) t 

17 

T = A + B exp[-CT -1 

T 

0 

y = y [t + $1 (26) 

A similar delayed failure equation results from the 

solution of equation (23) with the Chase-Goldsmith model. 

The resulting equation has the form 

G1 +- Y(t) (1 - exp [- - t11 

17 a0 

G 

G1 (exp[- - t] - I)]) Y(t) 

G n 

+- (27) 

0 

In equation (27) the exponential terms contain viscosity 

coefficients as viscoelastic models are used. In this case 

the viscosity affects the creep rate of the material. If 

the material constant C of equation (27) were taken to be a 

constant for all degrees of endcapping, increasing the 

viscosity would predict a longer time to failure for a given 

stress level. 

The asymptotic values of equations (25) and (27) 

represent the maximum safe creep stress levels below which 

19 
delayed failures are not expected to occur. Mathematical 

analysis reveals that equation (27) predicts a much faster 

approach to its asymptotic stress level in comparison to 

equation (25). These asymptotic values can be expressed a& 

a function of temperature to represent experimental data 

empirically. Such application of Crochet’s equation to the 

elavated temperature creep data is Justified by the 

applicability of the time-temperature superposition 

principle to viscoelastic creep relations. 

20 
CHAPTER 3 

EXPERIMENTAL PROCEDURES 

Materials and SDecimen Features 

The preparation and bonding of Thermoplastic 

Polyiafdesulfone has been described in detail by St. Clair 

C1,23. A polyinide acid solution is formed by mixiug 0.0569 

lbs (25.8 gna) of 3,3’,4,4’-Benzophenonetetracarboxylic 

dianhydride (BTDA) in a solution made of 0.0439 lbs (19.9 

gms) of 3,3’-diaminodiphenylsulfone (3,3’-DDS> and 0.5701 

lbs (258.6 gms) of bis (2-nethoxyethyl) ether. Endcapping 

is accomplished through the addition of a phthalic anhydride 

in concentrations varying from zero to three percent. The 

acid solution is then brushed onto 112-E-A100 glass carrier 

cloth. The cloth acts as a control for bondline thickness 

and as a carrier for the adhesive. The scrim is then heated 

to 212 OF (100 OC) for one half hour, 302 OF (150 OC) for 

one half hour, and to 392 OF(200 OC) for one half hour. 

This heating process removes the solvent and the water from 

the adhesive. 

Titanium (6-AL,4-V) strips are grit blasted with 120 

mesh aluminum oxide and treated with Pasa Jell 107. A 

primer coating of the polyimide acid is brushed onto 

titanium and heated to 212 OF (100 *C> for one hour and 392 

OF (200 OC) for one hour. Lap shear specimens are then 

formed by sandwiching the scrim between the titanium 

21 
adherends using a 0.5 inch (1.27 cm) overlap. The specimens 

are heated at 41 oF/min (5 oC/min) to 392 OF (200 OC); a 200 

psi (1.38 MPa) clamping force is applied at this point. The 

specimens ere then heated further to 635 OF (335 OC), held 

for 5 minutes, then allowed to cool to 302 OF (150 OC) 

before removing the bonding clamps. 

All test specimens were prepared at NASA - Langley 

Research Center, and are in accordance with ASTM D1002 

specifications. 

Molecular weights were determined by membrane osmometry 

on a Knauer osmometer at Arro Laboratories.111. Information 

on molecular weight distribution and number average 

molecular weight values could not be obtained because of the 

cloudiness of the material. The molecular weight values 

reported by NASA are shown in table (1). 

Table 1. Molecular Weight Values of Zero, One, and Three 

Percent Endcapped Thermoplastic Polyimidesulfone 

Adheai ve. 

Percent End-Cap 

0.0 

1 .o 

3.0 

The bondline 

Molecular Weight Carnu) 

19700 

18800 

17400 

thicknessee were measured using a 

22 
microscope with 40X magnification. Twenty five measurements 

were taken along each outside edge of the bonds. The 

bondline length, adherend width, and adherend thicknesses 

were measured with a Kanon dial caliper. Specimen 

dimensions are listed in tab198 ( A-1 and A-2 1 in Appendix 

<A). 

The dinensione for each specimen were needed to 

calculate the adhesive shear stresses and strains. The 

shear stress was assumed to be uniform and was calculated by 

dividing the applied load by the overlap area. The shear 

strain was calculated by the bondline elongation divided by 

the average bondline thickness. 

23 
Testinu Methods 

The mechanical testing of the single lap 

Polyimideeulfone samples was performed at Clarkson 

University. NASA - Langley Research Center supplied twenty 

of each one and three percent endcapped specimens. Testing 

on zero percent endcapp’d specimens was previosly performed 

at Clarkaon University by Schenck C143. All testing was 

done using a Model 1331 Instron servohydraulic materials 

testing machine. 

\ 

Each single lap specimen had notches milled 0.99 inches 

(2.51 cm) apart on the overlap edge of the bond. The 

notches were 0.01 inches (-0254 cm) deep by 0.01 inches 

(0.0254 ca) wide and were used for the placement of a high 

precision extensometer. The output signal from the 

extensometer was amplified through the Instron’s internal 

amplification system. The signals from the extensometer and 

the load cell were then recorded on a three channel strip 

chart recorder. The extensometer was calibrated using a 

high magnification extensometer calibrator. 

Output from the extensometer was used to calculate the 

strain in the adhesive joint. Deformations in the adherends 

were subtracted from the total deformation values to obtain 

adhesive deformations. The adherend deformations were 

calculated using equation: 

[F(D - 011 

[E (AT)W 1 Adherend Deformation = (28) 

24 
where 

F= load applied to the adherends 

D= the distance between the milled notches in the 

adherende of the aingle lap specimens, 0.99 inches 

(2.51 CR) 

O= overlap length 

E= Young’s Modulus for titanium, 15,000,000 psi 

(103,421 MPa) 

AT =adherend thicknese, 0.05 inches (0.127 cm) 

W- udherend width over the overlap area. 

Creep and constant strain rate tests were performed at 

temperatures of 70 OF (21 OC), 250 OF (121 OC), and 350 OF 

(177 OC). High temperature testing was performed with an 

environmental chamber mounted on the testing frame of the 

Instron. Room temperature constant strain rate tests were 

conducted with crosshead rates of 0.001, 0.1, and 10.0 

inches/minute (0.00254, 0.254, 25.4 cm/min). High 

temperature constant strain rate tests were ran at a 

crosshead rate of 5.0 incheslminute (12.7 cm/minute). Three 

to five specimens of each percent endcapping were tested in 

creep at each temperature level. The creep tests had an 

initial crosshead rate of 0.3 inches/minute (0.762 cm/ 

minute), and were performed et stress levels above the 

elastic limit stress found from the constant strain rate 

tests. 

25 
r 

Figure 2. 

0 

J 

OVERLAP 

AREA 

0 

SPRING ATTACHMENT TO 

SPECIMEN TO SECURE CLIPON GAGE 

I 

CLIP-ON 1 GAGE 

NOTCH MACHINED ONTO 

SINGLE LAP SPECIMEN 

\ 

Clip-on Gage Attachment to the Single Lap 

Specimen for Adhesive Deformation Yeasurement. 

26 

v- ... 
CHAPTER 4 

RESULTS AND DISCUSSION 

The constant strain rate behavior of zero, one, and 

three percent endcapped Thermoplastic Polyimidesulfone 

adhesive at room temperature is shown in figure 3. The 

experiments displayed are all at comparable initial elastic 

strain rates. Experiments at 100 and 10,000 times slower 

rates did not exhibit strong rate dependence based on 

ultimate stresses and strains. The observed regions of 

linear elastic behavior followed by regions of visco-plastic 

behavior suggest that the Chase-Goldsmith model may be used 

to describe the constant strain rate behavior of 

Thermoplastic Polyimidesulfone. Behavior predicted by the 

Chase-Goldsmith model is also shown in figure 3. The model 

provides a good fit to the data at all stress levels 

observed. The coefficients (Go, GI, T , y , and 0 1 used to 

fit the model at room temperature are given in table 2. 

0 

Y 

The elastic shear moduls, G, for each endcapped form of 

Thermoplastic Polyimidesulfone was calculated from the 

experimental data by averaging the slope of the shear stress 

strain curve between data points in the initial linear 

elastic region. G was observed to increaee as molecular 

weight increased (figure 4). The values of Go and C1 used to 

fit the Chase-Goldsmith model at room temperature followed 

27 
5.0 

4.0 

3.0 

2.0 

1.0 

0 

0 

& SPECfMEN FAILURE 

EXPERIMENTAL DATA 

OZERO PERCENT END-CAP 

AONE PERCENT END-CAP 

OTHBES PERCENT END-CAP 

T=70zF 

=21 c 

, 

\ 

2147-2 

- CHASE - GOLDSMITH 

MODEL 

20 40 

SHEAR STRAIN (%> 

60 

35 

30 

25 

10 

5 

Figure 3. Constant Strain Rate Stress Strain Behavior 

of Thermoplastic Polyimidesulfone Adhesive 

and Comparison with Theory. 

28 
Table 2 

Coefficients u8ed to fit the Chaee-Goldamith rodel to the constant 

strain rate behavior of zero, one, and three percent endcapped 

Thermoplastic Poliride8ulfone adhesive at 70 OF (21 OC) (see figure 3) 

T 

End-Cap KSI (HPa) KSI (HPa) KSI (HPa) %/sec KSI-sec (HPa-sec) 

rl ? Percent Go 61 Y 

0 17.1 (117.8) 4.3 (29.6) 3.0 (20.7) 24.6 4.1 (28.2) 

1 13.8 (95.1) 1.6 (11.0) 2.9 (20.0) 38.5 2.1 (14.5) 

3 7.0 (48.2) 1.2 (8.3) 2.3 (15.8) 125.0 0.8 (5.5) 

Chase-Goldsmith rodel: 

29 
25.0 

20.0 

15.0 

10.0 

5.0 

T = 70°F (21OC) 

-30 7.173 

-29 7.173 

- G(ksi) = (2.74xlO )M 

G(MPa) = (1.89~10 )M 

2147-2 

- 

180 

160 

140 

120 

100 

80 

60 

40 

20 

17,000 18,000 19,000 20,000 

MOLECULAR WEIGHT (amu) 

Figure 4. Variation of Elastic Shear Modulus with 

Molecular Weight for Thermoplastic 

Polimidesulfone Adhesive. 

30 
the same trend. 

The viscoeities found in fitting the Chase-Goldsmith 

model are plotted in figure 5. Equation (10) predicts a much 

slower increase in viscosity be molecular weight increases. 

The discrepancys between the two values may partially be 

attributed to the fact that the adhesive ueed for 

experimentation was in the bonded form rather than in the 

bulk fora. 

The constant strain rate behaviors of zero, one, and 

three percent endcapped Thermoplastic Polyinidesulfone at 

250 OF (121 OC) and 350 OF (177 OC) are shown in figures 6 

and 7 along with theoretical predictions by the 

Chase-Goldsmith model. The regions of linear elastic 

behavior followed by visco-plastic deformation are evident. 

It can be observed that the zero percent endcapped specimens 

reached higher rupture strains and lower rupture stresses in 

comparison to the one percent endcapped samples. This can 

be attributed to Schenck's 1141 observation of increasing 

interfacial failure with increasing temperature for the zero 

percent endcapped specimens. This phenomena causes the 

coefficients of the Chase-Goldsmith model to change. thus 

making it impractical to compare viscosities, shear moduli, 

etc. for the different molecular weights at elevated 

temperatures. The coefficients used to fit the 

Chase-Goldsmith model at 250 OF (121 OC) and 350 OF (177 OC) 

are listed in tables 3 and 4. 

31 
EXP ER I?IENTAL 

35 

30 

25 n 

U 

0) 

v1 

20 

15 

10 

5 

17,000 18,000 19,000 20,000 

NOLECULAR \;EIGHT (am> 

Figure 5. Variation of Viscosity with Molecular IJeight 

for Thermoplastic Polyimidesulfone Adhesive. 

I 

(3 

a 

E: 

W 

.. 

n 

> 

32 
5.0 

4.0 

3.0 

2.0 

1.0 

0 

0 

SPECIMEN FAILURE 1 EXPERIMENTAL DATA 

0 ZERO PERCENT END-CAP 

A ONE PERCENT END-CAP 

a THREE PERCENT END-CAP 

2142-3. 

CHASE - GOLDSMITH 

NODEL 

- 

T - 250'F 

* 121OC 

20 40 60 

SHEAR STRAIN (Z) 

35 

30 

25 

10 

Eigure 6. Constant Strain Rate Stress Strain Behavior 

of Thermoplastic Polyimidesulfone Adhesive 

and Comparison with Theory. 

33 
Table 3 

Coefficients used to fit the Chase-Goldsmith rode1 to the constant 

strain rate behavior of zero, one, and three percent endcapped 

Thermoplastic Polirideaulfone adhesive at 250 OF (121 OC) (see figure 6) 

rl T ? Percent Go Gl Y 

End-Cap KSI (HPa) KSI (HPa) KSI (HPa) %/sec KSI-sec (HPa-sec) 

0 24.2 (166.7) 2.1 (14.5) 2.3 (15.8) 59.1 1.2 (8.3) 

1 20.7 (142.6) 2.2 (15.2) 3.0 (20.7) 19.4 4.0 (27.6) 

3 16.5 (113.7) 2.1 (14.5) 2.7 (18.6) 20.2 2.9 (20.0) 

34 
5.0 

4.0 

3.0 

2.0 

1.0 

0 

0 

SPECIMEN FAILURE T = 350EF 

’ EXPERIMENTAL DATA = 177 C 

0 ZERO PERCENT END-CAP 

A ONE PERCENT END-CAP 

n THREE PERCENT END-CAP 

2140-4 

- CHASE - GOLDSMTH 

MODEL 

20 40 60 

SHEAR STRAIN (%) 

35 

30 

25 

n 

Q 

f 

I 

W 

VI 

m 20 

3 

X 

m 

15 

10 

5 

Figure 7. Constant Strain Rate Stress Strain Behavior 

of Thermoplastic PolyimidesuIfone Adhesive 

and Comparison with Theory. 

35 
Table 4 

Coefficients used to fit the Chaee-Goldarith rodel to the constant 

strain rate behavior of zero, one, and three percent endcapped 

Therroplastic PoliRidO6ulfOne adhesive at 350 OF (177 OC) (see figure 7) 

Percent Go G1 Y ? rl 

T 

End-Cap KSI (HPa) KSI (NPa) KSI (HPa) %/sec KSI-sec (NPa-sec) 

0 13.3 (91.6) 1.0 (6.9) 2.0 (13.8) 86.0 1.5 (10.3) 

1 12.38 (84.1) 1.7 (11.7) 2.9 (20.0) 26.1 1.8 (12.4) 

3 9.6 (66.1) 1.3 (9.0) 1.9 (13.1) 28.2 2.9 (20.0) 

36 
The ultimate shear stresses under constant strain rate 

conditions (figure 8) show the effects of the high 

temperature interfacial failures on the zero percent 

endcapped specimens. At room temperature the ultimate shear 

stresses increase with increasing molecular weight in a 

linear fashion. It should be noted that this data represent 

the maximum ultimate shear stress values even though other 

(lower) values were also obtained at different test rates. 

The ultimate shear stress values reported by NASA for one 

half and two percent endcapped Thermoplastic 

Polyinidesulfonc specimens at room temperature also agree 

with the linear interpolation of increasing ultimate shear 

stress with increasing molecular weight. A similar linear 

behavior between the ultimate shear stress and molecular 

weight is observed at elevated temperatures. This data, 

however, represent results obtained using only a single 

crosshead rate. 

Figures 9 and 10, respectively, show the variation of 

the ultimate shear stresses and maximum shear strains with 

the initial elastic shear strain rate for zero, one, and 

three percent endcapping. Data depicted in both figures is 

irregular, implying weak rate dependence which can be 

concealed by the effects of the fracture process and other 

experimental variations. 

The constant strain rate behavior of Thermoplastic 

Polyimidesulfone can also be predicted using a nonlinear 

37 
0 

N ul 

m 

0 el 

I I I I I I 

0 

7 k 

9 :! X z 

GI 

rl 

I 

N 

U 

4 

N 

U 

I 

0 

U 

rl 

N 

L) 

s 

bc 

al 

3 

l-l 

38 
0 

m N 

In 2 I I I I I I 

'\O \ 

I I 

e4 U 

ca e 

--I rl 

rl m 

I1 I1 

4 m h h 

h 

U 

n 

nN 

rlrl 

NN 

ww 

4 

I 

N 

U 

rl 

N 

4. 

U 

I 

0 

U 

ri 

N 

4 

I b 

U 

rl 

N 

2 

5 

L 

I\ 

4 

C 

0 

C 

a3 

rl 

.I 

0 

0 

0 

L 

b 

rl 
I I I I I 1 I I 1 I I 

0 

OQ 

U 

I 

d 

cu 

%a 

m 

I 

(5\ 

m 

a 

N 

I CQ 

5 

4 

cv 

\ 

\ \ 

h 

W 

r \ I 

\i 

\ 

oc 

0 

4 

n 

h 

cc) 

0 

+ 

W 

h 

Ln 

CJ 

I1 

n 

c: 

c 

5: 

c 

U 

39 
0 

hl 

I 

rU 

rl 

h( 

w 

tc 

0 

n 

9 

U 

LA 

+ 

0 

hl 

P 

(7 

!I 

i 

v 

I 

0 

U 

I h 

U 

4 

N 

U 

I 

m 

CL) 

4 

N 

a 

00 

3 C 

a G 

N 

0 

-4 

4 

0 

4 

c 

0 

i 

R 

I 

2 

1 4 

N 
elastic approach. In figures 11, 12, and 13 it can be 

observed that by assuming linear elastic behavior up to the 

elastic limit atress. and by using a nonlinear power law 

equation paet this limit, the constant strain rate behavior 

of zero, one, and three percent endcapped specimens can be 

predicted adequately at all temperatures tested. The 

coefficient, 11, of the nonlinear relation used (equation 13) 

is determined from the experimental data. Since m is the 

reciprocal of the coefficient d commonly used in a power law 

of the form 

i 29.) d 

~=ry 

it is a relative measure of the amount of work-hardening 

that the adhesive can withstand before failure occurs. In 

other words, low values of m indicate soft, ductile 

materials, and high values of m indicate hard brittle 

materials. Examination of data indicates that room 

temperature m decreased by 22.5 % between zero and one 

percent endcapping, and by 8.0% between one and three 

percent endcapping. As can be seen in figure 11, toughness 

increases considerably when the molecular weight is 

increased from 17400 amu to 19700 amu. This trend was not 

as pronounced at elavated temperatures. 

The creep results for three percent endcapped 

Thermoplastic Polyimidesulfone at 70 OF (21 OC), 250 OF (121 

41 
n 

.d 

rn 

24 

m 

m 

s 

2 

k 

m 

W 

5.0 

4.0 

3.0 

2.0 

SPECIMEN FAILURE T = 70:F 

EXPERIMENTAL DATA f 21 c - 

O ZERO PERCENT END-CAP 

A ONE PERCENT END-CAP 

TKREE PERCENT END-CAP 

1.0 

0 

0 20 40 60 

SHEAR STRAIN (%) 

35 

30 

25 

20 

15 

10 

5 

0 

Figure 11. Constant Strain Rate Stress Strain Behavior 

of Thermoplastic Polyimidesulfone Adhesive 

and Comparison with Theory. 

42 
5.0 

4.0 

3.0 

2.0 

1.0 

0 

0 

SPECIMEN FAILURE 5. EXPERIMENTAL DATA 

O ZERO PERCENT END-CAP 

A ONE PERCENT END-CAP 

THREE PERCENT END-CAP 

'- 3 

T = 250°F 

= 121OC 

NONL INEXR ELXST IC 

THEORY 

- 

20 40 60 

SHEAR STRAIN (Z) 

Figure 12. Constant Strain Rate Stress Strain Behavior 

of Thermoplastic Polyimidesulfone Adhesive 

and Comparison with Theory. 

35 

30 

25 

20 

1< .La 

10 

5 

0 

43 
5.0 

4.0 

3.0 

2.0 

1.0 

0 

0 

T = 350°F 

EXPERIMENTAL DATA = 177OC 

1 SPECIMEN FAILURE 

0 ZERO PERCENT END-CAP 

A ONE PERCENT END-CAP 

c] THREE PERCENT END-CAP 

2140-4 

- NONLINEAR ELASTIC 

THEORY 

20 40 60 

SHEAR STRAIN (%> 

35 

30 

25 

20 

15 

10 

5 

0 

Figure 13. Constant Strain Rate Stress Strain 3ehavior 

of Thermoplastic Polyimidesulfone Adhesive and 

Comparison with Theory. 

44 
Table 5 

Coefficients used to fit the nonlinear relation 

(equations 12 and 13) to the constant etrain rate atreea atrain behavior 

of Therroplaatic Polyirideeulfone Adhesive (see figures 11, 12, and 13) 

Percent 

End-Cap 

0 

1 

3 

0 

1 

3 

0 

1 

3 

T 

OF (OC) 

70 (21) 

70 (21) 

70 (21) 

250 (121) 

250 (121) 

250 (121) 

350 (177) 

350 (177) 

350 (177) 

f 

Y 

kai (HPa) 

3.0 (20.7) 

2.9 (20.0) 

2.3 (15.8) 

2.3 (15.8) 

3.0 (20.7) 

2.7 (18.6) 

2.0 (13.8) 

2.9 (20.0) 

1.9 (13.1) 

Nonlinear Elaetic Relation: 

G 

ksi (MPa) 

18.1 (124.7) 

14.3 (98.5) 

7.3 (50.3) 

24.2 (166.7) 

20.7 (142.6) 

16.5 (113.7) 

13.3 (91.6) 

12.2 (84.1) 

9.6 (66.1) 

D 

ksi (HPa) 

0.010 (0.069) 

0.027 (0.186) 

0.073 (0.503) 

0.016 (0.110) 

0.015 (0.103) 

0.018 (0.124) 

0.031 (0.214) 

0.034 (0.234) 

0.062 (0.427) 

1 

2.44 

1.89 

1.74 

2.16 

2.01 

2.18 

2.09 

1.80 

1.72 

y=D 

45 
OC), and 350 OF (177 OC) are shown in figures 14, 15, and 16 

respectively. The adhesive’s resistance to creep in the 

primary and secondary regions is characteristic of linear 

thermoplastics. Contrary to expectations, observations do 

not show a strong trend of increasing creep strains with 

increasing temperature. This may be caused by the level of 

molecular weight. The chain entanglement may be sufficiently 

low to force the strain to be more dependent upon the 

polymer network than temperature and secondary factors. The 

decrease in safe creep stress levels at 350 OF (177 OC) may 

be caused by decreased viscosity due to increased 

temperature, allowing chain slippage to occur at lower 

stress levels. 

Figures 17, 18, and 19 show the creep results of one 

percent endcapped specimens at 70 OF (21 OC), 250 OF (121 

oC),and 350 OF (177 OC), respectively. Creep strains 

increase as the temperature is increased. This is expected 

due to decreasing viscosity and increasing molecular motion. 

Decreases in the levels of safe creep stresses are also 

observed with increasing temperatures. 

Figure 20 illustrates creep strains at comprable stress 

levels applied for 50 minutes for zero, one, and three 

percent endcapped specimens. The strains plotted are all 

from secondary creep regions. The creep strains decrease as 

molecular weight is increased. This behavior is expected 

since increases in molecular weight result in increasing 

46 
viscosity and decreasing chain mobility. 

Figures 21, 22, and 23 show comparison of creep-rupture 

data with Crochet's equation baaed on the Maxwell (equation 

25) and Chase-Goldsmith (equation 27) models for three, one, 

and zero percent endcapped Theromoplastic Polyimidesulfone 

adhesives, respectively. Apparently, in most situations, 

Crochet's equation based on the Chase-Goldsmith model 

provides a better fit to the experimental data in comparison 

to the solution with the Maxwell model. The Chase-Goldsmith 

form of the equation predicts a much faster approach to the 

asymptotic safe creep stress levels. The appropriate 

constants used in fitting Crochet's equation based on the 

Maxwell and Chase-Goldsmith models are shown in tables 6 and 

7, respectively. 

The maximum safe stress levels under creep conditions 

at room temperature, 250 OF (121 OC), and 350 OF (177 OC> 

are shown in figure 24. At room temperature there is an 

increase of approximately 17% as the molecular weight 

increases by 7.4 %. The safe creep stress level remains 

relatively constant when the molecular weight is increased 

by another 4.6%. The increase is expected due to the 

increasing chain entanglement and increasing degree of 

intermolecular forces. The nearly constant values of maximum 

ultimate shear stress, between one and zero percent 

endcapping, on the other hand, may be caused by decreasing 

adhesion as the viscosity increases with molecular weight 

47 
c 

CD 

0 

0 

r) 

0 

a0 

h v 

a 

0 

U 

0 

54 

3 

3 

N 

I 73 .- e. 

, 
0 

U 

0 

0 

rl 

0 

00 

0 

9 

3 

U 

0 

N 

3 

3 

N 

49 
m 

I 

9 

0 

0 

rl 

0 

ca 

0 

a 

0 

U 

c 

c\I 

J 

a 

C 

0 u4 

rl 

3 

m 

Q) 

U 

50 
0 

2 

0 

co 

C 

\o 

0 

U 

0 

cu 

0 

a 

?. 

C Jd 

.r) C’J 

51 
, f '. 

m 

I 

m 

c 

0 

0, 

0 

a) 

c 

9 

0 

V 

3 

fl 

0 

52 
\-+ 

0 

0 

l-l 

0 

eo 

0 

,a 

0 

U 

0 

N 

0 

53 
0 4 0 

hl U e4 

I I I 

Q\ 

so 

h 

c7 

I 

N 

I I 

9 

u i c 

N N 

OQ 0 

I I I I 1 I I I I 

54 
Table 6 

Constants used in fitting Crochet's equation baaed on the 

Haxwell rodel for zero, one, and three percent endcapped 

Theraoplaatic Polylaidasulfone adheeive 

(8- figure8 21, 22, and 23) 

Percent Teap A B C 

End-Cap OF (OC) psi (HPa) pei (HPa) pui-ain (!!Pa-rin) 

0 70 (21) 3267 (22.5) 1054 (7.3) 46.7~10'6 (0.32~10'6) 

1 70 (21) 3239 (22.3) 579 (4.0) 7.3~10'6 (0.05~10'6) 

3 70 (21) 2769 (19.1) 467 (3.2) 10.8~10'6 (0.07~10'6) 

0 250 (121) 2765 (19.0) 685 (4.7) 27.4~10'6 (0.19~10'6) 

0 350 (177) 2270 (15.6) 1069 (7.4) 157.7~10'6 (1.09~10'6) 

1 350 (177) 2889 (19.9) 477 (3.3) 126.7~10'6 (0.87~10'6) 

3 350 (177) 2434 (16.8) 430 (3.0) 25.8~10'6 (0.18~10'6) 

Crochet's Equation Based on the Haxwell Model: 

t 

n 

T = A + B exp[-C.r -1 
h 

0 

aJ 

(0 

1 h 

4. 

N 

VI 

9 

h 

W 

n 

9 

m 

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d 

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ln 

d 

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d 

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56 
beyond a point where optimum Interfacial bond formation is 

no longer possible. The safe creep stress levels are also 

observed to Increase with molecular weight at higher 

temperatures for the one and three percent endcapped 

specimens. The significant decreases In the safe creep 

stress levels at zero percent endcap value under elavated 

temperatures Is attributed to the high temperature 

interfacial failures as discussed previously. 

Figure 25 Illustrates the variation of safe creep 

stress with temperature for zero, one, and three percent 

endcapped Thermoplastic Polyimidesulfone adhesive. 

Increasing safe creep stress levels are &served with 

increased molecular weight (excluding the zero percent 

end-cap values at elevted temperatures). The safe creep 

stress versus temperature behavior of the one and three 

percent endcapped specinens is very similar. This suggests 

that the degree of decrease in viscosity and increase in 

chain mobility with increasing temperature is relatively 

constant for the molecular weight values examined. 

57 
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61 
ENVIRONMENTAL TERMPERATURE (OC) 

3.4 

3.2 

3.0 

2.8 

2-6 

2.4 

2.2 

0 50 100 150 200 

0 ZERO PERCENT END-CAP 

A ONE PERCENT END-CAP 

THREE PERCENT END-CAP 

LOW VALUES ATTRIdUTED TO 

INTERFACIAL FAILL'RE RESULTISG 

FROM HIGH TENPERATURE 

24 

.. 

22 

20 

18 

16 

14 

12 

0 100 200 300 400 

EXVIROXMENTAL TENPERATURE (OF) 

Figure 25. Variation of ?laxinun (Safe) creep Stress 

with Environmental Temperature for Different 

End-Cap Forms of Thermoplastic Polyimidesulfone 

Adhesive. 

62 
CHAPTER 5 

CONCLUSIONS 

The effects of molecular weight on the shear creep and 

constant strain rate behavior of Thermoplastic 

Polyimidesulfone adhesive was studied in the bonded form. 

For this purpose, constant strain rate, creep, and 

creep-rupture experiments were performed on zero C141, one, 

and three percent endcapped specimens of the adhesive. 

Mathematical relations were proposed to represent the 

constant strain rate and delayed failure 

viscoelastic-plastic behavior of Thermoplastic 

Polyimidesulfone adhesive. Based on the mathematical 

relations, used in con3unctim-i with experiaental results, 

several conclusions can be derived: 

It is possible to describe the constant strain rate 

shear stress strain behavior of zero, one, and three percent 

endcapped Thermoplastic Polyimidesulfone adhesive using 

viscoelastic Chase-Goldsmith model or nonlinear elastic 

relations; 

Ludwik’s equation does not reveal changing rate 

dependence (pertaining to ultimate shear stresses and 

strains) with changing molecular weights: 

63 
Crochet’s equation based on the Chase-Goldsmith model 

provides a better description of the delayed failure data in 

comparison to Crochet’s equation based on the Maxwell model; 

Room temperature constant strain rate results show 

increases in viscosity with increasing molecular weight <it 

should be noted, however, that these viscosity coefficient 

values are calculated parameters obtained on the basis of 

the Chase-Goldsmith model and they do not represent values 

measured using standard viscosity measurement techniques): 

Nonlinear relations reveal increasing strain-hardening 

capacity with increasing molecular weight; 

At all temperature levels of experimentation, ultimate 

shear stress levels and safe creep stress levels increased 

with increasing molecular weight. 

Reductions in molecular weight revealed increased 

adhesion (less interfacial failure), but also reduced the 

cohesive strength of the model adhesive. In general, the 

results of this investigation imply that reductions in the 

molecular weight of Thermoplastic Polyimidesulfone adversely 

affect the mechanical properties studied. Reductions in 

molecular weight may increase wetting and diffusion during 

the bonding process, but these effects are insufficient to 

override the advantages of increased intermolecular and 

intramolecular forces created at higher molecular weights. 

64 
REFERENCES 

1. St. Clair, T.L. and Yaaaki, D.A., **A Thermoplastic 

Polyiaideeulfone," NASA Langley Research Center, 

Technical Memorandum No. 84574, Nov. 1982. 

2. Singh, J.J., St. Clair,'T.L. and Walker, S., ** A Study 

of the Effects of End-Cap Molecular Species on 

Enviromental Characteristics of Polyimidesulfones," 

NASA Langley Research Center, Technical Memorandum No. 

86283, Aug. 1984. 

3. Siggia, S. and Hark, H. F. , Polymer Science and 

Materials, Ed. Tobolsky, A.V. and Mark, H.F., John 

Wiley and Sons, Inc., New York (19711, pp 56-57. 

4. Alfrey, T. Jr., Mechanical Behavior of High Polymere, 

Interscience Publishers, Inc., New York (19481, pp. 

493-498. 

5. Tsuji, T., Masuoka, M. and Nakao, K., "Superposition of 

Peel Rate, Temperature and Molecular Weight for T Peel 

Strength of Polyisobutylene," Adhesion and Adsorption 

of Polymers, Plenum Publishing Corp.,1980. 

6. Zhurkov, S. N., '' Kinectic Concept of the Strength of 

Solids," Internl. J. of Frac. Mech., 1 (41, 331-323 

(1965). 

7. Mark, H. F., Polymer Science and P!aterials, Ed. 

Tobolsky, A.V. and Mark, H.F., John Wiley and Sons, 

Inc., New York (19711, p 236. 

8. Margolies, A. F. , " Effect of Molecular Weight on 

Properties of HDPE," SOC. Plas. Eng. J., 27, 44, June 

1971 . 9. Fellers, J. F. and Kee, B. F. , "Crazing Studies 

of Polystyrene. I. A New Phenomenological Observation," 

Journal of Applied Polymer Science, John Wiley and 

Sons, Inc., 18, pp 2355-2365, (19641. 

10. Knauss, W.G., "The Mechanics of Polymer Fracture," App. 

Mechanics Reviews, Jan. 1973. 

11. Gardner, R. J. and Martin, J. R., Humid Aging of 

Plastics: Effects of Molecular Weight on Mechanical 

Properties and Fracture Morphology of Polycarbonate," 

J. of App. P. Sei., 24, 1269-1280, 1979. 

12. Vlachopoulos, J., Had~is, N., and Hamielec, A. E., 

"Influence of Molecular Weight on the Tensile 

Properties of Nearly Monodisperse Polystyrenes," 

65 
Polymer, 19, 115, 1977. 

13. Shinozaki, D. M. , Woo, K. , Vlachopoulos, J. and 

Hanielec, A., "Influence of Processing Parameters and 

Molecular Weight on the Mechanical Properties of PVC," 

J. App. P. SCi., 21, 3345-3354, 1977. 

14. Schenck , S. C. and Sancaktar , E. , " Material 

Characterization of Structural Adhesives in the Lap 

Shear Mode,*' NASA Langley Research Center, NASA 

Contractor Report No. 172237, Sept. 1983. 

15. Sancaktar, E. and Padgilwar, S., I' The Rate and Time 

Dependent Materia 1 Characterization of LARC-3 

Structural Adhesive," Clarkson College, Potsdam, New 

York, Report No. HIE-069, April 1981. 

16. Nielsen, L. E. , Mechanical Properties of Polymers, 

Reinhold Publishing Corp., New York (19621,pp. 56-57. 

17. Fox, T.G. and Flory, P.J., "Further Studies on the Melt 

Viscosity of Polyisobutylene," The Journal of Physical 

and Colloid Chemiatry, Ed. Lind, S.C., 55, 221-234. 

1951 . 

18. Wu, S. , Polymer Interface and Adhesion, Marcel Dekker, 

Inc., New York (19821, p. 403. 

19. Renieri, M. P., Herakovich, C. T. and Brinson, H. F., 

"Rate and Time Dependent Behavior of Structural 

Adhesives," Virginia Polytechnic Institute, Report No. 

VPI-E-76-7, April 1976. 

20. Sancaktar, E. and Schenck, S.C., "Temperature Dependent 

Creep Rupture of Structural Adhesives," Proceedings of 

the Southeastern XIIth Conference on Theoretical and 

Applied Mechanics, Vol. 2, 148-153 (1984). 

21. Brinson, H. F., Renieri, M. P. and Herakovich, C. T., 

** Rate and Time Dependent Failure of Structural 

Adhesives," Fracture Mechanics of Composites, ASTM STP 

593, 1975,pp. 177-199. 

22. Brinson, H. F. , Deformation and Fracture of High 

Polymers, (H.H. Kaush et al. eds.1 Plenum Press, New 

York (1974). 

23. Sancaktar. E. and Padgilwar, S., '' The Effects of 

Inherent Flaes on the Time and Rate Dependent Failure 

of Adhesively Bonded Joints," Transactions of the ASME 

Journal of Mechanical Design, Vol. 104, 643-650, July 

1982. 

66 
APPENDIX A 

TABLES OF SINGLE LAP SPECIMEN DIMENSIONS 

67 
Table A-1 

Dimension6 for Single Lap Specimens Bonded with 

One Percent Endcapped Thermoplastic Polyimidesulfone 

Specimen 

No. 

2138-2 

2138-3 

2138-4 

2139-1 

2139-3 

2139-4 

2140-1 

2140-2 

2140-4 

2141-1 

2141-2 

2141-3 

2142-1 

2142-2 

2142-3 

2142-4 

2149-1 

2149-2 

2149-3 

2149-4 

Bondline Thicknees 

Inche8 (CHI 

0.00695 (0.01765) 

0.00644 (0.01636) 

0.00764 (0.01941) 

0.00731 (0.01857) 

0.00744 (0.01890) 

0.00908 (0.02306) 

0.00699 (0.01775) 

0.00610 (0.01549) 

0.00849 (0.02156) 

0.00976 (0.02479) 

0.00813 (0.02065) 

0.00819 (0.02080) 

0.00759 (0.02019) 

0.00746 (0.01895) 

0.00615 (0.01562) 

0.00674 (0.01712 1 

0.00653 (0.01659) 

0.00468 (0.01189) 

0.00547 (0.01389) 

0.00673 (0.01709) 

Overlap Length Overlap Width 

Inchea (CHI Inchea (CHI 

0.52705 (1.33871) 0.98767 (2.50868) 

0.52500 (1.33350) 0.98170 (2.49352) 

0.51920 (1.31877) 0.98213 (2.49461) 

0.54515 (1.38468) 0.99560 (2.52882) 

0.53305 (1.35395) 0.98827 (2.51021) 

0.52490 (1.33325) 0.99707 (2.53256) 

0.51805 (1.31585) 0.99413 (2.52509) 

0.52900 (1.34366) 0.99287 (2.52189) 

0.53995 (1.37147) 0.99877 (2.53688) 

0.52485 (1.33312) 0.99982 (2.53954) 

0.52660 (1.33756) 0.99140 (2.51816) 

0.52455 (1.33236) 0.99350 (2.52349) 

0.53445 (1.35750) 0.99633 (2.53068) 

0.53100 (1.34874) 0.98545 (2.50304) 

0.52745 (1.33972) 0.98600 (2.50444) 

0.52080 (1.32283) 0.98930 (2.51282) 

0.52411 (1.33124) 0.99143 (2.51823) 

0.52170 (1.32512) 0.99077 (2.51656) 

0.51945 (1.31940) 0.98466 (2.50105) 

0.51661 (1.31219) 0.99993 (2.53982) 

68 
Table A-2 

Dinendona for Single Lap Specinens Bonded with 

Three Percent Endcapped Thermoplastic Polyiridesulfone 

Specinen 

No. 

2143-2 

2143-3 

2143-4 

2144-1 

2144-3 

2144-4 

2145- 1 

2145-2 

2145-4 

2146-1 

2146-2 

2146-3 

2147-1 

2147-2 

2147-3 

2147-4 

2148-1 

2148-2 

2148-3 

2148-4 

Bondlfne Thickness 

Inche8 (CHI 

0.00321 (0.00815) 

0.00263 (0.00668) 

0.00452 (0.01148) 

0.00444 (0.01128) 

0.00352 (0.00894) 

0.00495 (0.01257) 

0.00450 (0.01143) 

0.00399 (0.01013) 

0.00435 (0.01105) 

0.00313 (0.00795) 

0.00322 (0.00818) 

0.00318 (0.00808) 

0.00317 (0.00805) 

0.00316 (0.00803) 

0.00309 (0.00785) 

0.00369 (0.00937) 

0.00308 (0.00782) 

0.00472 (0.01199) 

0.00547 (0.01389) 

0.00593 (0.01506) 

Overlap Length Overlap Width 

Inches (CM) Inchee (CHI 

0.54105 (1.37427) 0.98950 (2.51333) 

0.54420 (1.38227) 0.98717 (2.50741) 

0.54605 (1.38697) 0.98906 (2.51221) 

0.51605 (1.31077) 0.99463 (2.52636) 

0.52860 (1,34264) 0.99184 (2.51927) 

0.53295 (1.35369) 0.99297 (2.52214) 

0.52405 (1.33109) 0.99660 (2.53136) 

0.52174 (1.32522) 0.99213 (2.52001) 

0.52313 (1.32875) 0.99210 (2.51993) 

0.49911 (1.26774) 0.96990 (2.46355) 

0.51815 (1.31610) 0.97176 (2.46827 

0.53435 (1.35725) 0.96817 (2.45915 

0.51960 (1.31978) 0.99120 (2.51765 

0.52600 (1.33604) 0.98587 (2.50411 

0.53150 (1.35001) 0.99157 (2.51859) 

0.53541 (1.35994) 0.99501 (2.52733) 

0.51842 (1.31679) 0.99198 (2.51963) 

0.51973 (1.32011) 0.99372 (2.52405) 

0.51942 (1.31933) 0.98461 (2.50091) 

0.52601 (1.33607) 0.99104 (2.51724)