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Adhesive bonding of thick steel adherends for marine structures
Marine Structures 12 (1999) 405}423
Adhesive bonding of thick steel adherends for marine structures S.A. Hashim* Glasgow Marine Technology Centre, Department of Mechanical Engineering, University of Glasgow, James Watt Building, Glasgow G12 8QQ, UK Received 28 November 1998; received in revised form 19 August 1999; accepted 25 August 1999
Abstract The paper is intended to aid designers when considering structural adhesives as an alternative joining method for load bearing structures. Results from published and unpublished Glasgow Marine Technology Centre (GMTC) research are presented and reviewed in relation to important aspects of bonded steel joints. The main aspects include the understanding of adhesive properties and their limitations, design and behaviour of structural joints and the use of stress analysis. Research methodologies and results in relation to a hot-curing epoxy adhesive for steel grillage panels and attachments, with emphasis on experimental techniques, are highlighted. The topics are presented and discussed in separate but inter-related sections. Key conclusions to these discussions are given. Crown copyright ( 1999 Published by Elsevier Science Ltd. All rights reserved. Keywords: Marine structures; Epoxy adhesives; Joint performance; Grillage panel; Design tools
1. Introduction and background Modern epoxy adhesives o!er engineering designers #exibility to achieve economical and technical advantages for o!shore and ship construction, especially in grillage connections between plates and sti!eners [1}3]. Fig. 1 shows various concepts for connections between steel sti!eners and steel or polymeric composite plates but this paper is concerned with bonding of structural steel adherends (bonded substrates). Adhesives also o!er considerable potential in the repair and minor attachments on ships and o!shore structures [4}6]. Adhesive bonding, just as welding and fastening,
* Tel.: 0141-339-0969; fax: 0141-330-4015. 0951-8339/99/$ - see front matter Crown copyright ( 1999 Published by Elsevier Science Ltd. All rights reserved. PII: S 0 9 5 1 - 8 3 3 9 ( 9 9 ) 0 0 0 2 9 - 5
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Fig. 1. Possible bonded sti!ened structural panel designs.
has speci"c design and materials requirements and a successful application depends upon their understanding. These include optimum use of adhesive properties, appropriate bonding processes, an understanding of the behaviour of bonded joints and design tools to evaluate joint strength. The major advantages of adhesive bonding for steel applications are: f the elimination of thermal distortion associated with welding [7]; f improved fatigue strength, especially in low and long-term loading conditions [8], due to low stress concentrations; f the ability to create e$cient complex joints, such as in sandwich structures [9]; f a reduction of pitting corrosion due to the absence of weld defects and the additional bene"t of the adhesive acting as a sealant within a joint, thus minimising crevice corrosion. The main disadvantages of the use of adhesives are: f some surface pre-conditioning is required to obtain strong and durable joints [1,10]; f it is di$cult to combine the properties of maximum impact resistance and elevated temperature resistance in a single adhesive [1,11,12]; f long-term durability in wet/humid conditions is uncertain due to a shortage of design data at present [10,13];
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f a high-temperature sensitivity when compared with metals; f load bearing joints require new design skills and may require optimum modi"ed standard sections. Before designing a bonded joint, it is essential to obtain technical data which specify performance, environments, service life, dimensions, weight, materials, testing programme, cost limitations, standardisation and appearance. Failure to address one of these parameters could result in pursuing the wrong development. The purpose of this paper is to underline important properties and limitations of structural adhesive materials, the structural behaviour of bonded joints and the relevant design tools. The sections below address these elements with reference to the joining of thick steel adherends. 2. Adhesive properties Adhesive selection depends on design requirements and this may be assisted by an experimental programme based on test specimens to various BSI and ASTM standards [14,15]. The choice of suitable standard test specimens, to study adhesives for thick adherends applications, should take into account the need to produce cleavage and shear modes of failure in sti! joints. Fig. 2 shows standard shear and cleavage test
Fig. 2. Standard test specimens.
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specimens to ASTM D3165 and BS5350, respectively. A range of structural adhesives were examined in recent studies [1,6,10] for bonding steel to steel, steel to GRP (glass reinforced plastics) and GRP to GRP. Table 1 gives mechanical and processing properties for three typical epoxy adhesives. Whilst these come from a single manufacturer, other manufacturers have their equivalents. The adhesives in Table 1 are toughened epoxy adhesives which have the ability to resist brittle failure [16,17]. The single-part hot-curing Araldite 2007 adhesive (AV199) has the best mechanical properties and is suitable for steel bonding. The other two are cold-curing two-part adhesives suitable for bonding both steel and polymeric composites. Araldite 420 (Redux 420A/B) o!ers good cleavage strength while Araldite 2004 (AV138M#V998) has good temperature resistance. In general, properties exhibited by hot curing adhesives are superior to cold-curing ones because of their higher glass transition temperatures, greater density of crosslinking and superior wetting capabilities [16]. Curing of a single-part epoxy adhesive is typically carried out at 1503C for 60 min, whilst some two-part epoxy adhesives may take days to be fully cured, at room temperature. The two-part products can also be warm cured, typically at 703C for up to 2 h. Adhesive manufacturers are usually able to specify the curing schedule to give optimum adhesion. Data from mechanical testing of bulk adhesives (produced by casting of adhesive) and bonded joint specimens can be used by designers to estimate structural joint strength. Most frequently available strength data are obtained from tensile lap shear specimens with an overlap of 5 mm to determine shear strength and shear modulus of elasticity of the adhesive [18]. Bulk adhesive specimens can be used to determine tensile and compression strength as well as Young's modulus and Poisson's ratio. Compression strength is greater than the tensile strength due to the signi"cance of the crazing phenomena in the latter. In compression, microscopic defects in the adhesive experience less mode I (opening mode) loading. Fig. 3 shows typical test properties of bulk adhesive and a lap shear joint for a single-part epoxy adhesive. Higher tensile strength can be obtained from an axial butt joint where boundary conditions and
Table 1 General properties for three typical structural epoxy adhesives Adhesive trade name
Araldite 2007
Araldite 420
Araldite 2004
Number of parts Cure temp. (3C)/time (min)
One 180/20 120/180 1.5 110 4 0.35 65 48 19
Two 20/2 d 70/120 '0.5 70 2 0.35 35 25 17
Two 20/1 d 70/60 '0.5 150 4 0.35 30 34 8
Maximum gap "lling (mm) Max. service temp. (3C) Young's modulus (GPa) Poisson's ratio Bulk tensile strength (MPa) Av. shear strength (MPa) Av. cleavage strength (MPa)
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Fig. 3. Test results from bulk and joint specimens.
moulding quality are di!erent from that for bulk specimens. A butt joint with a 0.1 mm thick adhesive layer gave a tensile strength 50% higher than that from the bulk [1,10]. Micro defects in bulk adhesive could be the main reason for this di!erence. The bonding process for steel construction would typically require seven operations [19]. These are surface roughening, degreasing, marking, application of adhesive, positioning of clamps, curing and removal of clamps. Silane primers can also be applied to surfaces to provide chemical and physical protection mechanisms to inhibit corrosion and promote adhesion [10]. It is important that rules concerning safety precautions for working personnel are observed with reference to the COSHH regulations, particularly for skin protection and ventilation. Heat curing can be carried out by the use of conventional electrical resistance heating blankets or tapes [19]. Fig. 4 shows a laboratory prototype system [1] developed for heat curing of adhesive joints for 1.2]1.2 m sti!ened panels. This system could be developed further for a semi-automated production line in a shipyard environment. The system uses conventional low-voltage heating equipment (with control) and magnetic clamps. Other existing steel fabrication/production equipment for cleaning, painting and moulding processes can also be adapted and utilised in the bonding processes.
3. Durability in wet environments The main mechanisms [12,20}22] that contribute to strength reduction in bonded steel joints in wet environments are; interfacial attack to displace adhesive from the adherend, degradation of adhesive strength due to plastisisation and corrosion of adherends. It is believed that the "rst mechanism is the most in#uential. Factors
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Fig. 4. Layout of a steel panel during bonding.
associated with these mechanisms have been discussed widely [10,13,23,24]. These factors include (i) the level of humidity or water exposure and operating temperature, (ii) conditions of the bonded surface including roughness and whether primers have been used, (iii) the formulation of the epoxy adhesive and its curing conditions, (iv) corrosion protection of the adherends such as with external primers or electrical methods, (v) dimensions of the bonded area and (vi) the level of service loading on the joint. A large number of accelerated durability test methodologies [10,13] have been carried out on bonded steel joints for short periods of exposure, normally measured in weeks, to assess durability in wet/marine environments under various conditions. Most methods can give a good indication of the sensitivity of factors in#uencing joint
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strength. The major drawback with these tests is obtaining a relationship between laboratory test specimens/conditions and full-size joints subjected to natural service conditions. As part of the current study, a marine exposure site test at the University Marine Biological Station Millport, in the Lower Clyde Estuary, was used in a way similar to MoD (Navy) methodology [5]. Experiments were carried out to determine the residual strength in an adhesive when exposed to a natural marine environment over a period of "ve years [1]. In these experiments, 14 thick adherend lap shear specimens (Fig. 2) were bonded using standard bonding processes. Following curing, the specimens were coated externally with an epoxy paint to reduce corrosion. The specimens were then assembled into a chain in series using shackles and exposed to splash zone conditions, while under tensile loading representing 10% of the initial failure load. Samples were recovered at annual intervals over "ve years and tested in the wet condition. The results from these experiments are shown in Fig. 5, which illustrates an essentially linear strength reduction with time, approaching 50% after 5 yr. Visual examination of the fractured surfaces indicated no visible corrosion at the interface between the adhesive and the steel surface. Corrosion was, however, apparent around the edges of the small joint (25]15 mm). Samples tested after 4 yr su!ered considerable reduction of adherend thickness due to corrosion. This may have contributed to the strength reduction of the joints due to a reduction in adherend sti!ness. There are very few long-term published data on bonded steel joints that are available for comparison with the current results. However, results from experiments with bonded unstressed steel sti!eners, with hot-curing epoxy adhesives, carried out by Martin [24], indicated a similar reduction in tensile strength after 18 months of cycles of intermittent immersion in seawater under laboratory conditions. Other results [5] from exposure to a marine environment were of limited value due to severe
Fig. 5. Resistance of bonded joints to wet marine environments.
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corrosion of the steel lap shear joints, after 2 yr. In both of the above cases the specimens were not coated with a corrosion resisting primer. The exposure trial reported here may also, to a certain extent, be regarded as an accelerated test because such severe conditions are unlikely to be experienced by a practical application in marine structures. Using no adhesion primers and exposing the joints to intermittent immersion while under stress represents one of the most severe service conditions [13,24]. Bonded joints in ship and o!shore construction would experience much lower continuous stresses and be regularly painted to seal joints from wet conditions. Furthermore, structural joints would have a larger bonded area in comparison with the small standard lap shear joints tested here. Since moisture ingress is di!usion controlled [13], absolute size of a joint a!ects the degree of moisture ingress and time for total saturation.
4. Resistance to elevated temperatures The second main limitation of adhesively bonded joints is their relatively high sensitivity to temperature compared with that of structural metals. Bonded structures may experience high operating temperatures, be subject to high solar gain in a hot climate or be subject to accidental "re conditions. The resistance to elevated temperatures will depend, mainly, on the glass transition temperature of the adhesive. High-temperature commercial adhesives are now available and these include modi"ed epoxies, bismaleimide and cyanate-based adhesives. They have been used in many high-temperature applications [12,18,25,26] mainly in the aircraft industry. Their service temperatures are as high as 300 and 1503C for short- and long-term exposures, respectively. Their main drawback, however, is that the initial cleavage/peel strength at room temperature, is considerably lower than that for toughened structural adhesives. In addition they are relatively new and not widely evaluated. The thermal performance of a typical structural epoxy adhesive is demonstrated by Araldite 2007, which has a glass transition temperature of 1203C [18]. Lap shear specimens (Fig. 2) were tested over the temperature range from 15 to 2003C. Prior to loading, thermal equilibrium was achieved by enclosing the joint assembly within the test oven for 30 min at the required temperature. Fig. 6 shows test results that demonstrate the reduction in strength, with increasing temperature, especially as the lower glass transition temperature is approached. About 25% of the room temperature strength (48 MPa) is lost by 803C and 70% at 1203C. Beyond 1603C only marginal strength remains up to the char temperature of approximately 2503C. Comparison of these results with similar tests [26] using a high-temperature modi"ed epoxy adhesive shows that, while the modi"ed adhesive has lower shear strength (15}20 MPa) at room temperature, it su!ers only a 40% strength reduction at 1303C with no further decrease at 1803C. Continuous exposure to elevated temperature could have an even more dramatic e!ect on the strength of adhesive joints and should be limited to low-to-moderate loading conditions. For Araldite 2007 at 1003C the maximum stress should be less than 14% of the maximum room temperature strength to avoid creep rupture [1].
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Fig. 6. Resistance of bonded joints to elevated temperatures.
Even though joint strength is relatively low at elevated temperatures there are two factors which need to be taken into account. Firstly, the reduction in the modulus of elasticity of the adhesive leads to a reduction of peak stresses in bonded joints. Secondly, practical structural adhesive connections will have a large bond area and hence high inherent strength redundancy. Furthermore, resistance to elevated temperature is not just a function of the adhesive property but also joint design. Protective measures such as insulating a bonded joint from the source of heat and maintaining joint under compression and shear loading can be very e!ective to resist the e!ect of elevated temperature [16]. Another measure is to use a limited number of fasteners in critical locations in bonded structures [27]. It may be also feasible to complement adhesive bonding with limited welding [3,28], but the extent of the adhesive damage need to be investigated. It is worth noting that the thermal aspects of adhesive performance did not adversely a!ect their take-up in the aerospace industries.
5. Bending behaviour of bonded panels Lightly-loaded sti!ened panels such as minor bulkheads will mainly experience lateral loading, resulting in bending shear stress within the adhesive. Large shear strains in the adhesive result in increased stresses and de#ection in the bonded steel adherends. This in turn in#uences the scantling requirements in grillage structures. Sizes of, and spaces between, sti!eners and plate thickness are important considerations in optimising the fabrication cost of panels for ships and o!shore construction [29]. Therefore, it is essential to understand bending behaviour in a bonded panel in order to control structural design and fabrication cost. For this purpose, large-scale bonded panels were statically tested under four-point bending. The details of production and testing of these panels are reported elsewhere [1,7]. Fig. 7 shows some experimental details of the test panels. Results of load vs.
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Fig. 7. Four}point bending test.
Fig. 8. Load}de#ection curves for the panel in Fig. 7.
central de#ection curves are shown in Fig. 8, together with a theoretical curve based on classical beam theory. The experimental results show that the elastic central de#ection is higher than the theoretical value. The main reason for this is that the theoretical calculations assume that the panel/beam section is continuous i.e. they ignore the adhesive element between sti!ener and plate. Therefore, existing beam theories for solid beam sections require modi"cations to account for the materials discontinuity for the bonded beam section, i.e. steel/adhesive/steel. To study the above behaviour, simpli"ed bending models were idealised [1,30]. The details of these are shown in Fig. 9 which includes a bonded and a solid model to represent bonded and welded panels, respectively. Both models have similar dimensions except that the bonded is 5% thicker than the solid model due to the presence of the adhesive layer. Both have their neutral axis at the centre to simplify analysis. The two models were carefully produced and statically tested under three-point bending, while central de#ection and longitudinal bending strain at the lower surface were
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Fig. 9. Idealised bonded and welded beam models. Table 2 Dimensions and test results of beam models Model
l (mm)
h (mm)
I (mm4)
P (kN) x
De#ection (mm)
Strain (le)
Bonded Solid
30 30
10.5 10
2400 2100
15 15
0.065 0.040
1700 1400
measured. Table 2 shows test results including central de#ection and steel surface strain at a load of 15 kN. De#ection and strain for the bonded model were greater than that for the solid one. These results correspond to the behaviour pattern in the large panel (Fig. 8). However, unlike sti!ened panels, these models are not expected to fail in buckling of the sti!ener as a result of compression stresses. Thus, the behaviour is only representative within the elastic limit of the sections. In a theoretical estimate of the behaviour of bonded model it is assumed that there is no adhesion or friction between the two steel adherends of the beam, and that they behave like a leaf spring (stacked sheet strips). This assumption would ignore the shear #ow generated along the adhesive and predicts excessive stresses and de#ections [31]. Another, more realistic, approach is to consider analysis based on plane strain and static equilibrium of beam element using anisotropic plate theory [31]. This can be used for stress analysis of laminated composite materials including the consideration of the interface between laminae/plies. This can be applied to the bonded steel model in Fig. 9 and the relations between load and stresses and de#ection are given by [32] hM p "f , x 2 I
(1)
h2Q , q "f xz 8I
(2)
Pl3 , d "f 2 z 48EI
(3)
where f is the interface coe$cient, h the height of beam section, M the maximum bending moment, I the second moment of inertia, q the average interlaminar xz
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bending shear stress in adhesive, Q the shear force, P the force at centre of the beam, l the length of the beam, d the vertical de#ection of the beam, p the maximum z x bending stress of the beam, and E is Young's modulus of the adherends. For the bonded model, the calculated value is f + 1.2. f is a function of beam curvature which depends on the dimensions of section, boundary conditions and type of adhesive [33]. Further research into this area is required in order to produce typical values of f for bonded steel panels. However, a very recent study [37] on a bonded glass T-beam (plate and web) used a simple theory modi"ed from sandwich beam analysis and claims a good prediction for shear and bending stresses in the composite beam. The estimated f value from that study was approximately 1.1, considering the fact that the adherend is glass which has one-third the Young's modulus of the steel. The modi"ed theory seems very sensitive to adhesive shear modulus of elasticity and thickness. High stresses and de#ections in a bonded panel may imply that bonded structures should be slightly heavier than welded equivalents. This, however could be compensated for by the absence of welding residual stresses in the plating and the freedom to bond thinner plates with closer sti!eners without the technical and economical problems associated with controlling thermal distortion during welding of such geometries. Bonded sti!eners would also increase the e!ective breadth of panels due to the wide #ange attachment and absence of distortion.
6. Local behaviour of sti4ener/plate attachment Local stresses in adhesive joints are largely non-linear and typical local non-linear stress distributions in small bonded joints are shown in Fig. 10. This non-linearity arises from the loading direction and the great mismatch between the elastic modulus
Fig. 10. Local stresses in typical small joints [38].
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of the adherends and adhesive. The non-linearity makes detailed stress and strain analysis, for the purpose of fatigue prediction for example, di$cult without the use of suitable numerical methods. Because bonded panels are considered for light loading, local failure loads in sti!ener/plate joints may only arise as a result of extreme loading, such as in impact. To understand the cleavage behaviour of sti!ener/plate joints, a simpli"ed model was used (Fig. 11) to examine the cleavage failure under transverse loading on a sti!ener. A solid block, instead of a sti!ener, was used in this model to insure that cleavage failure takes place in the adhesive. Using a sti!ener such as rolled steel joist (RSJ) section would normally lead to plastic deformation of the web rather than adhesive joint failure [1]. The model was also used to study the in#uence of local plate sti!ness on cleavage stress. This was achieved by considering two lengths (200 and 300 mm) for the 8 mm plate, as a function of e!ective breadth in sti!ened panels. The model was tested statically until adhesive failure. Table 3 gives failure load results together with dimensions and contains adhesive cleavage stresses calculated from the numerical and analytical methods discussed in a later section. The above results may be related to the spacing of sti!eners in bonded panels (using the same plate thickness) that are subjected to transverse bending moment. A bonded panel with close sti!eners would produce lower cleavage stresses than one with widely
Fig. 11. Idealised bonded sti!ener/plate joint.
Table 3 Dimensions and results from analysis and test of sti!ener/plate models Plate length B (mm)
Joint length (mm)
Joint width (mm)
Failure load (kN)
Eq. (4) stress (MPa)
FE stress (MPa)
300 200
75 75
45 45
10.4 11.7
148 148
170 170
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spaced sti!eners. The closer the sti!eners the lower is the local de#ection of the plate. The de#ection largely controls the level of the cleavage stresses and hence the failure of the joint. Other factors which a!ect structural joint design are sti!ener foot shape, size and the detail of sti!ener ends. Experimental results based on RSJ sti!ener/plate joints are reported elsewhere [7]. In general, design of load bearing joints should consider the minimisation of cleavage stress and utilise shear and compression strength of the adhesive. These considerations are reported frequently in the literature [12,16] by producing `DOsa and `DON'Tsa guidance/recommendations for bonded joints. 7. Elastic stress analysis Elastic "nite element (FE) modelling has been widely used to assess strength in adhesive joints, with reasonable reliability. One approach to interpret adhesive failure criteria for bonded joints may be based on von Mises or principal stresses, occurring at a prescribed location from the edge of joint. Because of the mismatch of materials, FE models for adhesive joints exhibit points of theoretical singularity, especially at joint edges where stresses become in"nite. Errors in failure prediction of up to 20% might be expected with this approach [34]. Because high stress gradients occur in certain regions of bonded joints, these regions have to be modelled very accurately and economically. The adhesive layer is very thin (0.1}0.5 mm) compared with the thickness of the adherend (several mm), and to achieve reliable results, it is necessary to use several elements through the thickness of the adhesive. Owing to the resulting "ne mesh and the limitations of element aspect ratios, the number of elements and hence degrees of freedom in a joint is, typically, rather high. Fig. 12 shows a typical mesh and boundary conditions for the sti!ened joint (Fig. 11). Normally, the adhesive layer is modelled by at least "ve elements through the thickness with the highest mesh density nearer the tension edge of the joint. Many
Fig. 12. Finite element model for sti!ener/plate joint.
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packages are suitable for elastic FE analysis. Fig. 13 shows typical stress distributions along the 45 mm of the adhesive joint/line at 0.1mm above the plate interface with the adhesive. The maximum stress is expected to be at the plate surface, but would be di$cult to determine, due to the sudden change in material properties. The stress distribution is clearly non-linear and with high stresses concentrated in a small region of the joint. A similar pattern of stress distribution was also found in the thick adherend lap shear joint [1] (Fig. 2). This is shown in Fig. 14 where again the stresses along the adhesive line were taken at 0.1 mm above the lower adherend. When comparing the principal stresses at the ends of both joints (Figs. 13 and 14) good correlation between the peak principal stresses can be obtained. This indicates the suitability of the lap shear joint for the purpose of stress prediction for stresses arising from local loading on the sti!ened joint. An elastic analytical method to calculate local cleavage stress for sti!ened joints (L and T shapes) was developed by Crocombe et al. [35] from the classical lap shear analysis by Goland and Reissner [36]. This analysis considered the adhesive joint as
Fig. 13. Stress distribution along the adhesive line for a sti!ener/plate joint.
Fig. 14. Stress distribution along the adhesive line for a lap shear joint.
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Fig. 15. Analytical model for stress calculations.
adherend-adhesive sandwich subject to transverse shear forces and bending moments at the ends. Fig. 15 illustrates the basic model where the adherends are assumed to behave as cylindrically bent plates of di!erent #exural sti!ness. As a result of the di!erential in vertical displacement, the adhesive transmits cleavage stresses, with the maximum at the ends of joints. The expression for the maximum cleavage stress at the joint edge is p "Ak , (4) z a where p is the cleavage (normal) adhesive stress, and k the e!ective spring sti!ness of z a the adhesive. w E k " a a, a ¹ a where w is the adhesive joint width, E the adhesive Young's modulus, ¹ the a a a adhesive thickness, and A the constant function of forces and moments and adhesive thickness [1,35]. This design formula was applied to the models (joints) shown in Fig. 11. The calculated results are given in Table 3 which show realistic stress values in comparison with equivalent results from "nite element analysis. This formula, however, is sensitive to changes in adhesive thickness and does not account for variation in stresses through the thickness of the adhesive. It also, unlike the "nite analysis method, requires the designer to determine the forces and moments at the boundaries of the joint.
8. Conclusions and further work A number of materials and design requirements have been discussed, in relation to steel adherends and structural epoxy adhesives. The key conclusions and recommendations are; f Cleavage and shear strength values are very useful comparative data for selecting a structural adhesive. f Formulation of relevant small and large experiments, suitable for thick steel adherends, is a key to understanding and evaluating joint design and behaviour.
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f High strength structural adhesives are typically limited to operate at a temperature up to 1003C, which is acceptable for many marine applications. f Long-term exposure of loaded small lap shear joints to a severe wet environment causes a strength reduction of approximately 10% per year. Further work is required to establish whether the e!ect would be smaller in larger, practical joints and structures exposed to normal marine service conditions. f Beam theory for calculating stresses and de#ections in sections made up of two longitudinal elements bonded together may be simply modi"ed to account for the presence of the adhesive through the use of an interface coe$cient f. Further experimental work is necessary to obtain appropriate f values for bonded panels and materials. f FE analysis is suitable to evaluate local strength in sti!ened joints, especially when existing analytical techniques are correlated with results from lap shear joints. f Existing analytical methods, while less accurate than FE analysis, can be a very useful design tool to determine local stresses.
Acknowledgements This work was funded by Science and Engineering Research Council (Grant No. GR/C/7139.2/1G95) through the Marine Technology Directorate Ltd and the Ministry of Defence. Such support is gratefully acknowledged.
References [1] Hashim SA. Assessment of adhesive bonding for structural design with thick adherends. Ph.D. Thesis, University of Glasgow, Scotland, 1992. [2] Dodkins AR, Judd GE, Maddison A. Adhesively bonded aluminium superstructures. Proceedings of the International Conferance on Lightweight Materials in Naval Architecture. RINA, Paper No. 9, Southampton, 1996. p. 1}21. [3] Bochkarev VP, Glevitskaya TI. Adhesive bonded and welded joints in shipbuilding. Weld Prod 1970;17(4):43}7. [4] Oliver RA. Underwater applications of adhesives within o!shore and marine industries. Proceedings of the Seminars on Engineering Applications of Adhesives. London: Butterworths, 1988. p. 53}7. [5] Allan RC, Bird J, Clark JD. The use of adhesives in the repair of cracks in ship's structures. Proceedings of the First International Conference on Structural Adhesives in Engineering. Bristol: IMechE, 1986. p. 169}78. [6] Hashim SA, Winkle IE, Knox EM, Cowling MJ. Advantages of adhesive bonding in o!shore marine structural applications. In: Faulkner D, Cowling MJ, Incecik A, Das PK, editors. Proceedings of the Fifth International Symporium Integrity of O!shore Structures. Scotland: EMAS, 1993. p. 417}37. [7] Hashim SA, Winkle IE, Cowling MJ. A structural role for adhesive in shipbuilding. The Naval Architect, October 1990. p. 203}20. [8] Cowling MJ, Smith EM, Hashim SA, Winkle IE. Performance of adhesively bonded steel connections for marine structures. Proceedings of the International Conferance Evaluation of Materials Performance in Severe Environments. Japan: ISIJ 1989. p. 827}34.
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[9] Knox EM, Cowling MJ, Winkle IE. Adhesively bonded steel corrugated core sandwich construction for marine applications. Mar Struct 1998;11:185}204. [10] Knox EM. Marine applications for structural adhesives. Ph.D. Thesis, University of Glasgow, Scotland, 1996. [11] Perry HA. Adhesive bonding of reinforced plastics. London: McGraw-Hill, 1959. [12] Adams RA, Wake WC. Structural adhesive joints in engineering. London: Elsevier Applied Science, 1984. [13] Beevers A. A critical appraisal of durability test methods. Proceedings of the Fourth International Conference on Structural Adhesives in Engineering. Bristol: Inst of Materials, 1995. p. 137}44. [14] BS5350. Methods of tests of adhesives Part C1:Determination of cleavage strength of adhesive bonds. London: BSI, 1986. [15] ASTM D3165}79. Strength properties of adhesives in shear by tension loading of laminated assemblies. Philadelphia: ASTM, 1979. [16] Lees WA. Adhesive selection-Engineering applications of adhesives. London: RAPRA, 1988. [17] Bishop JA. The chemistry and properties of a new generation of toughened epoxy matrices. Int J Adhesion Adhesives 1992;12(3):178}84. [18] Jeandrau J. New concepts for designing structural adhesively bonded joints. Proceedings of the Second. International Conference on Structural Adhesives in Engineering. Bristol: Butterworth, 1989. p. 149}56. [19] Shields J. Adhesives handbook. London: Butterworth, 1984. [20] Comyn J. Kinetics and mechanisms of environmental attack. In: Kinlock AJ, editor. Durability of structural adhesives. London: Applied Science Publishers, 1983. p. 85}131. [21] Gledhill R, Kinloch AJ. Environmental failure of structural adhesive joints. Int J Adhesion 1983;6:315}30. [22] Bowdich MR, Hiscock D, Moth AD. The role of substrate in the hydrolytic stability of adhesive joints. Int J Adhesion Adhesives 1992;12(3):164}70. [23] Fay PA, Maddison A. Durability of adhesively bonded steel under salt spray and hydrothermal stress conditions. Proceedings of the Second International Conference on Structural Adhesives in Engineering. Bristol: Butterworth, 1989. p. 112}21. [24] Martin DM. Bonded shear sti!eners for steel bridges. Ph.D. Thesis, University of Dundee, Scotland, 1985. [25] Millard EC. Use of structural adhesives in aircraft turbine engine nacelles. Proceedings of the International Conference on Adhesion. Nottingham: PRI, 1984. p. 341}46. [26] Dixon DG. Adhesive bonding for high temperature applications. Proceedings of the Fourth International Conference on Structural Adhesives in Engineering. London: The Inst Materials, 1995. p. 257}60. [27] Smith CS. Design of marine structures in composite materials. London: Elsevier Applied Science, 1990. [28] Mortimer AA. Applications for structural adhesives within James Howden and Company Limited. Company report, Scotland, 1988. [29] Winkle IE Baird D. Towards more e!ective structural design through synthesis and optimisation of relative fabrication costs. The Naval Architect, November, 1986. p. 313}36. [30] Hashim SA, Winkle IE, Cowling MJ. Stress analysis of bonded thick adherend beam in bending. Proceedings of the Third Internationa Conference on Structural adhesives in Engineering. Bristol: PRI, 1992. p. 231}36. [31] Evans JH. Ship Structural design concepts. London: Cornell Maritime Press, 1975. [32] Whitney JM. Structural analysis of laminated anisotropic plates. Basel: Technomic, 1987. [33] Beevers A, Kho ACP. The performance of adhesive bonded thin gauge sheet metal structures with particular reference to box section beam. In: Mittal KL, editor. Adhesive joints formation, characteristics and testing. New York: Plenum Press, 1982. [34] Richardson G, Crocombe AD, Smith PA. Failure prediction in adhesives joints by various techniques including the modelling of crack development. Proceedings of the Fourth International Conference on Structural Adhesives in Engineering. Bristol: The Inst of Materials, 1995. p. 45}50.
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[35] Crocombe AD. A general bonded joint peel analysis. Proceedings of the First International Conference on Structural Adhesives in Engineering. Bristol: IMechE, 1986. p. 1}13. [36] Goland M, Reissner E. The stresses in cemented joints. J Appl. Mech. 1944;66:17}27. [37] Pye A, Ledbetter S. The selection of an adhesive for construction of a glass-adhesive T-beam. Int J Adhesion Adhesives 1998;18:159}65. [38] Beevers A. The principles of adhesive bonding. Proceedings of the Seminar The Engineer's Guide to Industrial Adhesives. London: National Tribology Centre, 1995. p. 2}18.
Adhesive bonding of thick steel adherends for marine structures S.A. Hashim* Glasgow Marine Technology Centre, Department of Mechanical Engineering, University of Glasgow, James Watt Building, Glasgow G12 8QQ, UK Received 28 November 1998; received in revised form 19 August 1999; accepted 25 August 1999
Abstract The paper is intended to aid designers when considering structural adhesives as an alternative joining method for load bearing structures. Results from published and unpublished Glasgow Marine Technology Centre (GMTC) research are presented and reviewed in relation to important aspects of bonded steel joints. The main aspects include the understanding of adhesive properties and their limitations, design and behaviour of structural joints and the use of stress analysis. Research methodologies and results in relation to a hot-curing epoxy adhesive for steel grillage panels and attachments, with emphasis on experimental techniques, are highlighted. The topics are presented and discussed in separate but inter-related sections. Key conclusions to these discussions are given. Crown copyright ( 1999 Published by Elsevier Science Ltd. All rights reserved. Keywords: Marine structures; Epoxy adhesives; Joint performance; Grillage panel; Design tools
1. Introduction and background Modern epoxy adhesives o!er engineering designers #exibility to achieve economical and technical advantages for o!shore and ship construction, especially in grillage connections between plates and sti!eners [1}3]. Fig. 1 shows various concepts for connections between steel sti!eners and steel or polymeric composite plates but this paper is concerned with bonding of structural steel adherends (bonded substrates). Adhesives also o!er considerable potential in the repair and minor attachments on ships and o!shore structures [4}6]. Adhesive bonding, just as welding and fastening,
* Tel.: 0141-339-0969; fax: 0141-330-4015. 0951-8339/99/$ - see front matter Crown copyright ( 1999 Published by Elsevier Science Ltd. All rights reserved. PII: S 0 9 5 1 - 8 3 3 9 ( 9 9 ) 0 0 0 2 9 - 5
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Fig. 1. Possible bonded sti!ened structural panel designs.
has speci"c design and materials requirements and a successful application depends upon their understanding. These include optimum use of adhesive properties, appropriate bonding processes, an understanding of the behaviour of bonded joints and design tools to evaluate joint strength. The major advantages of adhesive bonding for steel applications are: f the elimination of thermal distortion associated with welding [7]; f improved fatigue strength, especially in low and long-term loading conditions [8], due to low stress concentrations; f the ability to create e$cient complex joints, such as in sandwich structures [9]; f a reduction of pitting corrosion due to the absence of weld defects and the additional bene"t of the adhesive acting as a sealant within a joint, thus minimising crevice corrosion. The main disadvantages of the use of adhesives are: f some surface pre-conditioning is required to obtain strong and durable joints [1,10]; f it is di$cult to combine the properties of maximum impact resistance and elevated temperature resistance in a single adhesive [1,11,12]; f long-term durability in wet/humid conditions is uncertain due to a shortage of design data at present [10,13];
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f a high-temperature sensitivity when compared with metals; f load bearing joints require new design skills and may require optimum modi"ed standard sections. Before designing a bonded joint, it is essential to obtain technical data which specify performance, environments, service life, dimensions, weight, materials, testing programme, cost limitations, standardisation and appearance. Failure to address one of these parameters could result in pursuing the wrong development. The purpose of this paper is to underline important properties and limitations of structural adhesive materials, the structural behaviour of bonded joints and the relevant design tools. The sections below address these elements with reference to the joining of thick steel adherends. 2. Adhesive properties Adhesive selection depends on design requirements and this may be assisted by an experimental programme based on test specimens to various BSI and ASTM standards [14,15]. The choice of suitable standard test specimens, to study adhesives for thick adherends applications, should take into account the need to produce cleavage and shear modes of failure in sti! joints. Fig. 2 shows standard shear and cleavage test
Fig. 2. Standard test specimens.
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specimens to ASTM D3165 and BS5350, respectively. A range of structural adhesives were examined in recent studies [1,6,10] for bonding steel to steel, steel to GRP (glass reinforced plastics) and GRP to GRP. Table 1 gives mechanical and processing properties for three typical epoxy adhesives. Whilst these come from a single manufacturer, other manufacturers have their equivalents. The adhesives in Table 1 are toughened epoxy adhesives which have the ability to resist brittle failure [16,17]. The single-part hot-curing Araldite 2007 adhesive (AV199) has the best mechanical properties and is suitable for steel bonding. The other two are cold-curing two-part adhesives suitable for bonding both steel and polymeric composites. Araldite 420 (Redux 420A/B) o!ers good cleavage strength while Araldite 2004 (AV138M#V998) has good temperature resistance. In general, properties exhibited by hot curing adhesives are superior to cold-curing ones because of their higher glass transition temperatures, greater density of crosslinking and superior wetting capabilities [16]. Curing of a single-part epoxy adhesive is typically carried out at 1503C for 60 min, whilst some two-part epoxy adhesives may take days to be fully cured, at room temperature. The two-part products can also be warm cured, typically at 703C for up to 2 h. Adhesive manufacturers are usually able to specify the curing schedule to give optimum adhesion. Data from mechanical testing of bulk adhesives (produced by casting of adhesive) and bonded joint specimens can be used by designers to estimate structural joint strength. Most frequently available strength data are obtained from tensile lap shear specimens with an overlap of 5 mm to determine shear strength and shear modulus of elasticity of the adhesive [18]. Bulk adhesive specimens can be used to determine tensile and compression strength as well as Young's modulus and Poisson's ratio. Compression strength is greater than the tensile strength due to the signi"cance of the crazing phenomena in the latter. In compression, microscopic defects in the adhesive experience less mode I (opening mode) loading. Fig. 3 shows typical test properties of bulk adhesive and a lap shear joint for a single-part epoxy adhesive. Higher tensile strength can be obtained from an axial butt joint where boundary conditions and
Table 1 General properties for three typical structural epoxy adhesives Adhesive trade name
Araldite 2007
Araldite 420
Araldite 2004
Number of parts Cure temp. (3C)/time (min)
One 180/20 120/180 1.5 110 4 0.35 65 48 19
Two 20/2 d 70/120 '0.5 70 2 0.35 35 25 17
Two 20/1 d 70/60 '0.5 150 4 0.35 30 34 8
Maximum gap "lling (mm) Max. service temp. (3C) Young's modulus (GPa) Poisson's ratio Bulk tensile strength (MPa) Av. shear strength (MPa) Av. cleavage strength (MPa)
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Fig. 3. Test results from bulk and joint specimens.
moulding quality are di!erent from that for bulk specimens. A butt joint with a 0.1 mm thick adhesive layer gave a tensile strength 50% higher than that from the bulk [1,10]. Micro defects in bulk adhesive could be the main reason for this di!erence. The bonding process for steel construction would typically require seven operations [19]. These are surface roughening, degreasing, marking, application of adhesive, positioning of clamps, curing and removal of clamps. Silane primers can also be applied to surfaces to provide chemical and physical protection mechanisms to inhibit corrosion and promote adhesion [10]. It is important that rules concerning safety precautions for working personnel are observed with reference to the COSHH regulations, particularly for skin protection and ventilation. Heat curing can be carried out by the use of conventional electrical resistance heating blankets or tapes [19]. Fig. 4 shows a laboratory prototype system [1] developed for heat curing of adhesive joints for 1.2]1.2 m sti!ened panels. This system could be developed further for a semi-automated production line in a shipyard environment. The system uses conventional low-voltage heating equipment (with control) and magnetic clamps. Other existing steel fabrication/production equipment for cleaning, painting and moulding processes can also be adapted and utilised in the bonding processes.
3. Durability in wet environments The main mechanisms [12,20}22] that contribute to strength reduction in bonded steel joints in wet environments are; interfacial attack to displace adhesive from the adherend, degradation of adhesive strength due to plastisisation and corrosion of adherends. It is believed that the "rst mechanism is the most in#uential. Factors
410
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Fig. 4. Layout of a steel panel during bonding.
associated with these mechanisms have been discussed widely [10,13,23,24]. These factors include (i) the level of humidity or water exposure and operating temperature, (ii) conditions of the bonded surface including roughness and whether primers have been used, (iii) the formulation of the epoxy adhesive and its curing conditions, (iv) corrosion protection of the adherends such as with external primers or electrical methods, (v) dimensions of the bonded area and (vi) the level of service loading on the joint. A large number of accelerated durability test methodologies [10,13] have been carried out on bonded steel joints for short periods of exposure, normally measured in weeks, to assess durability in wet/marine environments under various conditions. Most methods can give a good indication of the sensitivity of factors in#uencing joint
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strength. The major drawback with these tests is obtaining a relationship between laboratory test specimens/conditions and full-size joints subjected to natural service conditions. As part of the current study, a marine exposure site test at the University Marine Biological Station Millport, in the Lower Clyde Estuary, was used in a way similar to MoD (Navy) methodology [5]. Experiments were carried out to determine the residual strength in an adhesive when exposed to a natural marine environment over a period of "ve years [1]. In these experiments, 14 thick adherend lap shear specimens (Fig. 2) were bonded using standard bonding processes. Following curing, the specimens were coated externally with an epoxy paint to reduce corrosion. The specimens were then assembled into a chain in series using shackles and exposed to splash zone conditions, while under tensile loading representing 10% of the initial failure load. Samples were recovered at annual intervals over "ve years and tested in the wet condition. The results from these experiments are shown in Fig. 5, which illustrates an essentially linear strength reduction with time, approaching 50% after 5 yr. Visual examination of the fractured surfaces indicated no visible corrosion at the interface between the adhesive and the steel surface. Corrosion was, however, apparent around the edges of the small joint (25]15 mm). Samples tested after 4 yr su!ered considerable reduction of adherend thickness due to corrosion. This may have contributed to the strength reduction of the joints due to a reduction in adherend sti!ness. There are very few long-term published data on bonded steel joints that are available for comparison with the current results. However, results from experiments with bonded unstressed steel sti!eners, with hot-curing epoxy adhesives, carried out by Martin [24], indicated a similar reduction in tensile strength after 18 months of cycles of intermittent immersion in seawater under laboratory conditions. Other results [5] from exposure to a marine environment were of limited value due to severe
Fig. 5. Resistance of bonded joints to wet marine environments.
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corrosion of the steel lap shear joints, after 2 yr. In both of the above cases the specimens were not coated with a corrosion resisting primer. The exposure trial reported here may also, to a certain extent, be regarded as an accelerated test because such severe conditions are unlikely to be experienced by a practical application in marine structures. Using no adhesion primers and exposing the joints to intermittent immersion while under stress represents one of the most severe service conditions [13,24]. Bonded joints in ship and o!shore construction would experience much lower continuous stresses and be regularly painted to seal joints from wet conditions. Furthermore, structural joints would have a larger bonded area in comparison with the small standard lap shear joints tested here. Since moisture ingress is di!usion controlled [13], absolute size of a joint a!ects the degree of moisture ingress and time for total saturation.
4. Resistance to elevated temperatures The second main limitation of adhesively bonded joints is their relatively high sensitivity to temperature compared with that of structural metals. Bonded structures may experience high operating temperatures, be subject to high solar gain in a hot climate or be subject to accidental "re conditions. The resistance to elevated temperatures will depend, mainly, on the glass transition temperature of the adhesive. High-temperature commercial adhesives are now available and these include modi"ed epoxies, bismaleimide and cyanate-based adhesives. They have been used in many high-temperature applications [12,18,25,26] mainly in the aircraft industry. Their service temperatures are as high as 300 and 1503C for short- and long-term exposures, respectively. Their main drawback, however, is that the initial cleavage/peel strength at room temperature, is considerably lower than that for toughened structural adhesives. In addition they are relatively new and not widely evaluated. The thermal performance of a typical structural epoxy adhesive is demonstrated by Araldite 2007, which has a glass transition temperature of 1203C [18]. Lap shear specimens (Fig. 2) were tested over the temperature range from 15 to 2003C. Prior to loading, thermal equilibrium was achieved by enclosing the joint assembly within the test oven for 30 min at the required temperature. Fig. 6 shows test results that demonstrate the reduction in strength, with increasing temperature, especially as the lower glass transition temperature is approached. About 25% of the room temperature strength (48 MPa) is lost by 803C and 70% at 1203C. Beyond 1603C only marginal strength remains up to the char temperature of approximately 2503C. Comparison of these results with similar tests [26] using a high-temperature modi"ed epoxy adhesive shows that, while the modi"ed adhesive has lower shear strength (15}20 MPa) at room temperature, it su!ers only a 40% strength reduction at 1303C with no further decrease at 1803C. Continuous exposure to elevated temperature could have an even more dramatic e!ect on the strength of adhesive joints and should be limited to low-to-moderate loading conditions. For Araldite 2007 at 1003C the maximum stress should be less than 14% of the maximum room temperature strength to avoid creep rupture [1].
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Fig. 6. Resistance of bonded joints to elevated temperatures.
Even though joint strength is relatively low at elevated temperatures there are two factors which need to be taken into account. Firstly, the reduction in the modulus of elasticity of the adhesive leads to a reduction of peak stresses in bonded joints. Secondly, practical structural adhesive connections will have a large bond area and hence high inherent strength redundancy. Furthermore, resistance to elevated temperature is not just a function of the adhesive property but also joint design. Protective measures such as insulating a bonded joint from the source of heat and maintaining joint under compression and shear loading can be very e!ective to resist the e!ect of elevated temperature [16]. Another measure is to use a limited number of fasteners in critical locations in bonded structures [27]. It may be also feasible to complement adhesive bonding with limited welding [3,28], but the extent of the adhesive damage need to be investigated. It is worth noting that the thermal aspects of adhesive performance did not adversely a!ect their take-up in the aerospace industries.
5. Bending behaviour of bonded panels Lightly-loaded sti!ened panels such as minor bulkheads will mainly experience lateral loading, resulting in bending shear stress within the adhesive. Large shear strains in the adhesive result in increased stresses and de#ection in the bonded steel adherends. This in turn in#uences the scantling requirements in grillage structures. Sizes of, and spaces between, sti!eners and plate thickness are important considerations in optimising the fabrication cost of panels for ships and o!shore construction [29]. Therefore, it is essential to understand bending behaviour in a bonded panel in order to control structural design and fabrication cost. For this purpose, large-scale bonded panels were statically tested under four-point bending. The details of production and testing of these panels are reported elsewhere [1,7]. Fig. 7 shows some experimental details of the test panels. Results of load vs.
414
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Fig. 7. Four}point bending test.
Fig. 8. Load}de#ection curves for the panel in Fig. 7.
central de#ection curves are shown in Fig. 8, together with a theoretical curve based on classical beam theory. The experimental results show that the elastic central de#ection is higher than the theoretical value. The main reason for this is that the theoretical calculations assume that the panel/beam section is continuous i.e. they ignore the adhesive element between sti!ener and plate. Therefore, existing beam theories for solid beam sections require modi"cations to account for the materials discontinuity for the bonded beam section, i.e. steel/adhesive/steel. To study the above behaviour, simpli"ed bending models were idealised [1,30]. The details of these are shown in Fig. 9 which includes a bonded and a solid model to represent bonded and welded panels, respectively. Both models have similar dimensions except that the bonded is 5% thicker than the solid model due to the presence of the adhesive layer. Both have their neutral axis at the centre to simplify analysis. The two models were carefully produced and statically tested under three-point bending, while central de#ection and longitudinal bending strain at the lower surface were
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415
Fig. 9. Idealised bonded and welded beam models. Table 2 Dimensions and test results of beam models Model
l (mm)
h (mm)
I (mm4)
P (kN) x
De#ection (mm)
Strain (le)
Bonded Solid
30 30
10.5 10
2400 2100
15 15
0.065 0.040
1700 1400
measured. Table 2 shows test results including central de#ection and steel surface strain at a load of 15 kN. De#ection and strain for the bonded model were greater than that for the solid one. These results correspond to the behaviour pattern in the large panel (Fig. 8). However, unlike sti!ened panels, these models are not expected to fail in buckling of the sti!ener as a result of compression stresses. Thus, the behaviour is only representative within the elastic limit of the sections. In a theoretical estimate of the behaviour of bonded model it is assumed that there is no adhesion or friction between the two steel adherends of the beam, and that they behave like a leaf spring (stacked sheet strips). This assumption would ignore the shear #ow generated along the adhesive and predicts excessive stresses and de#ections [31]. Another, more realistic, approach is to consider analysis based on plane strain and static equilibrium of beam element using anisotropic plate theory [31]. This can be used for stress analysis of laminated composite materials including the consideration of the interface between laminae/plies. This can be applied to the bonded steel model in Fig. 9 and the relations between load and stresses and de#ection are given by [32] hM p "f , x 2 I
(1)
h2Q , q "f xz 8I
(2)
Pl3 , d "f 2 z 48EI
(3)
where f is the interface coe$cient, h the height of beam section, M the maximum bending moment, I the second moment of inertia, q the average interlaminar xz
416
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bending shear stress in adhesive, Q the shear force, P the force at centre of the beam, l the length of the beam, d the vertical de#ection of the beam, p the maximum z x bending stress of the beam, and E is Young's modulus of the adherends. For the bonded model, the calculated value is f + 1.2. f is a function of beam curvature which depends on the dimensions of section, boundary conditions and type of adhesive [33]. Further research into this area is required in order to produce typical values of f for bonded steel panels. However, a very recent study [37] on a bonded glass T-beam (plate and web) used a simple theory modi"ed from sandwich beam analysis and claims a good prediction for shear and bending stresses in the composite beam. The estimated f value from that study was approximately 1.1, considering the fact that the adherend is glass which has one-third the Young's modulus of the steel. The modi"ed theory seems very sensitive to adhesive shear modulus of elasticity and thickness. High stresses and de#ections in a bonded panel may imply that bonded structures should be slightly heavier than welded equivalents. This, however could be compensated for by the absence of welding residual stresses in the plating and the freedom to bond thinner plates with closer sti!eners without the technical and economical problems associated with controlling thermal distortion during welding of such geometries. Bonded sti!eners would also increase the e!ective breadth of panels due to the wide #ange attachment and absence of distortion.
6. Local behaviour of sti4ener/plate attachment Local stresses in adhesive joints are largely non-linear and typical local non-linear stress distributions in small bonded joints are shown in Fig. 10. This non-linearity arises from the loading direction and the great mismatch between the elastic modulus
Fig. 10. Local stresses in typical small joints [38].
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417
of the adherends and adhesive. The non-linearity makes detailed stress and strain analysis, for the purpose of fatigue prediction for example, di$cult without the use of suitable numerical methods. Because bonded panels are considered for light loading, local failure loads in sti!ener/plate joints may only arise as a result of extreme loading, such as in impact. To understand the cleavage behaviour of sti!ener/plate joints, a simpli"ed model was used (Fig. 11) to examine the cleavage failure under transverse loading on a sti!ener. A solid block, instead of a sti!ener, was used in this model to insure that cleavage failure takes place in the adhesive. Using a sti!ener such as rolled steel joist (RSJ) section would normally lead to plastic deformation of the web rather than adhesive joint failure [1]. The model was also used to study the in#uence of local plate sti!ness on cleavage stress. This was achieved by considering two lengths (200 and 300 mm) for the 8 mm plate, as a function of e!ective breadth in sti!ened panels. The model was tested statically until adhesive failure. Table 3 gives failure load results together with dimensions and contains adhesive cleavage stresses calculated from the numerical and analytical methods discussed in a later section. The above results may be related to the spacing of sti!eners in bonded panels (using the same plate thickness) that are subjected to transverse bending moment. A bonded panel with close sti!eners would produce lower cleavage stresses than one with widely
Fig. 11. Idealised bonded sti!ener/plate joint.
Table 3 Dimensions and results from analysis and test of sti!ener/plate models Plate length B (mm)
Joint length (mm)
Joint width (mm)
Failure load (kN)
Eq. (4) stress (MPa)
FE stress (MPa)
300 200
75 75
45 45
10.4 11.7
148 148
170 170
418
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spaced sti!eners. The closer the sti!eners the lower is the local de#ection of the plate. The de#ection largely controls the level of the cleavage stresses and hence the failure of the joint. Other factors which a!ect structural joint design are sti!ener foot shape, size and the detail of sti!ener ends. Experimental results based on RSJ sti!ener/plate joints are reported elsewhere [7]. In general, design of load bearing joints should consider the minimisation of cleavage stress and utilise shear and compression strength of the adhesive. These considerations are reported frequently in the literature [12,16] by producing `DOsa and `DON'Tsa guidance/recommendations for bonded joints. 7. Elastic stress analysis Elastic "nite element (FE) modelling has been widely used to assess strength in adhesive joints, with reasonable reliability. One approach to interpret adhesive failure criteria for bonded joints may be based on von Mises or principal stresses, occurring at a prescribed location from the edge of joint. Because of the mismatch of materials, FE models for adhesive joints exhibit points of theoretical singularity, especially at joint edges where stresses become in"nite. Errors in failure prediction of up to 20% might be expected with this approach [34]. Because high stress gradients occur in certain regions of bonded joints, these regions have to be modelled very accurately and economically. The adhesive layer is very thin (0.1}0.5 mm) compared with the thickness of the adherend (several mm), and to achieve reliable results, it is necessary to use several elements through the thickness of the adhesive. Owing to the resulting "ne mesh and the limitations of element aspect ratios, the number of elements and hence degrees of freedom in a joint is, typically, rather high. Fig. 12 shows a typical mesh and boundary conditions for the sti!ened joint (Fig. 11). Normally, the adhesive layer is modelled by at least "ve elements through the thickness with the highest mesh density nearer the tension edge of the joint. Many
Fig. 12. Finite element model for sti!ener/plate joint.
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419
packages are suitable for elastic FE analysis. Fig. 13 shows typical stress distributions along the 45 mm of the adhesive joint/line at 0.1mm above the plate interface with the adhesive. The maximum stress is expected to be at the plate surface, but would be di$cult to determine, due to the sudden change in material properties. The stress distribution is clearly non-linear and with high stresses concentrated in a small region of the joint. A similar pattern of stress distribution was also found in the thick adherend lap shear joint [1] (Fig. 2). This is shown in Fig. 14 where again the stresses along the adhesive line were taken at 0.1 mm above the lower adherend. When comparing the principal stresses at the ends of both joints (Figs. 13 and 14) good correlation between the peak principal stresses can be obtained. This indicates the suitability of the lap shear joint for the purpose of stress prediction for stresses arising from local loading on the sti!ened joint. An elastic analytical method to calculate local cleavage stress for sti!ened joints (L and T shapes) was developed by Crocombe et al. [35] from the classical lap shear analysis by Goland and Reissner [36]. This analysis considered the adhesive joint as
Fig. 13. Stress distribution along the adhesive line for a sti!ener/plate joint.
Fig. 14. Stress distribution along the adhesive line for a lap shear joint.
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Fig. 15. Analytical model for stress calculations.
adherend-adhesive sandwich subject to transverse shear forces and bending moments at the ends. Fig. 15 illustrates the basic model where the adherends are assumed to behave as cylindrically bent plates of di!erent #exural sti!ness. As a result of the di!erential in vertical displacement, the adhesive transmits cleavage stresses, with the maximum at the ends of joints. The expression for the maximum cleavage stress at the joint edge is p "Ak , (4) z a where p is the cleavage (normal) adhesive stress, and k the e!ective spring sti!ness of z a the adhesive. w E k " a a, a ¹ a where w is the adhesive joint width, E the adhesive Young's modulus, ¹ the a a a adhesive thickness, and A the constant function of forces and moments and adhesive thickness [1,35]. This design formula was applied to the models (joints) shown in Fig. 11. The calculated results are given in Table 3 which show realistic stress values in comparison with equivalent results from "nite element analysis. This formula, however, is sensitive to changes in adhesive thickness and does not account for variation in stresses through the thickness of the adhesive. It also, unlike the "nite analysis method, requires the designer to determine the forces and moments at the boundaries of the joint.
8. Conclusions and further work A number of materials and design requirements have been discussed, in relation to steel adherends and structural epoxy adhesives. The key conclusions and recommendations are; f Cleavage and shear strength values are very useful comparative data for selecting a structural adhesive. f Formulation of relevant small and large experiments, suitable for thick steel adherends, is a key to understanding and evaluating joint design and behaviour.
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f High strength structural adhesives are typically limited to operate at a temperature up to 1003C, which is acceptable for many marine applications. f Long-term exposure of loaded small lap shear joints to a severe wet environment causes a strength reduction of approximately 10% per year. Further work is required to establish whether the e!ect would be smaller in larger, practical joints and structures exposed to normal marine service conditions. f Beam theory for calculating stresses and de#ections in sections made up of two longitudinal elements bonded together may be simply modi"ed to account for the presence of the adhesive through the use of an interface coe$cient f. Further experimental work is necessary to obtain appropriate f values for bonded panels and materials. f FE analysis is suitable to evaluate local strength in sti!ened joints, especially when existing analytical techniques are correlated with results from lap shear joints. f Existing analytical methods, while less accurate than FE analysis, can be a very useful design tool to determine local stresses.
Acknowledgements This work was funded by Science and Engineering Research Council (Grant No. GR/C/7139.2/1G95) through the Marine Technology Directorate Ltd and the Ministry of Defence. Such support is gratefully acknowledged.
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