Proceedings

W056 - Special Track

18th CIB World Building Congress

May 2010 Salford, United Kingdom

CIB W056 - Sandwich Panels

CIB Publication 342

CIB WORKING COMMISSION

W056 ‐ SANDWICH PANELS

PAPERS AND POSTGRADUATE PAPERS FROM THE SPECIAL TRACK

HELD AT THE CIB WORLD BUILDING CONGRESS 2010, 10‐13 MAY 2010

THE LOWRY, SALFORD QUAYS, UNITED KINGDOM

Selected papers from the Proceedings of the 18th CIB World Building Congress. Proceedings edited by: Professor Peter Barrett, Professor Dilanthi Amaratunga, Dr. Richard

Haigh, Dr. Kaushal Keraminiyage and Dr. Chaminda Pathirage

W056 Special Track Papers (excluding Postgraduate Papers) reviewed by: Professor Jose

Amorim Faria, Professor Jukka Aalto, Professor Jörg Lange, Dr. Lars Heselius and Professor J. Michael Davies

CIB Publication 342

W056 ‐ SANDWICH PANELS

PAPERS AND POSTGRADUATE PAPERS FROM THE SPECIAL TRACK

The objectives of this Commission are to exchange information and coordinate research programmes in all technical aspects of light‐weight constructions using traditional and new building materials and to prepare synthesised reports on matters of particular interest.

CONTENTS

Papers

Stabilisation of Beams by Sandwich Panels ‐ New Regulations and Recent 1 Research Results ‐ Misiek, T. Kapplein, S. Durr, M. Saal, H.

A Unified Approach for the Local Buckling of Sandwich Panels and Trapezoidal Sheeting 14 Misiek, T. Hassinen, P.

ECCS/CIB Joint Committee on Sandwich Constructions, Recent European 27 Recommendations on Design and Testing Davis, J.M. Hassinen, P. Heselius, L. Misiek, T.

Methods to Measure the Durability of Structural Sandwich Panels 38 Hassinen, P. Pfeiffer, L.

Reliable Composite Roofing: Learning from Experience 51 Roberts, K.

Postgraduate Papers

Openings in Sandwich Elements 61 Warmuth, F. Lange, J.

Optimization of Geometry and Core Materials of Sandwich Panels with Metallic Faces 73 Kurpiela, A. Lange, J. Berner, K.

CIB Brochure 85

Disclaimer 87

5

A Unified Approach for the Local Buckling Of Sandwich Panels and Trapezoidal Sheeting

Misiek, T.

Karlsruhe Institute of Technology, Karlsruhe, Germany

(email: thomas.misiek@kit.edu)

Hassinen, P.

Faculty of Engineering and Architecture, Helsinki University of Technology, Espoo, Finland and

Pontek Consulting Engineers Ltd, Espoo, Finland

(email: paavo.hassinen@tkk.fi)

Abstract

Failure of thin-walled building components like sandwich panels or trapezoidal sheeting is normally

initiated through a local buckling of the plane elements of the cross-section. For trapezoidal sheeting,

EN 1993-1-3 gives an equation for the determination of the effective width of these plane elements

and thus for the calculation of the load-bearing capacity of these components. The cross-sectional

parts of a lightly or strongly profiled facing of a sandwich panel can be regarded as elastically

supported plane elements, whereas the elastic support is provided by the core material. For the

determination of the load-bearing resistance of sandwich panels with lightly or strongly profiled

facings, a calculation procedure to determine the effective width of the elastically supported plane

elements is needed. Some approaches for the calculation of the effective width already exist. The

papers published so far are using a modified buckling coefficient for the calculation of the buckling

strength, following Winter’s approach such as given in EN 1993-1-3. Because no generally accepted

design procedures exist, the load-bearing resistance of sandwich panels is determined

experimentally. Based on the basic principles of structural stability, the buckling strength of the

elastically supported plane element can be calculated, taking into account the material properties of

the core material and the associated buckling wavelength for minimum buckling strength. Then the

design procedures of EN 1993-1-3 for thin plate buckling of trapezoidal sheeting can be used,

expanded by the procedures of EN 1993-1-5 for taking into account the column type buckling

behaviour for buckling wavelengths smaller than the total width of the plane element. Comparison of

the test results with different arrangements to the calculated values shows a good consistency. For

lightly profiled faces, depending on the depth of the profiling, failure will finally take place through a

plate buckling of the plane elements or by a column buckling of the stiffening profiles, whereas the

latter failure mode is looking similar as the wrinkling failure of a flat facing. Comparison of the

experimental results obtained with different test arrangements to the calculated values show the

present limits of the applicability of the proposed design procedure. The differences in failure modes

are discussed.

Keywords: sandwich panels, face layer, buckling, elastic foundation, elastic half-space

6

b

1. Introduction

The determination of the load-bearing capacity required for the design of sandwich panels is to a

large degree based on test results. In contrast, there is a large number of references for the

determination of the load-bearing capacity of the trapezoidal sheeting, documented in national

standards such as the Swedish StBK-N5 or the German DIN 18807 as well as international standards

such as EN 1993-1-3. This is astonishing because both building components are com-parable in

materials, geometry and load-bearing behaviour. Compared to trapezoidal sheeting the foam core

provides an additional elastic foundation to the plane thin-walled elements.

In the following we will show that the calculation procedures developed for trapezoidal sheeting can

be modified to be applied for sandwich panels. This will start with the basic module, of which all

cross-sections consist: a simple plane element, like a flange of a strongly profiled face of a sandwich

panel.

2. Theoretical background and standardized design procedures

The local buckling of the plane cross-sectional elements shall be taken into account in the

determination of the load-bearing capacity of the thin-walled building components. Because of the

local buckling of the plates of a medium or high slenderness only an effective width, which is smaller

than the total width, can be taken into account.

x

a

y

EIF

= f (EF, t,

F)

x

c = f (EC, G

C,

C)

z

Figure 1: Plate on elastic foundation

The calculation of the effective width is based on the critical buckling stress cr. An additional

support provided by an elastic foundation increases the critical buckling stress cr. This increase is an

7

E t 2

a 2 2 2 2

F

2 2 2

p

addition to the local buckling stress of the plate without foundation calculated according to EN 1993-

1-3. The increase of the critical buckling stress leads to a decrease of the relative slenderness of the

plate and thus to an increase of effective width and load-bearing capacity. Up to now, most of the

present studies capture the effect of the elastic foundation by adjusting the buckling value k as

proposed by Davis, Hakmi and Hassinen (1991). This adjustment is based more on statistical

evaluations of test results. The determination of the critical buckling stress can be based on the elastic

potential, which is shown in Timoshenko and Woinowsky-Krieger (1959) for the plate without

foundation. The calculation of the stiffness of the Winkler’s foundation out of the material

parameters EC and GC of the core material is shown in Stamm and Witte (1974). We obtain the

following equation:

cr , p

2

F m n

12 1 2 m a b

2 1 E

a m n

C C

1 C 3 4 C t m a b

with

E C EC GC 2 1 C

according to Stamm and Witte (1974) to take into account very approximately the anisotropy of the

material properties of core materials such as for example polyurethane and polystyrene foam.

The effective width of the plate can be calculated from the critical buckling stress and the relative

slenderness. EN 1993-1-3 gives the following expressions:

f y p

cr , p

beff , p

1,0 1 0,22 for p

p

0,673

b 2 p

p p

0,673

These equations derived by Winter assume pure plate-like behaviour. This can be found for ratios =

a/b > 1,0. Because of the elastic foundation we obtain a buckling wave pattern with an aspect ratio

a/b which is smaller than 1,0. This has to be taken into account when calculating the critical buckling

stress, leading to an increase of the buckling coefficient k . Unfortunately this also results in a minor

support of the mid-part of the plate by the supports at the longitudinal edges compared to buckling

patterns with higher aspect ratios. Therefore the load-bearing capacity is lower than the one

calculated according to Winter’s equation. Column-like buckling behaviour of the plate has to be

taken into account. This can be done by utilising the procedure given in EN 1995-1-5. In this case an

interpolation between the plate buckling curve according to Winter and the column buckling curve c

8

2 2 2

F

is done. This interpolation is based on the ratio of the elastic critical buckling stress of the plate and

the column:

c p c 2 c

with

0 1,0 where cr , p

1

cr ,c

and

cr ,c

EF t m

12 1 2 a

2 1 C

1 C 3 4 C

E C a t m

f y c

cr ,c

0,5 1

0,21 2

c 0,2 c

1 c

2 2

c

1,0

3. Comparison with results from numerical calculations and tests with flat plates

The presented approach for the determination of the load-bearing resistance of a flat plate on an

elastic foundation was verified using finite-element-calculations and comparisons with test results.

The plates were modelled using four-node shell elements with a linear-elastic ideal-plastic stress-

strain relationship. For the elastic foundation, spring elements with a linear elastic stress-strain

relationship were used. The parameters such as sheet thickness t, yield strength fy and stiffness of the

foundation c (via EC, GC and C) were varied. Figure 2 shows the results of the comparison between

the theoretically derived critical buckling stress and the buckling stress based on FE-calculations.

Using this FE-model and the initial imperfection corresponding to the first eigenvalue of the model,

the ultimate load has been computed. The initial imperfection was scaled to w0 = 0,1 t. The

comparison of the computed results with the results of the procedure based on the interpolation

between the plate buckling curve according to Winter’s equation and the column buckling curve c is

shown in Figure 3.

9

cr,p

,FE

M [

N/m

m²]

c,F

EM

[-]

600

500

400

300

200

100

0

0 100 200 300 400 500 600

cr,p,calc [N/mm²]

Figure 2: Comparison of critical buckling stress based on the finite-element-calculations and on the

analytical expressions

1,0

0,8

0,6

0,4

0,2

0,0

0,0 0,2 0,4 0,6 0,8 1,0

c,calc [-]

Figure 3: Comparison of the results of finite-element-calculations with results of the calculation

procedure: normalised effective width c

To verify the calculation procedure, a series of tests was performed. These tests were done with

trapezoidal sheeting (five ribs) and with cross-sections cut out from the sheeting (one rib). The

trapezoidal sheeting was the upper face of a panel with one strongly profiled face (Figure 4). The

10

inner flat face was removed. Thus, the trapezoidal sheeting provided the only load-bearing

component of the specimen. Due to the remaining foam core, the plane elements of the face still were

supported by an elastic foundation. The normalized effective width c was back-calculated using the

calculation procedure for trapezoidal sheeting according to EN 1993-1-3, supplemented with the

procedures of EN 1993-1-5.

20 mm

23 mm

114 mm

tF1

= 0,47 mm

35 mm

61mm

no face

Figure 4: Trapezoidal sheeting with elastic foundation

According to the calculation procedure, the aspect ratio of the buckling waves should be about =

0,3. Figure 5 shows the buckling waves in the upper flange confirming the calculated value.

Figure 5: Local buckling of the upper flange

An additional series of tests was done with the upper face layer without any foam material. The

geometry and material was identical to the one used in the tests described above. In this case no

elastic foundation for the plate is available and we have a cross-section of a standard trapezoidal

sheeting. Results of the tests are shown in Figure 6.

In addition, the tests results of Davis and Hakmi (1991) and the results of Pokharel and Mahendran

(2003) were used for comparison (Figure 6). Davis and Hakmi performed bending tests on C-shaped

11

c,t

est

[-]

members with a foam core. Pokharel and Mahendran glued flat sheets of different thicknesses and

steel grades to layers of plastic foam material.

1,0

0,8

0,6

0,4

0,2

0,0

Davis and Hakmi 1991

Pokharel and Mahendran 2003

trapezoidal sheeting with elastic

foundation

trapezoidal sheeting without

elastic foundation

0,0 0,2 0,4 0,6 0,8 1,0

c,calc [-]

Figure 6: Comparison of calculation procedure with test results: normalised effective width c

All the test results show a good correlation with the calculated values. The calculated values make

the lower boundary to the test results. Interestingly the values with the highest discrepancy between

test and calculation being on the unsafe side are the ones for the standard trapezoidal sheeting

without any elastic foundation. This is a result of the very high slenderness of the upper flange, the

relative slenderness being of the order of 5,5 to 5,6.

It is interesting to discuss about the effect of the material properties used in the calculation procedure.

The first difference arises in the test method for the determination of the shear modulus GC. Davis

and Hakmi performed tests with small cubes of the core material which were loaded with a pure shear

load. No information is available about the tests of Pokharel and Mahendran. However, it can be

assumed that the tests were done in a similar way based on small cubes. Our tests for the

determination of the shear stiffness were performed with small short-spanning beams according to

EN 14509. The latter test method might lead to higher discrepancies between the measured value and

the actual material properties, resulting in a higher discrepancy between the calculated and

experimental values c,calc and c,test.

The elastic modulus EC of the core material is normally determined as an average value

corresponding to the total thickness of the specimen, determined separately in a compression and

tensile test. For plastic foam materials made in a continuous process the local values strongly deviate

from this mean value, depending on the relative position of the sample in the thickness direction of

the panel. For these sandwich panels, investigations on the distribution of the stiffness presented in

12

Dürr (2008) show that the elastic modulus near to the faces is higher than in the centre plane of the

panel. In addition, different values of the modulus can be obtained close to the faces depending on

their position (top or bottom) during the production. These effects have to be kept in mind when

looking for the results of our tests (which were done with real panels) and presumably also for the

tests of Davies and Hakmi. For the tests of Pokharel and Mahendran which were performed with

foam material glued to the face this effect can be neglected. In fact, when calculating the load bearing

resistance with the values from the tests, calculated values are on the safe side, because the higher

stiffness near the faces gives a stiffer elastic foundation than assumed by using the mean values over

the total core thickness.

4. Strongly profiled sandwich panels

In addition to the buckling tests with trapezoidal sheeting, tests with sandwich panels were also

performed. The panels had a similar geometry as the specimens introduced before. In these tests, the

inner face was not removed, so real sandwich-type load-bearing behaviour existed. The results of the

tests were back-calculated, too. Now, the lower flange of the outer face is also under compression

loading. This can be easily seen by comparing the pictures of this flange taken after the failure

(Figure 7).

Figure 7: Upper and lower flange of the specimens after failure: trapezoidal sheeting with elastic

foundation (left) and sandwich panel (right)

The effective cross-section of the outer face can be calculated with the introduced procedure. At first,

the effective cross-section of the outer face under compression has to be determined by strictly

following the calculation procedure introduced in chapter 2 (approach A). In fact, there will be

always a difference between the calculated effective width and the test results. These differences in

the determination of the effective width will cause further differences in the calculation of the load-

bearing capacity of the sandwich cross-section. In our case, we have also the possibility to use the

effective width of the upper flange from the tests on tests described in chapter 3 (approach B) to

minimize the effect of the former differences. The use of both the both approaches allows us to

compare the results. The calculation has to be done iteratively. For the first iteration we assume a

13

constant compression stress distribution over the whole cross-section of the outer face. We obtain the

following results:

20 mm

23mm 114 mm

11,1 mm

35 mm

11 mm

11,1 mm Approach A: 15,4 mm Approach B: 22,8 mm 11 mm

Figure 8: Effective cross-section of the outer face after the first iteration

Table 1: load-bearing capacity of the sandwich panel after the first iteration

Approach A

Approach B

Test result

Ultimate bending moment

5,27 kNm

6,74 kNm

6,94 kNm

According to this calculation, the compression stress in the lower flange reaches a value of = 148

N/mm² which is below the yield strength. The effective width in the lower flange is larger than

assumed at first. Same is true for the lower part of the webs. Therefore the effective cross-section is

larger than calculated in this first step. The accuracy of the calculated load bearing capacity can be

improved iteratively. For the lower flange, this can be done by calculating the effective width with

= 148 N/mm² instead of fy = 400 N/mm². Special considerations are required for the web. We can use

the simple assumption given in EN 1993-1-3 that the effective width at the lower edge is 1,5-times

the one at the upper width were = fy. After the first iteration, the stresses in the lower flange are

further decreasing, and the stresses in the inner face increase. Load-bearing capacity is increasing

with every iteration. However, the values are converging very soon. Stress distribution based on the

fifth iteration can be accepted to represent the ultimate load:

20 mm

23mm 114 mm

11,1 mm

35 mm

16,7 mm Approach A: 15,4 mm

Approach B: 22,8 mm

Approach A: 20,0 mm

Approach B: 19,0 mm

Approach A: 20,0 mm

Approach B: 19,0 mm

Figure 9: Effective cross-section of the outer face after the fifth iteration

14

Table 2: Load-bearing capacity after five iterations

Approach A

Approach B

Test result

Ultimate bending moment

5,94 kNm

7,14 kNm

6,94 kNm

This is a considerable improvement compared to the results of the first iteration. However, in the

example we have a difference between the values of the pure calculational approach A and the test

result. The difference is in the same order as for the simple flat plate element.

5. Resistance of lightly profiled faces

A local buckling phenomena can also be observed in tests with panels having a slightly profiled face.

Figure 10 shows the face of a lightly profiled sandwich panel (depth of lining approx. 2,0 to 2,5 mm)

in a bending test before reaching the failure load. Finally, failure will occur by global buckling or

wrinkling of the face over the complete width of the panel, depending on the support conditions at the

longitudinal edges of the compressed face.

Figure 10: Local buckling of the slightly profiled face of a sandwich panel

For an extension of our approach to lightly profiled faces on an elastic support, the results of the

buckling tests of Pokharel and Mahendran (2005) were used. Pokharel and Mahendran (2005) varied

the thickness of the faces and the width b of the plane elements of the profiling. The total dimensions

of the faces were 400 mm and 1200 mm in the width and length directions, respectively. All the

edges were fully supported. The effective width of the plane elements between the stiffeners can be

calculated according to the presented calculation procedure. After incorporating the restrictions in

EN 1993-1-3 requiring some reductions in width also the cross-sectional values of the stiffeners itself

can be calculated. The part of the critical buckling stress provided by the stiffeners can now be

calculated using

15

b b

Nu

,test

[k

N]

cr ,s

1,8 E

I s t 2 3

e

Assuming a global buckling failure of the stiffened plate and recalculating the tests with the

equations introduced in section 2 of this paper, the results shown in Figure 11 were obtained.

40,0

4,25 mm 4,25 mm

13 mm b

32,0 1,0 mm

24,0

16,0

8,0 b = 78,5 mm

b = 28,5 mm

0,0

0,0 8,0 16,0 24,0 32,0 40,0

Nu,calc [kN]

Figure 11: Comparison of results of calculation procedure with test results

It can be shown that for the geometry investigated in Pokharel and Mahendran (2005) the effects of

the stiffeners can be neglected and the support at the longitudinal edges as well as the elastic

foundation of the core material are the dominating stabilising effect. In this case, failure occurred by

local buckling of the entire stiffened plate. The overall width dominated against the width of the

lining between the stiffeners.

Typical flat or lightly profiled faces of sandwich panels do not have supports at the longitudinal

edges. The wrinkling stress of these panels is usually calculated using the so-called Plantema-

equation.

cr 0,82 2 E

F E

C G

C

The coefficient 0,82 is in most cases modified to take into account the effects of imperfections etc.

By doing so the transition form the elastic critical buckling tress to the ultimate buckling stress is

done. Further modifications of this coefficient, sometimes denominated as a wrinkling factor, can be

justified by taking into account the geometry of the faces, leading to a value derived from tests.

Further adjustments of this equation are discussed in Pokharel and Mahendran (2005).

16

This problem concerning the flat or lightly profiled faces of typical sandwich panels without a

support on the longitudinal edges could also be calculated within the framework of EN 1993-1-3. The

wrinkling failure can be interpreted as a buckling failure of the stiffeners of the face. After

calculating the elastic critical buckling load of the stiffener, a buckling curve for the stiffener

available in EN 1993-1-3 could be used. The relative slenderness and the buckling curve are given by

f y

d

cr ,s

and

1,0 d 0,65

d 1,47 0,723 d for 0,65 d 1,38

0,66 d 1,38

d

This buckling curve was developed for trapezoidal sheeting using the assumption that the elastic

foundation of the stiffener is provided by the support of the neighbouring longitudinal edges of the

stiffened element.

Further investigations have to be done before applying the design procedures of EN 1993-1-3 for the

design of lightly profiled sandwich panels. At present, the observations can be used to make a

differentiation between lightly and strongly profiled faces of sandwich panels. When calculating the

critical buckling load of the stiffener, the width bP = sW of the web has to be fully effective, thus c =

1,0. As a second criteria, column-buckling of this stiffener must not be the significant failure mode,

so d = 1,0 for column buckling of the stiffener. Then, the face can be regarded as strongly profiled

and the load-bearing capacity of the panel can be calculated as presented in the previous chapters.

6. Conclusion

The determination of the load-bearing capacity required for the design of sandwich panels is to a

large degree based on test results. In contrast, there is a large number of references for the

determination of the load-bearing capacity of trapezoidal sheeting. Because both building

components are comparable in materials, geometry and load-bearing behaviour, a calculation

procedure for sandwich panels which relies on the design principles developed for trapezoidal

sheeting seems to be possible, by taking into account the effects of the elastic foundation provided by

the core layer.

The basic principles of the design model are introduced and compared with Finite-Element-

calculations as well as with test results from different sources, showing the applicability of the

procedures. Apparently this calculation procedure can be applied to strongly profiled faces of

17

sandwich panels as well as for trapezoidal sheeting with bonded plastic foam insulation. The last-

mentioned building components have found increasing dissemination on the market in the last years.

Finally, an overview to the design of lightly profiled sandwich panels was made. Comparison with

test results showed the present limits of the applicability of the design procedures of EN 1993-1-3 if

applied to sandwich panels. Lightly profiled faces show a complex behaviour with interactions from

plate buckling and column buckling failure. This can not be captured on a sufficient safety level at

the moment.

Acknowledgements

ThyssenKrupp Steel Europe AG provided specimens for the experimental investigations. We express

our sincere gratitude for this support.

References

Timoshenko S P Woinowsky-Krieger S (1959) Theory of Plates and Shells, New York, McGraw-

Hill.

Stamm K and Witte H (1974) Sandwichkonstruktionen: Berechnung, Fertigung, Ausführung, Wien

New York, Springer Verlag

Davies M J, Hakmi M R and Hassinen P (1991) “Face buckling stresses in sandwich panels”

Contributions 1991 ECCS Nordic Steel Colloquium, p. 99-110.

Davies M J and Hakmi M R (1991) “Postbuckling behaviour of foam-filled thin-walled steel beams”

Journal of Constructional Steel Research 20: 75 - 83.

Pokharel N and Mahendran M (2003) “Experimental investigations and design of sandwich panels

subject to local buckling effects” Journal of Constructional steel research 59: 1533-1552.

Pokharel N and Mahendran M (2005) “An investigation of lightly profiled sandwich panels subject to

local buckling and flexural wrinkling effects” Journal of Constructional steel research 61: 984-1006.

Dürr M (2008) Die Stabilisierung biegedrillknickgefährdeter Träger durch Sandwichelemente und

Trapezbleche, Berichte der Versuchsanstalt für Stahl, Holz und Steine der Universität Fridericiana in

Karlsruhe, 5. Folge – Heft 17. Karlsruhe 2008.

EN 1993-1-3: Eurocode 3: Design of steel structures – part 1-3: General rules - Supplementary rules

for cold-formed members and sheeting

EN 1993-1-5: Eurocode 3: Design of steel structures – part 1-5: Plated structural elements

18

ECCS/CIB Joint Committee on Sandwich Constructions: Recent European Recommendations on

Design and Testing

Davis, J.M.

83 Park Road Hale, Altrincham Cheshire WA15 9LQ UK

(email: profjmdavies@lineone.net)

Hassinen, P.

Helsinki University of Technology, Finland

(email: paavo.hassinen@tkk.fi)

Heselius, L.

LHH Consulting Oy Ab Ltd

(email: lars.heselius@lhconsulting.eu)

Misiek, T.

Karlsruhe Institute of Technology, Karlsruhe, Germany

(email: thomas.misiek@kit.edu)

Abstract

Light-weight sandwich panels, made of two metal faces separated by an insulating core material, are

modern pre-fabricated construction components. The design, manufacture and use of these

composite structural elements have required continuous development of regulations and standards.

The Joint Committee is the combination of two working groups; CIB W056 “Lightweight

constructions” and ECCS TWG 7.9 “Sandwich panels and related structures”. It therefore provides a

powerful forum for the consideration of international research and development of the design, testing

and end-use of sandwich panels. This paper introduces the Joint Committee and its work by outlining

its history and its most important publication, “European Recommendation for sandwich panels”.

The significant influence of the Committee on the recently published European Product Standard EN

14509 is also described. Focus is then directed to the latest publications of the Joint Committee,

namely the “Preliminary European Recommendations for Testing and Design of fastenings for

sandwich panels” and the “State of the art report for Design of Sandwich Panels with Openings”.

The latter is still under preparation and will probably be published in 2010.

Keywords: sandwich panel, light-weight, fastening, opening, CIB, ECCS, recommendations

19

1. The joint committee on sandwich constructions

1.1 CIB steering group S56 and working commission W056

The first sign of the existence of the CIB Steering Group S56 was at the International Symposium on

Low-Rise Lightweight Constructions held in Budapest in April, 1971. A significant product of this

era was CIB Publication 59 “Recommendations for the Structural Design of Lightweight Sandwich

Panels” which was published in 1978. This document gives principles for design with regard to static

out-of-plane and in-plane actions, impact loads and dynamic loads. It also gives recommendations

for design for environmental effects and fire. Finally, it discusses the properties of connections and

construction systems including the principles of the quality control. The report was based on ISO

standard 2394-1973 which introduces the principles of the modern limit state design. The report,

which was far ahead of its time, was written as a framework for national regulations and it probably

provided the first formalised recommendations for the design of sandwich panels. Nowadays, the

emphasis in Europe has moved away from national guidelines and towards harmonized standards for

the whole continent. However, historically, “Recommendations” produced by European-wide

Committees of experts have led to a similar unified outcome.

The name of Steering Group S56 was changed to Working Commission W056 and the membership

reconstituted. It held its first meeting in Espoo, Finland in May, 1984. Initially, developments in

lightweight constructions were introduced and discussed without limitation as to materials or types of

construction. Later, the subjects of W056 became focused more towards mineral wool cored

sandwich panels because of the need for guidelines for this new sandwich panel product. The

coordinators of W056 during this period were Professor J M Davies in 1986 - 89 and Lars Heselius in

1990 - 2006. Thus, W056 extended the Preliminary European Recommendations for Sandwich

Panels written originally by ECCS TWG 7.4 (see 1.2) and the outcome was CIB Publication No 148

which was published in 1993 and reprinted in 1995.

1.2 ECCS technical working group TWG 7.4

The European Convention for Constructional Steelwork (ECCS) has a long and distinguished record

of pre-standardisation work and its Technical Committee TC7 “Cold-formed thin-walled sheet steel

in buildings” has been one of its most successful working committees and, since its inception in

1974, it has produced numerous significant documents. The workload of TC7 quickly increased to

the point where a number of separate Working Groups were required and in 1983, noting the

increasing interest of the market in light-weight metal sheet faced sandwich construction, it was

decided to form a new technical working group TWG 7.4 “Sandwich Panels” to produce European

Recommendations for design, testing and good practice in the use of sandwich panels. The work of

TWG 7.4 started in 1983 under the chairmanship of Dieter Stemman taking full advantage of his

experience with German “Zulassung” documents.

20

TWG 7.4 produced “Preliminary European Recommendations for Sandwich Panels” in two parts.

Part I “Design” was published in 1991 and relates to the analysis, design and testing for the load-

bearing capacity of the sandwich panels. This significant document has provided the basis for a

number of subsequent Recommendations and Standards. Part II “Good Practice” was published in

1990 and introduces subjects such as building physics, fire, installation and erection work.

In the late nineties needs to update the European Recommendations became obvious and ECCS TC7

founded a new Technical Working Group TWG 7.9 “Sandwich Panels and related Subjects”.

Because of the mutual interest, in 1998, ECCS TWG 7.9 and CIB W056 combined to form the Joint

Committee on Sandwich Construction, which is still active today. The first two years of work

resulted in the final manuscript of the “European Recommendations for Sandwich Panels, Part 1:

Design”, which was published as CIB Publication No 257 in 2000 and ECCS Publication No 115 in

2001. The authors pay tribute to the work of Antti Helenius who acted as secretary of W056 and the

Joint Committee for more than ten years.

Based on the work of the CIB W056 Working Commission, a group of members combined together

to write a book covering all aspects of sandwich panel design and construction. Although individual

members took responsibility for individual chapters, this book is unique in that the authors took joint

responsibility for the whole and each chapter was discussed in detail at a series of meetings. The

contributions were edited by Professor J M Davies and the result was published in 2001 entitled

“Lightweight Sandwich Construction”.

The above Recommendations have formed the basis of the European product standard for self-

supporting double skin metal faced insulating panels, EN 14509, which was published in December

2006 and cited in the Journal of CEN in December 2008. Drafting of the product standard was in the

hands of CEN Committee TC128/SC11, which includes most members of the Joint Committee. The

product standard extends the European Recommendations to include all of the “Essential

Requirements” of the Construction Product Directive in accordance with formal mandates.

Because it is a European “product” standard, EN 14509 defines the requirements and the methods of

verification of the essential properties of the factory made sandwich panels. The standard does not

and cannot specify requirements for operations carried out after manufacture such as cutting, fixing

and installation. Consequently, in 2004, the reactivated Joint Committee identified a number of

subjects in which further direction and harmonization is needed. The first items to be considered

were the fastening and openings in sandwich panels.

21

2. The recent European recommendations

2.1 Testing and design of fastenings of sandwich panels

2.1.1 Background

The product Standard EN 14509 does not give any requirements regarding the joints and fastenings

of the sandwich panels. In order to remedy this, the Joint Committee has produced preliminary

recommendations which discuss the experimental determination of the load-bearing resistance of

fastenings, the evaluation of the test results and the principles of the design of fastenings. These

preliminary recommendations, now available as ECCS Publication No 127 and CIB publication No

320 in 2009, are based on the earlier versions of the European Recommendations for Sandwich

Panels ECCS & CIB (2000) updated to accord with current knowledge and experience.

2.1.2 Fastening of sandwich panels

Figure 1: Fixings of sandwich panels: typical application and fastener (right hand Figure courtesy of

Würth).

Because of the inherent rigidity of the sidelap, seam fasteners are not generally used in sandwich

construction. Attention is therefore concentrated on the connections to the supporting structure.

These may be concealed in the sidelap or may be visible and pass through the complete panel.

Whichever solution is adopted, self-tapping or self-drilling screws are generally used. Both of these

are able to cut their own threads into the substructure, however self-tapping screws require pre-

drilling. The screws have a formed drilling bit that allows the screwing and drilling in a single

22

F

operation. In order to produce a rainproof joint, sealing washers with a vulcanised EPDM layer are

used between the screw head and the face. The washers of the screws fixed to the upper flange of the

trapezoidally profiled facing, may take the form of saddle washers to tighten and to support the

profile.

The fasteners between the sandwich panel and the substructure may be loaded by tensile and shear

forces as well as by bending moments (head deflection) due to the thermal movements of the faces.

2.1.3 Tensile and shear resistance of fastenings

The load-bearing performance of fastenings is mainly determined experimentally. Therefore, tensile,

shear and bending tests have to be performed. Attention is generally concentrated on the strength

(capacity) of the connection. However, it is important to appreciate that stiffness and ductility are

also important. In particular, sandwich panel assemblies are generally extremely rigid with regard to

in-plane displacements so that any such movements (due, for example, to thermal elongation or

parasitic stressed skin action) must be accommodated primarily by ductility in the fastening system.

F

e1 e2

direction of span

eC e3

Figure 2: Test arrangement for pull-through resistance based on small-scale specimens at an end

support (ECCS & CIB, 2009).

eC e3

e4 e4

direction of span

Figure 3: Test arrangements for pull-through resistance with small-scale specimens at an intermediate

support (ECCS & CIB, 2009).

The tensile resistance of a fastening represents the minimum value of the pull-through resistance and

the pull-out resistance. The load-bearing capacity of the screw itself does normally not play an

important role in the resistance of the fastenings of sandwich panels. The Recommendations ECCS

& CIB (2009) deal with the pull-through resistance only. For determining the pull-out resistance

reference is made to ECCS (2008).

23

The distance between the end of the panel and the fastener has an influence on the tensile resistance

of the fastening. Therefore, two separate test arrangements have been developed to take into account

the different nature of the fastening at an end support and at an intermediate support as shown in Figs

2 and 3. Figure 4 shows the principal set-up of a test for the determination of the shear resistance of

a screw fastening. Since the influence of the external face decreases with the increasing thickness eC

of the core layer, the tests are to be performed with the largest envisaged panel thickness. As an

alternative, the sole direct load transmission between the internal face and the substructure may be

investigated.

eC e1

t1 F

u

Figure 4: Shear test assembly for screws passing through the panel (ECCS & CIB, 2009).

For the determination of the resistance of a screw to the deflection of the screw head, the shaft of the

fastener is subjected to a repeated deflection of „u‟, where u is the maximum lateral displacement of

the head. The deflection spectrum is as follows: 20.000 cycles with a deflection of 4/7 u, 2.000

cycles with a deflection of 6/7 u and 100 cycles with a deflection of u. This load spectrum is based

on the assumption of a service life of 50 years in a location in central Europe. During the test, the

screw shall not fail and, after cyclic loading, the screw has to achieve at least 80 % of the mean value

of the pull-out resistance without the repeated load.

2.2 Openings in sandwich panels

2.2.1 Background

In the majority of buildings, functional requirements necessitate openings in the wall and roof

cladding. The current state of the art with regard to openings within the sandwich panel wall and roof

construction generally requires the addition of reinforcement in the form of additional framing to

replace the load-bearing capacity which has been removed by the opening (Fig 5). Recent research

results indicate that this additional framing is not always needed. The state of the art report of the

Joint Committee will introduce the possibility of designing panels with openings without the need for

any additional strengthening.

2.2.2 Remaining resistance of panels with openings

The cross-section of the panel which remains after cutting an opening may be able to carry the

applied loads. In this case, the opening is classified as „small‟. The evaluation of the remaining

24

resistance is evidently the primary design task and this includes verification of the resistance of the

net section to the design bending moments and shear forces at the critical points in the vicinity of the

opening taking account of any stress concentrations.

A section A - A

framework A replacement with cross-

beams, e.g. C-sections

Figure 5: Openings requiring additional supporting members (ECCS & CIB, 2010).

remaining cross-section

gross cross-section

Figure 6: Remaining cross-section in the vicinity of an opening (ECCS & CIB, 2010).

The ECCS & CIB report (2010) gives design formulas to calculate the remaining resistance of a panel

with openings (Fig 6).

2.2.3 Activation of the load-bearing resistance of neighbouring panels

If the remaining cross-section of an individual panel is not able to carry the applied loads, the load-

bearing resistance of the neighbouring panels can be activated. Due to the difference of the stiffness

between the panels with and without openings, the whole or a part of the loads applied directly to the

panel with openings will be transferred via the longitudinal joints to the adjacent panels (opening „A‟

in Figure 7). The most severe case is the load transfer from a completely cut sandwich panel

25

A

(opening „B‟ in Figure 7). While in case A the panel might have sufficient capacity to withstand the

load, case B relies entirely on the load transfer. In both cases, the neighbouring panels will receive

additional loads which can be calculated on the basis of the compatibility of the deflections in the

longitudinal joints (Figure 8). These loads are transferred by these joints and therefore, the

assessment of the strength and stiffness of the joints is essential. Furthermore, the load transfer will

result in an eccentric line load applied to the adjacent panel without an opening thus activating its

torsional rigidity and causing additional shear stresses due to the torsional moment.

A

B

Figure 7: Small and full-width openings in sandwich panel walls (ECCS & CIB, 2010).

Figure 8: Line loads on the longitudinal joint between neighbouring sandwich panels with different

bending and shear stiffness (ECCS & CIB, 2010).

The intensity of the load, which can be transferred through the longitudinal joints to adjacent panels,

depends on the bending, shear and torsional rigidity of the complete panels and, in addition, on the

shear rigidity and shear resistance of the longitudinal joint. Simple framework-software can be used

to compute the internal forces and deflections. ECCS & CIB (2010) presents the basic principles of

the design procedure. A test set-up for the determination of the stiffness and resistance of the

longitudinal joints is presented as well as formulas to determine the required bending and torsional

stiffnesses of the panels.

26

2.2.4 Large-sized openings

In some cases load transfer from the panel with openings is not possible due to the limited shear

resistance of the longitudinal joint or the limited load-bearing resistance of the neighbouring

sandwich panels. An additional frame has then to be designed to carry the whole load of the sandwich

panel with openings. This frame may be placed within the longitudinal joints of the panel (Figure 9).

longitudinal beam

Fc

Fc

F1,c

Fc

cross-beam Fc

F1,c

Ft

Ft

longitudinal beam

Figure 9: Frame-construction within the panel using special aluminium profiles with web parts made

of plastic sections (ECCS & CIB, 2010).

3. The EASIE project

The Joint Committee is the birthplace of the European research Project “Ensuring Advancement in

Sandwich Construction through Innovation and Exploitation” (EASIE). The „EASIE‟ project

commenced in October 2008 and is funded to approximately 4 million Euro with financial support

from the European Community's Seventh Framework Programme FP7/ NMP2-SE-2008 under grant

agreement No 213302.

EASIE provides support to small and medium sized enterprises to develop the design, manufacturing

and use of products related to sandwich constructions. The Joint Committee first defined subjects and

tasks of the project reflecting the needs and interests of the industries were collected. In addition,

EASIE will also study subjects relevant to the updating of the product standard EN 14509. It

includes a number of innovative subjects such as the performance of sandwich panels subject to in-

plane shear and axial resistance.

The second goal of EASIE is to disseminate existing and new information for use in practice. This

will be carried out through seminars and the production of practical guidelines and e-learning

modules. The subject areas include the principles of the design and use of sandwich panels, fastening

and fixing, properties in extreme situations and the information given by the research results of

EASIE. Practical guidelines and seminars will help and broaden the correct and safe use of sandwich

panels in Europe and ICPC.

27

4. Future tasks of the Joint Committee

The earlier work of the Joint Committee has produced independent technical harmonised guidelines

and relevant background documents for the European product standard. The technology of sandwich

panels is still evolving and the Committee seems still to have a useful role as author of technical

guidelines based on new and reliable information without pressure from commercial or other

interests. Because of this continuous development, keeping the available documents up to date is an

important task of the Committee. Future tasks are likely to include the preparation of new guidelines

on the basis of the technical information emerging from projects such as EASIE. Finally, questions

arising from environmental issues in buildings and building products may open new technical tasks

requiring a measure of European collaboration.

References

CIB (1978) Recommendations for the structural design of lightweight sandwich panels. Rotterdam:

CIB - International Council for Research and Innovation in Building Construction, S56 Lightweight

Constructions.

International Organization for Standardization (1973) General principles for the verification of the

safety of structures. ISO 2394.

Davies J M et al. (2001) Lightweight sandwich construction, Oxford, Blackwell Science.

ECCS & CIB (2000) European Recommendations for Sandwich Panels – Part 1: Design.

Brussels/Rotterdam: ECCS - European Convention for constructional steelwork, TWG 7.9 Sandwich

panels and related structures & CIB - International Council for Research and Innovation in Building

Construction, W056 Lightweight Constructions. ECCS Publication No 115 and CIB Publication No.

257.

ECCS (2008) The Testing of Connections with Mechanical Fasteners in Steel Sheeting and Sections.

Brussels: ECCS - European Convention for constructional steelwork, TWG 7.10 Connections in

Cold-formed Steel Structures. ECCS Publication No 124.

ECCS & CIB (2009) Preliminary European Recommendations for Testing and Design of Fastenings

of Sandwich Panels. Brussels/Rotterdam: ECCS - European Convention for constructional steelwork,

TWG 7.9 Sandwich panels and related structures & CIB - International Council for Research and

Innovation in Building Construction, W056 Lightweight Constructions. ECCS Publication No 127

and CIB Publication No. 320.

ECCS & CIB (2010) State of the art report for Design of Sandwich Panels with Openings.

Brussels/Rotterdam: ECCS - European Convention for constructional steelwork, TWG 7.9 Sandwich

panels and related structures & CIB - International Council for Research and Innovation in Building

Construction, W056 Lightweight Constructions. (will be published in 2010?)

28

EN 14509 (2006). Self-supporting double skin metal faced insulating panels - Factory made products

- Specification. Brussels: CEN - Comité Européen de Normalisation.

Ensuring Advancement in Sandwich Construction through Innovation and Exploitation:

www.easie.eu

29

Methods to Measure the Durability of Structural Sandwich Panels

Hassinen, P.

Aalto University Department of Structural Engineering and Building Technology, Finland

(email: paavo.hassinen@tkk.fi)

Pfeiffer, L.

ThyssenKrupp Steel Europe AG Profit Center Color/Construction, Germany

(email: lars.pfeiffer@thyssenkrupp.com)

Abstract

Determination of the intended length of the service life is an increasingly important question and may

be soon a crucial parameter in the development and design of civil engineering structures. Length of

the service life may be limited by a reduction of the resistance or by an unsatisfactory function at the

serviceability state. To determine the service life, knowledge about the deterioration mechanism and

their influence on the structural behaviour is needed. In some cases, information about the

development of the mechanical and environmental loads during the service life is important, too.

European standard EN 14509 introduces experimental methods to measure the changes of the self-

supporting sandwich panels and further, to classify the panels to be fit for the use in external walls

and roofs. The methods are based on an indirect parameter by measuring the changes of the cross

panel tensile strength and not the properties, which are directly used in the analysis and in design.

Ageing of the cross panel tensile test specimens is based on accelerated ageing histories using a high

temperature and moisture. The method is practical and relatively easy to use, however, it does neither

tell about the changes caused by a real natural environment nor the real changes of the ultimate

resistance or the function at the serviceability limit state.

The contribution introduces the parameters which are essential in structural design and then, studies

the possibilities to measure the changes of the parameters in long-term use. The essential parameters

are the wrinkling strength of a face under compressive stresses and the shear strength of the core and

bond. These are followed by a number of other failure modes, such as the resistance at the supports

to pressure and suction loads. The information is based on tests with small-scale specimens, the

results of which will be compared to a limited number of results on full-scale specimens having been

in service and exposed to natural ageing conditions. The contribution concludes the promising

possibilities and draws outlines for development of methods for the use in practice.

This publication was prepared within the scope of the EASIE research project. The project is

sponsored in EU 7th framework programme.

Keywords: sandwich panel, resistance, long-term, durability, ageing, test

30

1. Introduction

1.1 Definition of a sandwich panel

Sandwich panels in the context of this report are composite structures, consisting of a minimum of

two deck layers and a core. The term “panel” is derived from the Latin, French and Dutch word

“paneel” and means “a flat piece of construction” (Brockhaus, 1991, a). This describes the general

shape of a sandwich panel. The word “sandwich” describes generally a combination of different

layers (Brockhaus, 1991, b). A typical sandwich layer in construction consists of three layers. A

rather thick rigid core material is laminated between two thin faces. There are many possible

combinations of facing and core materials, depending on the intended use of the sandwich panel. This

work, however, deals primarily with the most common form of a sandwich panel used currently in the

building construction: factory made engineered elements, consisting of thin metallic facings of a

thickness between 0.4 and 0.75 mm and rigid foam or inorganic wool core.

1.2 Load bearing behaviour of a sandwich panel

The individual layers of sandwich panels by themselves have almost no load-bearing capacity. In a

sandwich panel the layers no longer act separately but are connected in a process of adhesion. Only

this rigid connection makes the structure a sandwich panel and increases the load-bearing capacity

substantially.

While the metal sheet facing alone is almost without bending capacity as is the core by itself, the

connection of the three layers creates a completely new composite structure with widely enhanced

capacities. The composite sandwich panel possesses substantially higher load-bearing capacities with

regard to bending, shear and torsion impact than the sum of its individual components.

Compared to the metal faces, the core in a sandwich panel generally possesses very limited tension or

compression stiffness. The bending moment is therefore distributed to the two faces. For a single

span, three-layer sandwich panel loaded by a distributed load, this leads to compression and tension

in the two faces that are kept apart at a fixed distance by the core.

Based on the properties of the layers and the distribution of the internal stresses, the core carries

through almost all of the shear force. Additionally, the core provides foundations to the faces of the

sandwich panel. Because of that, the properties of the core have a significant influence on the over all

performance of the sandwich panel and need to be carefully evaluated in sandwich panel design.

Losses in the stiffness and strength of the core have immediate influences on the performance of the

whole panel. Such losses can, for example, be caused by durability related degradation.

31

1.3 Durability of construction products

A survey by Sarja (2002) on the typical share of civil engineering products for European countries

shows that:

Civil engineering products represent about 70 to 80% of the national assets.

The energy use for these products during production and maintenance is about 40% of the total

national energy consumption.

Civil engineering products produce about 35% of the total waste.

These figures emphasise the enormous economical importance of civil engineering products for the

societies. Due to the importance and costs of these products, it becomes obvious that there are high

demands to the service ability during the life span of the investment.

Economical effects are dominated by the fact that buildings generally require large investments. The

expectations in the investment can only be fulfilled if the building keeps its functionality over a

certain time period. The estimated time period depends on the type of building. For an industrial

building like a warehouse, for example, it is generally assumed that the building is designed for a

service period of 25 to 30 years. After this period a substantial refurbishment of the building is

generally required.

The most important technological influence for a civil engineering product is the guarantee of the

safe use over a time. Building materials need to be designed in a way, which guarantees an adequate

structural performance. If a structural building material cannot resist against the received load

impacts, the building is damaged or may even collapse. In a worst case scenario, this may lead to a

loss of lives and properties. Therefore, the main task in the regulation is to ensure an adequate

structural performance of the building materials throughout the whole life cycle of the structure.

Legally binding design procedures, which are very much harmonized across Europe, are the base to

guarantee a safe structural performance. For sandwich structures, such procedures are defined in the

European standard on sandwich panels EN 14509, where also requirements towards the durability of

the sandwich panels together with experimental evaluation procedures are described.

In practice, requirements to the fitness for use and to the withstanding the extreme actions during the

use are set down to a structure. The previous requirement defines the serviceability limit state and the

latter one the ultimate limit state. The structure shall maintain the properties in its environment

during its design working life, which defines the requirements on the durability of the structure. The

durability requirements may be achieved by choosing correct materials and manufacturing

techniques. As a result, the deterioration of the structure will stay small enough or the function and

resistance will be guaranteed by regular inspections and refurbishment. In constructions, the first

principle is used in the major cases of structures.

The available durability tests give information about the changes of the strength on a macroscopic

level without explaining the physical and chemical mechanisms affecting the deterioration of the

material. It is believed, that the durability test is able to classify if the product is to fit for the purpose

32

or not. The durability test methods and the loading histories are therefore strongly dependent on the

application.

2. Durability of sandwich panels

The European standard, EN 14 509, makes in chapter 5.2.3 a statement concerning durability;

“Durability and other long term effects.” Detailed information on testing scenarios are written down

in the normative annex B, “Durability testing method for sandwich panels.”

Currently, the durability performance is only evaluated qualitatively. This means that the loss in cross

panel tensile strength under an artificial, durability accelerating climate must stay within certain

boundaries over a time. The cross panel tensile strength is seen as an indicator for over all panel

performance. Depending on the type of degradation pattern (Fig 1), a core material is either fit for the

application or not. The boundaries chosen for an acceptable loss of the strength are, however, lacking

the scientific foundation. Even a loss of 60% of the initial tensile strength can be acceptable in

accordance with the standard. In particular, such a loss in the tensile strength is allowed without any

impact on other relevant design parameters. Because it is not possible to determine the remaining

wrinkling stress after the artificial ageing of the tensile strength, a new small-scale test to study the

development of the wrinkling stress is developed.

Figure 1. Possible degradation patterns for tensile strength (Davies et al.).

3. Wrinkling stress of sandwich panels

3.1 Wrinkling failure modes

The wrinkling of the face under compression is a typical and very important failure mode for

sandwich panels with membrane-like faces (Fig. 2). As stated before, in a loaded sandwich panel, one

33

u w

face is in tension while the other one is loaded by compressive stresses. A membrane under

compression has, on its own, no stability, but it fails immediately through buckling. In a sandwich

panel, the core provides a lateral support for the thin face and prevents the early buckling failure. The

lateral support is activated when the face deforms in a wave-like pattern and induces stresses in the

core material. The loss of the tensile strength of the core or the bond results in an immediate

wrinkling failure of the face. Wrinkling is one of the most critical failure modes in a sandwich panel

construction and it very often determines the ultimate limit state.

Figure 2. a) Wrinkling failure in a span and b) on an intermediate support.

3.2 Determination of wrinkling stress

The wrinkling stress is defined as the ultimate stress, a face can take in compression before it fails in

buckling. The compression in the panel face is, in practice, induced by a bending moment that can, be

caused by wind loads or temperature loads in multi span panels. Determination of the wrinkling stress

of a sandwich panel generally requires full scale testing. The full scale test may be carried out by

subjecting a simply supported beam to four line loads.

In a test setup according to figure 3 the ultimate bending moment (Mu) and the wrinkling stress ( w)

for a lightly profiled panels are given by:

M FBu L M u

8 and eC A1 (1), (2)

where Fu

L

is ultimate load including self weight of panel and loading equipment

span of the test specimen

eC

A1

distance between the centroids of the faces

cross-section area of the face in compression

In order to be able to determine wrinkling stress after an artificial ageing, it is necessary to develop a

small scale wrinkling test set-up which allows the determination of the wrinkling stress or more

precisely the time dependent change of the wrinkling stress. The small sample can then fit into a

climate chamber and tests on the aged samples can be performed.

3.3 Small-scale wrinkling test

The mechanical parameters necessary for the design may be obtained on the basis of the small scale

testing. An exception is the wrinkling strength. For a durability research this poses the problem that it

is not possible to determine the wrinkling strength after an artificial ageing. Therefore, a small scale

test setup determining wrinkling strength has to be developed. When developing a small scale

34

wrinkling test for sandwich panels, it is the general idea to no longer induce compressive stresses in

the face through a bending moment but to cut out a small panel area and expose it to a direct axial

load in the direction of the plane. Baehre (1988) and Pfeiffer (2000) have previously tried to establish

small scale wrinkling tests for the design of sandwich panels. However, it was not possible to

determine the exact wrinkling capacity of a sandwich panel through small scale testing. The main

problem in the earlier work was in finding a way to introduce the in-plane loads without damaging

the thin face locally, a problem that does not occur in a full scale test. Special load application

devices help to overcome the problem. During the test series it has been found that the full-scale

bending test loads the sandwich panel twofold. In the upper face, the in-plane bending stress meets

the deformation caused by the line loads, which evoke the actual bending moment and compressive

stresses in the face. For a direct comparison between six point bending test and a small scale

wrinkling test, this load combination must not be neglected. The results gathered in this project

suggest that the research conducted by Baehre (1988) would have had a positive outcome if the local

deformations were taken into consideration. With the results on hand, it was not possible to find a

small scale test defining the wrinkling strength in the same way as a full scale test does.

3.4 New test set-up

Figure 3. Arrangements of a full-scale and a small-scale wrinkling test with descriptions of the

distribution of the bending moment and compressive stress.

The new test combines the two previously used test methods. Instead of causing compression in the

upper face through a bending moment, the small scale test loads directly the face in a vertical, in-

plane direction (Fig. 3). While the authors in their earlier work used a wooden T-shaped load

application device, which was glued to a rectangular specimen with completely flat faces causing

compression in one face, Baehre (1988) used a micro-profiled specimen with T-shaped load

application devices made of steel. The load application device was not glued to the specimen but

rather clipped onto it. In both cases the T-shaped device was chosen to gradually introduce the load to

the specimen without causing local failures at the point of the loading (Fig. 4). Baehre failed to find a

35

small-scale test, giving similar results as obtained in a full-scale bending test. The reasons for this

were the local deformations in the compressed face under the loading points in the full scale test.

Such deformations weaken the observed wrinkling strength as they cause an additional local moment

and additional stresses in the face. For a small scale test comparable to a full scale test this effect

must be accounted for.

Figure 4. Test set-up of a small-scale wrinkling test from the left to right; Baehre (1988), Pfeiffer

(2000) and Pfeiffer (2004).

Figure 5. Introduction of the load in a small-scale wrinkling test. For illustration purposes, the

specimen is separated from the load introduction device. The total load (F) is transferred to the

specimen directly (Fdirect) and through the shear force (Fshear).

For this work, a combination of the two previously described small-scale tests, determining the

wrinkling stress, was chosen. The profiled specimen from Baehre was enlarged to a bone-shaped

sample, again using T-shaped load application devices at both ends of the specimen. This time the

application devices were both glued and clamped to the samples. Only the combination of gluing and

36

clamping together with the bone shape hinders the specimen from failing early through local

crushing. Gluing the load application device to the samples allows the load to be introduced to the

panel faces over a larger surface area, reducing concentrated load effects. The load is introduced

through a shear force in the glue area and directly to the top and bottom ends, where the specimen

touches the T-shaped device (Fig 5).

To study the effect of the local deformation at the load introduction areas in the full scale test,

samples with and without local deformations were studied. A three step approach was taken. First,

samples with completely undisturbed faces were tested, giving the maximum wrinkling stress that can

be obtained. In a second step, the faces were dented prior to executing the small scale test. The depth

of the imprinted deformations was of the same dimensions as observed in the full scale test. In a third

step, the effect was simulated through a bi-axial loading in the small scale test. An additional load

device was clamped to the specimen loading it in a perpendicular direction. All obtained results were

compared to the results of the full scale bending test. The last procedure resulted in the most accurate

results.

3.5 Test set-up based on rectangular specimen

Preparing of the bone-shaped specimens requires careful machining of the two metal sheets and the

core. To simplify the preparation work, a test set-up has been developed, which base on a rectangular

specimen. To avoid the failure at the end and to create a wrinkling failure in the mid-part of the

specimen, the specimen is loaded by an additional transversal load (Fig. 6).

a) b) c) d)

Figure 6. a) Rectangular small-scale test specimen, b) initial displacement caused by the transversal

load (initial deformation equal to the face thickness, w0 tF), c) displaced mesh and d) displacement

caused by the transversal and axial load (wu 3 tF).

37

ax

ial st

ress

in

rel

ati

on

t

wri

nk

lin

g s

tres

s

Transversal load causes an imperfection in the face and compressive stresses in the core layer. The

level of the transversal load has to be high enough to cause an initial imperfection to the wrinkling

failure and on the other hand, lower than the elastic limit of the core layer (Fig. 7). Experiments on

polyurethane-foam and mineral wool cored specimens have shown the test set-up to give acceptable

modes of wrinkling failures.

1

0,8

0,6

0,4

0,2

Fdef.1

Fdef.2

Fdef.3

Fdef.4

0

0 0,0005 0,001 0,0015 0,002 0,0025 0,003

axial compressive deformation

Figure 7. Influence of the transversal load on the stress-deformation behaviour of the face in the test

described in Fig. 6. Transversal loads Fdef.i cause the transverse stress of 83, 63, 42 and 21 kN/m2

in

the core. The axial stress of the face is compared to an analytically determined ideal wrinkling stress

of 125 N/mm2

in the case.

Small-scale specimens will be exposed to an accelerated influence of the environmental actions.

Because of the small size of the specimen, it is possible to create a large quantity of reliable

information about the changes of the wrinkling stress in function of time and environmental loading

history. The data will make the basis to the development of the design models in order to take into

account the possible deterioration of the strength already in the design phase.

3.6 Comparing small-scale and full-scale tests

In order to calibrate the test setup, a test series, comparing the large-scale and small-scale test

arrangements of the wrinkling stress was undertaken. All specimens in this series were taken from the

same production batch. The full-scale panels were first tested as simply supported beams as described

previously. The test panels had a PUR core between the two steel faces. The wrinkling failure

obtained in the full scale tests resulted in an average wrinkling stress of 180.94 N/mm² (individual

test results between min. 180.75 N/mm² and max. 181.27 N/mm²). The wrinkling failure, obtained

from the small scale test without the additional effect of the transversal load, averaged to 228.73

N/mm² (individual test results between min. 223.06 N/mm² and max. 234.76 N/mm²).

38

Figure 8. Local deformation in the loading point causing an early failure in the six point bending test.

The discontinuity (d) occurs under all loading points but interacts with maximum bending moment at

the indicated point.

Comparing the results given by the two methods, a difference of 26.4 % can be found. This difference

is not acceptable if the determination of the accurate wrinkling stress is required. The reason for the

observed difference lies in the test setup of the full scale bending test. In a six point bending test the

load is transferred to the panel in four line loads (Fig. 8). At the point of the load introduction, this

results in deformations of the face causing a failure at the inner point of the loading beam, where

maximum local deformations and the maximum bending moment interact. In order to achieve the

same results from the small scale testing as determined through full scale testing, this deformation

effect needs to be taken into account. An additional loading device generating a transverse load, was

introduced in the test. This additional device loads the panel perpendicular to the faces, causing the

same local deformation as obtained in the full scale test. The load is applied through two steel

profiles, having the same dimensions as the ones used in the full scale test. The loading is controlled

with the help of two screws and measured in a load cell. At an average of 178.79 N/mm² (individual

test results between min. 177.07 N/mm² and max. 180.50 N/mm²) the results obtained from the setup

show good correlation with the results of the full scale bending test. When comparing the average

results, 98.8 % of the strength determined in the full scale bending tests could be reached.

The results indicate clearly that it is possible to use the small scale testing to determine the accurate

wrinkling strength of a sandwich panel. It is, however, necessary to simulate the full-scale test

accurately and to take into account even the small details, like the local deformations at the load

introduction points.

When lacking knowledge of the maximum load obtained in the full scale test (Fmax), which is the case

when no full scale test has been carried out on a particular panel, it is necessary to gradually increase

39

the deformation load relative to the load parallel to the face. The load needed on the deformation

device can be calculated on the basis that the stress parallel to the face shall be equal in the both tests.

This means, that a certain load in the small scale test, corresponds to a certain load in the full scale

test, both causing the same stresses parallel to the surface. At the same time this load determines the

load on the additional load device and can be calculated as described in the following:

The membrane stress in the flat face in the six point bending test can be written by

F6 P L (3) f 6 P

8eC Af 16 P

where σf6P is the stress in face in 6 point bending test

F6P total load on panel in 6 point bending test

Af16P: cross sectional area of the face in compression in 6 point bending test

The corresponding stress in the small scale test is given by

FSC

(4) fSC A f 1SC

where σfSC is the stress in the face in the small scale wrinkling test

FSC total load parallel to the face in the small scale wrinkling test

Af1SC cross sectional area of the compressed face in a small-scale test

A simulation of the full scale test through the small scale test requires an equal stress parallel to the

face (wrinkling stress). Equating the equation 3 and 4 gives the load F6P

F 8 FSC eC b6 P

which can also be written as F6 P 2 FSC eC b6 P

(5)

6 P L

bSC 4 L bSC

in which tf1 net thickness of the face

b6P width of the panel in the 6 point bending test

bSC width of the sample in the small-scale wrinkling test

Local stresses in perpendicular direction to the face in the load introduction point are

def .6P

F6P

4 b6P B

full-scale test and (6)

def .SC

Fdef

bSC B

small-scale test (7)

where B is the width of load application device in the set-ups

40

For comparison of the two tests, the local transverse stresses need to be equal, which results in

Fdef

F6 P bSC

4b6 P

(8)

Combining the equations 8 and 5 leads to

Fdef 2 FSC eC b6 P bSC 2 FSC eC

(9)

L bSC b6 P L

The load in the additional deformation device can now be calculated, depending on the axial loading

of the sample, the distance between the centroids of the two panel faces, and the span length of the

static system in the six point bending test.

4. Conclusions

The results presented in the paper indicate that the small-scale wrinkling test can be used in the

evaluation of the wrinkling stress. The test requires extensive preparation work such as bone shape or

rectangular cutting and gluing and fixing of the load application devices. Laboratories dealing with

sandwich panel technology on a regular base can adopt the test with some additional work. For

further work on the durability related degradation of the wrinkling stress, it is sufficient to consider

only the durability related changes and thus the relative values. In future research work the test

arrangements without the additional deformation device will be studied. Also a simplified rectangular

wrinkling test set-up will be developed further.

The shear and compression strength of the core is determined in most cases on the basis of the small-

scale tests. Thus, the durability of the shear and compressive strength can be studied by exposing the

specimens to accelerated deterioration processes in a climate chamber. These studies will contribute

further information to the evaluation of the durability of the structural sandwich panels.

References

Sarja A (2002) „Neue Anforderung an die Dauerhaftigkeit von Konstruktionen“ (New requirements

towards durability of constructions), Der Prüfingenieur 20: 46-53.

EN 14 509 “Self-supporting double skin metal faced insulating sandwich panels – Factory made

products – Specification”

Davies JM et al. (2001) Lightweight Sandwich Construction Blackwell Science, Oxford

Baehre R (1988) Experimentelle Ermittlung der aufnehmbaren Knitterspannung von ebenen und

leicht profilierten Stahldeckschichten, Research Report, University of Karlsruhe

41

Pfeiffer L (2000) Der Einfluss der Zugfestigkeit auf die Knitterspannung bei Sandwichelementen,

Final Diploma Thesis, Fachhochschule Mainz (D) and Häme Polytechnic (FIN)

Pfeiffer L (2005) Durability Assessment of Sandwich Panel Construction, Doctoral Thesis,

University of Surrey (GB) and Fachhochschule Mainz (D)

EASIE Ensuring Advancement in Sandwich construction through Innovation and Exploitation, EU 7

th

FP, 2008 to 2011, www.easie.eu

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