International Symposium on Tubular Structures, Delft, The Netherlands, 26-28 June 1991: Addendum

International Symposium

On Tubular Structures

Delft, The Netherlands

June 26,27

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28, 1991

Department of Steel and Timber Structures

Faculty of Civil Engineering

Delft University of Technology

Delft, The Netherlands

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International Symposium

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June 26,27

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Department of Steel and Timber Structures

FacuIty of Civil Engineering

Delft University of Technology

Delft, The Netherlands

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LOAD TEST ON A SINGLE LAYER BRACED DOME FOR THE PRAGATI MAIDAN,

NEW

DEHLI, INDIA

G.S. Ramaswamy, Mick Eekhout and Jan Faber

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ABSTRACT

The paper reports the analysis of observations made and conclusions

drawn from a laad test carried out on a full scale prototype dame

typical of 16 dames designed by the authors for an Exhibition Hall,

now under construction, at the Pragati Maidan Exhibition Grounds at

New Delhi for the Trade Fair Authority of India.

The Exhibition structure comprises a cluster of 16 single layer

spherical dames on square ground plans of 22.10 m x 22.10 m with a

rise of 8.599 m,

supported on four boundary arches with a rise of

4.936 m.

The dame members are of steel tubes with a yield strength of

450 Njmm2

_{joined together by specially designed node connectors. The}

arches, also fabricated out of tubes, are of welded construction.

The contractcalled for a laad test on a full scale prototype under

the full design laad of 170 kgjm

2_•

_{The loads on the test dame were}

applied at the nodes through loading rods carrying specially made

steel cradles stacked with pre-weighed bricks. Deflexion observations

were made at a few selected points in the dame and arches by means of

specially fabricated devices. The design laad of 170 kgjm

2

_{was applied}

in 5 equal increments and the full design laad was kept susta

.

ined for

nearly 36 hours befare unloading began in 5 equal decrements.

The authorities concerned had demanded such a test on a full scale

prototype, by way of abundant caution, because of the uncertainties

inherent in analytical calculations and the well-known vulnerability

of single layer dames to snap-through buckling. The dame was idealized

as pin-jointed for purposes of stress analysis carried out on a

computer, using SAP 90 software. The bending moments caused by purlins

transferring loads between nodes were, however, taken into account.

,

A pilot test carried out on the bolted dame revealed the need for

additional rigidity at the nodes. Consequently, the junction between

the dame members and the arms of the connectors were reinforced by

welding and the dame retested.

The analytically predicted deflections were compared with the measured

deflexions and they were in reasonable agreement.

The test was wholly successful. The dame exhibited almast total

elastic recovery and the measured deflections were very much less than

those prescribed by the acceptance authority. The measured horizontal

deflexions were negligible. The dame has been pronounced to have

successfully met the acceptance criteria.

INTRODUCTION

A cluster of 16 single layer braced domes on square ground plans of

22.1 m x 22.1 m with a rise of 8.599 m,

supported on four baundary

arches with a rise of 4.936 m resting on reinforced concrete columns

of varying heights, is now under construction to house an Exhibition

Hall at Pragati Maidan Exhibition Grounds at New Delhi. The braced

dome and the arches are specified to be of tubular construction. The

domes are to be clad by 5 cm thick wood wool slabs resting on purlins.

The slabs are to be topped by 4 cm thick mesh-reinforced concrete. The

live load specified is 50 kg/m

2

•

As a part of the contract, the

acceptance authority has specified a load test on a full scale

prototype dome under the total design load of 170 kg/m

2

•

This is a

rather unusual requirement. Such a test was demanded, possibly because

single layer domes are known to be prone to snap-through buckling.

The authors were cal led in as consultants af ter an earlier dome,

designed by a different consultant, collapsed during the load test.

The authors, in redesigning the dome,

reduced its radius from 22 to

18.5 mand the number of horizontal rings from 9 to 6 to ensure

stability. The authors had, in fact, recommended that the radius be

reduced to 14.5 mand the rings to 4. Architectural constraints ruled

out such a radical change. The geometry of the redesigned dome may be

seen in fig. 1.

STRUCTURAL ANALYSIS AND DESIGN

For purposes of analysis, the dome was idealized as pinjointed.

structural analysis was carried out on a computer utilizing SAP 90

sofware. Bending moments caused by purlins transferring loads, between

nodes, were taken into account and a second order analysis was carried

out as laid down in reference [1]. The dome members are of tubes of

450 N/mm

2

_{yield strenght and they are flattened at the ends and}

connected to the arms of specially designed connectors by means of

shear balts. (fig.2)

In the design of the dome members,

the effects of 20 mm column

deflexions in the x and y horizontal directions and a rise in

temperature of 30· were taken into account. The arches, being of

welded construction, were analyzed as stiff-jointed. Local analysis

for reinforcing the junction between tubes was carried out in

accordance with the recommendations found in references [2] and [3].

PILOT TEST

A pilot test was carried out on the dome to assess its behaviour by

loading it to 60 % of the design load. For this purpose, the dome was

erected as follows:

(a) The arches were first erected over the columns.

(b) Using a central derrick, the dome was erected from top downwards,

adding one ring at a time.

(c) The pendentives were finally forced into position to rest on the

arches.

On

loading the dome to 60 % of the design load, unacceptably large

deflexions and local dimples developed in the region of the

pendenti-ves. These were most pronounced at measuring stations 10, 11, 12 and

13 (fig. 3). These were clearly danger signals pointing to imminent

snap-through. The pilot test was therefore halted to make a diagnosis

of the contributing causes and to work out the corrective action to be

taken. These may be summed up as follows:

(a) The node connectors with shear balts obviously do not provide

adequate rigidity against snap-through buckling, under the local

Indian conditions of quite large fabricication tolerances. This is

grey area on which the published literature is sparce [4], [5] and

codes of practice provide little or no guidance. Wright's observation

[5] that 'without bending strength an rigidity at nodes or joints,

snap buckling willoccur at small loads, but with effective joints the

problem disappears' was recalled and the authors decided to explore

ways and means of enhancing the rigidity at the nodes.

(b) The erection sequence followed of working from the top toward the

battom, adding one ring at a time, made it extremely difficult to

measure the initial geometry accurately at a height over the ground.

Moreover, when the pendentives were forced into position over the

arches, dimples had developed in the region of the pendentives.

Although the contractor removed them before the test started, they

reappeared during the test at a load of 60

%

of the design load.

Clearly, the ere ct ion sequence had introduced inadvertent errors ln

the initial geometry.

It is well known that such initial errors ln

geometry can be a contributing cause in triggering snap-through. It

was therefore decided to alter the erection sequence.

MODIFICATIONS IN DESIGN AND ERECTION SEQUENCE

The following modifications in design and erection sequence were made

in the light of the experience gained during the pilot Test:

Changes in Design

(1) An additional eye ring of 2,4 m diameter was provided to improve

rigidity. It was connected to the previous eye of 5 m diameter by a

triangulated network of tubes of 63.5 mm diameter and 2.9 mm

thickness. Connectors, reinforced with welding, were provided at the

nodes of the network.

(2)

In some members,

the two nos. of M12 balts of 8.8 quality were

replaced by 3 nos. of M12 bolts of 10.9 quality. As this was not found

to be reliable enough due to the poor deformation properties of the

high yield tubes which did not give a proper flat friction surface,

another solution was chosen: the junction of the flattened tubes with

the connector arms were reinforced by welding (fig. 4)

(3) In some other members, the 2 nos. of M12 balts of 8.8 quality were

replaced by 2 nos. of M12 bolts of 10.9 quality.Later also these node

connectors were additionally reinforced by welding.

(4) The tors ion rods provided in the arches were replaced by tubes to

impart additional rigidity.

Before these changes were decided upon,

the alternative of upgrading

the number and quality of the balts as described under (2) and (3)

abave and pretensioning them in addition was studied to obviate

welding. Taking the ground realities into account, welding instead of

pretensioning high yield stress balts was ultimately chosen as the

fail-safe solution.

Changes in Erection Seguence

To overcome the drawbacks observed during the pilot Test, the erection

sequence was modified as follows:

(a) The pendentives were first positioned on the arches by

balting

and subsequently firmly fixed by welding.

(b) The dome up to ring 5, including the previous eye of 5 m diameter,

was assembled over the arches and pendentives on the ground,

permit-ting the accurate checking of the initial geometry.

(c) The dome and arches were next positioned on the columns and the

remaining rings were added from the battom upward using the central

derrick (photo 1).

Af ter the dome was reerected, the load test was carried out as

described in the next section.

DESCRIPTION OF THE LOAD TEST

Dome and Arches

The test dome and arches were an exact replica of the actual

structures and the details of the support of the arches on the

reinforced columns were also faithfully reproduced.

It was not,

however,

practicable to replicate the columns as they are, because

some of them are 16 m high. To facilitate testing and observation, the

column heights were scaled down to 1 m abave their foundations. The

dome and arches were erected for the test as already described.

Loading Method and Seguence

The loads were applied at the nodes by means of loading rods which

carried specially made cradles on which preweighted bricks were

stacked in a symmetrical sequence (photo 2). The load of 170 kg/m

2

was

applied in five equal increments. The full design load was kept

sustained for nearly 36 hours, before unloading began in five equal

decrements.

.

Deflexion Observations

The deflexion measurements were taken at the stations shown in fig.3.

These measurements were between nodes to avoid interference with the

loading rods located at the nodes. The vertical deflexions were

measured by means of specially fabricated devices involving vertical

rods welded to the dome members at the points where observations are

to be made.

These telescoped into vertical tubes, carrying scales,

firmly fixed to the ground (photo 3). The measured deflexions were

cross-checked by means of simple water levels.

ANALYTICAL PREDITION OF DEFLEXIONS

It was considered desirable to compute deflexions at the observation

stations analytically sa that they may be compared with measured

values. Because the columns had been scaled down in height, the

analysis was made assuming the arches were regarded as truly

stiff-jointed. The assumption made regarding the end conditions of the dame

members is shown in fig.5 which is selfexplanatory.

OBSERVATIONS AND CONCLUSIONS

(1) No member of connector failed or showed at any signs of distress

during the test.

(2) The maximum measured deflexion of 17 mm was very much less than

the SPAN/360

=

62.5 cm.

(3) The residual deflexions measured on completely unloading the dame

were of insignificant magnitude, indicating almast total elastic

recovery.

(4) The measured horizontal deflexions were negligible.

(5) There was hardly any increase in the deflexions under sustained

laading.

(6) The measured and computed deflexions, shown compared in Tabel 1,

are in reasonable agreement. As is only to be expected, the

measured deflexions are slightly higher than computed deflextions

for the following reasans:

(a) The columns are not infinitely stiff as assumed.

(b) Nonlinear effects have not been considered in the analysis.

(c) The initial geometry may not be 100

%

accurate.

(d) The stiffness at the nodes may not exactly be as modelled.

(e) Small ground settlements underneath the measuring devices

cannot be ruled out.

The structure was consequently declared to have passed the laad test

by meeting all the acceptance criteria.

ACKNOWLEDGEMENTS

The authors express their gratitude to the client, The Trade Fair

Authority of Indiai RITES,

the supervising agencYi

the National

Building construction corporation and Nagarjuna Steel Ltd. which are

respectively the general and Sub Contractors for the Project. They

also record their thanks to stein Bhalla and Doshi, the architects who

made crucial contributions to the success of the Project. To Mr P.K.

Madhav,

President and

Mr

V.

Rama Rathnam and Mr K.

Venkateswarlu,

General Managers of Nagarjuna Steel Ltd. and Mr V.R.

Rajan, Vice

President, Nagarjuna Steel Ltd. special appreciation is due for the

many courtesies and facilities extended to the authors for carrying

out and documenting the Load Test. The authors acknowledge the

contribution of Ir. Jan Faber who was responsible for the analysis and

design of the domes and arches.

REFERENCES

[1] Regulations for the calculation of building structures, Design of

steel structures, NEN 3851

[2] Eurocode No.3:

"Design of Steel Structures Part 1, General Rules

and Rules for Buildings", Vol. 2 Annex K

[3] Wardenier, J.; "Hallaw Sectian Jaints", Delft University

Press,

1982.

[4] Saare, M.V.: "Investigatian af the Collapse af large

Span Braced

Dame",

Chapter 5 af "Analysis, Design and Canstructian af braces

Dames" edited by Z.S. Makawski, Granada, 1984, pp. 161

ta

171-[5] Wright, W.T.:

"Membrane Farces and Buckling in Reticulated

Shells, J. Struct. Div. ASCE 91 (ST 5), 1965, pp 173 ta

211.

Measuring station 1,2 3,7 photol. Computed Va1ues

Dome considered Dome considered

pin-jointed rigid-jointed 12.8 13.5 14.5 12.3 Average measured 17.0 16.5 4,5,6,8 8.9 7.0 10.0 9 5.5 3.7 3.5 10, 11, 12 1.9 1.4 4.2~ 14,15 17.5 14.3 14.0 --- --- ---H1,H2,H3,H4 2.3 1.6 2.2

~ Deflexion at station 11 was not taken into account because it was

excessive and amounted to 10 mmo

TABLE 1 Comoarison of comouted and measured deflexions

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- - -

-INVESTIGA TIONS ON THE FA TIGUE BEHAVIOR OF HOLLOW SECTION JOINTS SUBJECTED TO SPECTRUM LOADING (FATIGUE LIFE)

Ö. Bucak and F. Mang

Versuchsanstalt fûr Stahl, Holz und Steine Universität Karlsruhe

1. General

The object of the use of realistic load spectra is to make a more economical construction possible, and to all ow an accurate safety assessment of such structures. The big number of constant amplitude (CA) tests with different notch cases andjoints made ofhollow sections has been made available to the user as S-N-line catalogue [2] some years ago. In the following, the results ofthe fatigue investigations on hollow sectionjoints are announced and the application of the Miner-rule for variabIe load is shown, as it is required in Eurocode 3 [1] for considering the

variabIe load.

2. Testing Program

Systematic tests on hollow sectionjoints subjected to spectrum loading have been carried out at first with X- and K-type test specimens (fig. land table 1) under axialload applying load spectra known from literature [4, 7,8,11 etc.).

The basis for this work are fatigue tests carried out under constant amplitude load (S-N-lines). Based on the CA-tests, block tests were performed under the load sequence P = 1/3 and P = 2/3

for cranes and cranerailways which are known from literature [7, 8, 11 etc.) ..

By means ofthe load spectra "S-III Laplace" [7], North Sea" [7]and "North Sea Sequence" [12], aconnection to the European Offshore Research was established. Two further load spectra, one structure (mast, 200 m high) subjected to wind, the other one to the Gulf of Mexico completed the testing program.

Fig. 2 shows a compilation ofload spectra applied. The test values under load spectra have been compared with those ofthe Wöhler curves (CA) and the applicability ofthe Miner-rule has been checked.

Endurance tests will be evaluated by means ofthe Miner-rule whose accuracy and applicability will be examined in the course ofthe present investigations with regard to the treated specimen form and the load spectra applied.

The investigations ofCHS X-joints we re carried out with test pieces ofthree different dimensions (2 diameter ratios -d/dO = 0.5 and 0.8; and two wall thickness ratios - tolt = 1.4 and 2.0). Using these specimens, the dependenee of su eh hollow section joints on the existing geometrie parameter ratios could be considered. In addition, the given parameter ratios lead to three different types offailure:

a) Joint failure in chord shear (see fig. 3)

b) Crack starting from the weId toe ofthe web me mb er

c) Crack starting from the inside surf ace ofthe tube in the non-welded area ofthe chord

A detailed report has been given in [4, 5 and 6]. In the following, further investigations will be presented. The failure mode for rectangular hollow section (RHS) joints is identical to th at of

--

- - -

-

- -

-circular hollow section (CHS) joints. Figure 4 shows the test specimen with the geometrie parameter b/bo = 0.5 and tolt = 1.0 after the failure.

3. Test Results

From figure 5, the result ofthe investigations on X-type CHS-joints, and from figures 6 and 7, the results ofX-type RHS-joints can be seen. The stress range (SR = 0') is plotted in the vertical axis of these diagrams. The test results under load sequence have been plotted on the level of the maximum stress range in a colleetive. The tests were perfonnedwith the same testing machines, under the same test conditions without corrosion impact.

The detennination ofthe lines for different load collectives has the same slope as the S-N-lines for constant amplitude tests. Af ter completion ofthe first tests under the collectives P = 2/3,

Laplace Sm and P = 1/3, we found out that the slope ofeach S-N-line is the same. Therefore, we

decided to carry out the tests under load collectives on one level, and to draw a fatigue life line with the same slope as that used under constant amplitude by means of the mean value of the load cycle numbers ofthe test points. From tests on small specimens and L-type specimens tests it is known that the slope can become smaller, e.g. bigger values for m. In favor of considering the scatters, this positive phenomenon has been ignored.

An important question to be answered concerns the size ofthe crack, when ajoint type loses its serviceability. Since the de fini ti on ofthe bearing capacity ofthe applied testing load proved to be only a bulk value, measurements of the crack growth were conducted. As a criterion for disconnecting the testing machine, an elongation ofthe test pieces from 0.1 up to 3 mm, divided into various stages, was fIXed. The limitation of the maximum elongation of the test pieces 10 3 mm proved to be sufficient, since the crack reached over half ofthe width ofthe web chord when using this value. Nevertheless, even in this failure condition, the test piece could still bear the maximum testing force.

The EGKS-standards [7] indicate four different load cycle number or failure criteria for such investigations.

Ni Load cycle by a strain reduction of15% (e -15% = 0,85 e) close to the first crack. N2 Load cycle when the first discernible crack occurs.

N3 Load cycle through wall cracking (crack goes through the whole wall thickness. N4 Load cycle at the end ofthe test.

where the load cycles Ni and N2 only slightly differ from each other.

Before defining the cut-off criteria, the relation between the load cycle numbers have been detennined by means of single tests. On ofthe test results are given in figure 8.

In order to clarify the effect ofthe shape influence on the service fatigue strengths, besides X-type joints, K-type joints made of rectangular hollow sections were also tested under identicalloading collectives up 10 now according to fig. 2. A test rig, which has been used for previous investigations of K-joints, was applied in order to gain the truss forces as they occur for lattice structures.

The results of the investigations on the fatigue strength of K-type truss joints are presented in figure 9. Figure 10 shows one ofthe specimens af ter failure. The type offailures is identical 10 that

ofX-typejoints according to figures 3 and 4.

- - - -

-4. Proposal for the utilization of test results

Tbrough the comparison of stress ranges for a given load cycle, increase factors are obtained from the test values for the stress range in dependence on the collectives.

They are for the collectives 2/3

1/3

60

Collective tJOWöhler North Sea LBF North Sea WG3 Wind

5. Application ofthe Palmgren-Miner-Rule

= ) = ) =) =) = ) fE = 1.3

tE

= 1.8 fE= 3.0 f_E = 3.4 f_E= 3.0

Most ofthe new design rules recommend the application ofthe Miner-rule for the consideration of varying load (load sequence), e.g. for the proof of fatigue strength. For this reason, the meaningfulness ofthe Miner-rule has been checked in the scope ofthis work.

Tbe results ofthe Miner evaluation are presented graphically by means offig. 11.

Tbe application of the Palmgren-Miner-Rule did not result in a good accordance in the field investigated (fig. 11).

This statement is applicable to all possibilities of modified S-N-lines (different cut-off limits; different slopes etc.) [5] and fig. 12.

From figure 13, the frequency of the Miner sum determined from the tests can be seen. The results of the investigations made in Karlsruhe are recorded comparatively into the diagram according to Schütz/Zenner [3].

It can be concluded from available tests that the cumulative damage results in values below 1.0 for load collectives with a higher fullness, as for example 2/3, and in values between 3.0 and 4.0 for a smaller fullness.

For the practical application of the Miner-rule, the cumulative damage indicated in table 2 is recommended for different types of collectives.

These indications are valid for structures with sm all tube dimensions as they occur for crane systems, masts etc. Presently, experimental investigations are being carried out with hollow sections of a bigger dimension and bigger wall thickness.

5. Proposal for the design ofhollow section component parts under load spectrum

As the design calculation of the structural component parts by using the Miner-rule or the relative Miner-rule did not give satisfactory results in the past, new design methods were sought for practical application.

While considering the load spectrum collectives with the area under the corresponding staircase curves, and the endurance lines for these collectives determined by tests according to figures 5 to

---~---7 and 9 as weIl as the test result given in [4], it occurs to one that the maximum stresses or altematively the endurance (load cycles) for certain stress level increase with the decreasing fuIlness coefficient.

The fullness ofa collective is an integration ofeach load level overthe number ofcycles (N). The following hypothesis is formulated using thls reverse effect:

"The ratio ofthe fullness (area) oftwo load spectra behaves inversely proportional to the

load increment coefficients of these load spectra".

According to the hypo thesis, the following is valid for the load spectrum collectives A and B:

fullness ofthe collective A fullness ofthe collective B

load increment coefficient for the collective A load increment coefficient for the collective B

The load increment coefficient is defined as follows:

maximum stress range in the spectrum for ajoint under collective i (fatigue life line)

f= ₁

maximum stress range for an identicaljoint according to the Wöhler tests (CA) Fig.

14

shows the determination offi-values (schematically).

In general, the above mentioned hypothesis can be expressed for any load spectrum collective i asfollows:

Y·₁ ·f=C ₁

where C is a constant, which has to be determined empirically by tests.

The argument for the hypothesis is that the fullness coefficient or altematively the product ofthe fullness coefficient and the corresponding load increment coefficient is considered as physical

value and defined as follows:

The fullness coefficient represents the quantity of energy applied on the structural component part in the form of deformation energy.

The product of the fullness coefficient and the load increment coefficient represents a physicallimit, a measuring number, which shows the load bearing capacity ofa structural component part for any load spectrum collective, where the constant amplitude (Wöhler) test is taken for reference collectives.

Advantage by using this method is that only one reference S-N-line (Wöhler = CA) is required in total, in this case the Wöhler-line for the structural component part is to be designed. The proof for the fatigue strength for any other load spectrum can be done using this recalculation factor.

In order to obtain the life endurance line for any other load spectrum collective, it is only necessary to carry out constant amplitude tests for a certain structure or use constant amplitude tests already available.

Disadvantages ofthis method

The lowest value is obtained by using this method.

This hypo thesis was checked for various load spectrum collectives by means of extensive investigations.

Table 3 contains the maximum and minimum values ofthe load increment coefficient calculated using the test data and the measuring number for the design according to the figures 5, 6, 7,9 and other test values according to [4].

A constant close to 7.5 . 107 for a collective range of 106 loading cycles is obtained by multiplying the corresponding fullness coefficients with the measuring values on the right hand column of the tabie. This value is, however, only an aporoximative value, that represents an acceptable calculation value for all described load collectives.

Ifthe fullness coefficient for a new collective be now obtained by integrating the staircase curve, the load increment coefficient can be calculated by determining the ratio of the limiting value

7.5

'

lQ

7 to the given fullness coefficient.

6. Example forthe calculation

The proof ofthe fatigue strength is to be given for aportal crane made of circular hollow sections. The goveming collective is P = 0/3. The load cycle number to be calculated is 2 . 106.

For the given joint geometry, a load cycJe number of 2*106, a probability of survival of 50 % (Pü = 50 %), and a bearable stress range of 42.0 N/mm2 are determined from the Wöhler test

(CA-test).

42,ON/mm2

32,1 N/mm2 LW

2'106

Conversion ofthe bearable stress range ofPü = 50 % to Pü = 97.5 % can be done by means ofthe factors from the table in [5].

6050% I 6097.5%= 1.31

Thus, this results in C~PPü=97 .5% = 42.0/1.31 = 32.06 N/mm2. (This value can be also taken from the standards)

For a collective P = 0/3 _{the fullness V O/3 is obtained (for a maximum stress range of} SR = 100 N/mm2 and for a collective range of! 06 cycles.

V₀₁₃ = 2.115 *107

The load increment coefficient fOI3 is calculated through 7.5 *107 2.115 *107

3.54

The stress range L:PPQ=97 .5% = 32.06 N/mm2 calculated for Pü = 97.5 % is multiplied with this load increment coefficlent.

~,aPü=97.5%

* fOI3 = 32.06 * 3.54 = 113.5 N/mm2 '{ m = 1.25 (from the chapter "Proposal for a saftey concept") [5]

C:~.a97.5;0/3

= 113.5 90.8 N/mm2

'[; m 1.25

Thus, the allowable maximum stress range for this joint geometry under spectrum loading P = 0/3; for 2'106 cycles, is 90.8 N/mm2.

References

[1] N.N.: Eurocode 3, Design of Steel Structures, August 88

[2] Mang, F., Bucak, Ö. and Klingier, 1.: Wöhlerlinien-Katalog fUr Hohlprofilverbindungen (S-N-Line Catalogue for Hollow Section Joints)

Studiengesellschaft fUr Anwendungstechnik von Eisen und Stabl e.V., Düsseldorf, June 1987, January 1988

[3] Schütz and Zenner: Schadensakkumulationshypothesen zur Lebensdauervorhersage bei

schwingender Beanspruchung Zeitschrift fûr Werkstofftechnik 1973 Pan 1, No. 1 p. 25-33

Pan 2, No. 2 p. 97-102

[4] Mang, F. and Ö. Bucak: Investigations into the Fatigue Behaviour ofHollow Section Joints Subjected to Spectrum Loading

ISOPE '91, Edinburgh, 11.-15.081991

[5] Bucak, Ö.: Ermüdung von Hohlprofilknoten (Fatigue Behaviour ofHollow Section

Joints) PhD Thesis, University ofKarlsruhe, May 1990

[6] Bucak, Ö. and F. Mang: Investigations into the Fatigue Behaviour of CHS-Joints Subjected to Spectrum Loading, Third International Symposium on Tubular Structures, Lappeenranta, September 1 and 2,1989

[7] N.N.: Stahl in Meeresbauwerken

Proceedings ofthe International Conference in Paris, Oct 15-18, 1981 EUR-Bericht Nr. 7347 (1981), S. 439-483 und Plenary Session No. 5

-

- - -

-[8] Bierett, G.: Einige wichtige Gesetze der Betriebsfestigkeit geschwei13ter Bauteile aus Stah!. Schwei13en und Schneiden, Heft 11, 1972, S. 429-434

[9] Haibach, E.: Modifizierte lineare Schadensakkumulationshypothese zur Berück

-sichtigung des Dauerfestigkeitsabfalls mit Fortschreiten der Schädigung LBF Darmstadt, TM 50170

[10] Lipp, W.: Zur Lebensdauerabschätzung mit dem Blockprogramm und Zufalls

-lastenversuch

LBF Darmstadt, TM 76/80

[11] DIN 15018: Krane, Grundsätze fUr Stahltragwerke; Berechnungen, Ausgabe April 1974

[12] Sonsino, CM., H. Klätscke, W. Schütz and M. Hück: Standardized Load Sequence for Offshore Structures

WASH-1

LBF-Report No. FB 181 (1988) IABG-ReportNo. TF2347 (1988

[13] Sonsino, CM. and K. Lipp: Übertragbarkeit des an Winkelproben ermittelten Betriebs-festigkeitsverhaltens aufgro13e Rohrknoten fUr die Offshore Technik

5. LBFKolloquiumin Darmstadt, March 8-9,1988 Bericht Nr. TB-180 (1988)

- - - -

-G

~p

I I I I I I

~

p

Fig. 1 X-typejoints made ofhollow sections (circular and rectangular) forthe tests carried out in Karlsruhe

Type of joints chord member bracing material R

0101,6x5,6 051 x4,0 +0,1

0

X 0101,6x5,6 0101,6x8,0 082x4,0 082x4,0 St 37-2 +0,1 -1,0 K 0101,6x5,6 051 x4,0 +0,1

D

X 100x 100x4,0 50x50x4,0 +0,1 100x 100x6,0 90x90x3,0 St 37-2 +0,1 K 100x 100x4,0 50x50x4,0 +0,1

Table 1 Data ofthe test specimens

Fig. 2 Fig. 3 f-- 1_ _ 1_ •. ,

1

=

1=

-=-

=

-=-

=

1=

/-- - - - -1 -f-- /

-1=

'=

-

/-r=

=

1=

f--- - t-- - / -.. W' .' . ' . . . ' .' Wohlpr Ipsl

•

~

--

=

. ' . .

--

. " '

..

--

--- P,U) - - -- - po1n _ -- - - - -- - - -• -• ' . ' . ' W' .' .' ·'H • • ' . ' . ' . ' . - .. ·'N • • ' . ' . ' . ' . - . ' .'w

SPfclrUIII P ~ lil _{SPfctrum p: 11]}

,. ,ur wind SptrlrUiIJI

.

~

..

~

.

~~"'

--~.~%--.. ", .. ' .. ' "'"

..

..

....

..'...'...

'

..

..,.'.'...'...'.

.

..

,

.,.,.,....

....

Norlh Sfiilload SpfclrUfll GtllA8G/tBFI Horlh Sn 10. spfc'rulIIlIABG/WG)1

tor 1 year tor 1 yur

Compilation of spectra applied

Joint failure in chord shear (CHS; dld

_o

= 0.5 and tolt = 1.4)

A19

A typlul 1.'1t}Jf 1Ndrw} spKtru. 101' SC J'fVS 11'1 ... GoH ol tit.l( 0

-Fig. 4 Joint failure in the chord shear RHS; b/b

_o

= 0.5 and tolt = 1.0)

700 600 500 400 300 200 150 100 Sa IN/mm')

r-Ip'fl

rl

S -N-L.ne

r-

r--:-I

Lapla,·1 LP

,t

~ ... I"'-~ + r.~f-I _~ waO.6 r'" _r ~ ~~ (hord I('\lorf~ (hord: .101>15.6

~

-48G I", web member: • 5'"4.0

~p~

/Vor I ' G 1II web membf.'r

~I materoal:Sl37 ~ ' - - ' l8" I-~'-F=:r--:::t--

---_

...

~~

r=:::r--

lt nI lil..; ~ ~ • b-. r---~ - -...ot

r---

r--~

I--

r--~

F==i:: - -'

r--r=::~-r-

I---I-t--

--

-

~

... ___ 50 40 • $-N-lml'

..

--....

•

...

r::::--:t: +

:::----

-30 + p.21l

~---

• •

-

::::::::",

"-~ )( loplo(t A p.lIl

•

•

--

-!

~--

"-~

"

'Wind ::::" -I---20

0 North Sec IlBF

• Nor'h S.o I ABG I WG 111

10 !

10 2 4 6 810' 2 4 6 810' 2 4 6 8 10' 2.,0' 4 6 8 10'

N

Fig.5Results ofthe fatigue investigations on CHS X-type joints underload spectrum (dld

_o

= 0.5 and tolt = 1.4)

Fig. 6 Fig. 7 300 200 150 100 70 50 40 30 20 10

f-[jo IN· mm·1_{] (Maximum stress as nominal stress of the bracmg member)}

) ) _{] I} I rw=3.omml11 chord 100.100.0.0 "b ••• b"

".;0.

~

l

p,

~/

_r-...

mater,al: St 37 +- ' , _

I

--I'--

r--...

l

p,

1/

~ ~

1'1--.

t---.... cb , _...

r-....

r----.

S-N I,ne j-....

r::---

-...; t--_t--r--...

""'"

"Ëilli~

t--

1---,

_I;-...

I

f...

_....

r-

t--t---.... ... I a.

t--... --!t--....

I

R • 0.1 10> 2 4 6 810' 2 4 6 810' 2 4 6 810" 2.10" 4 6 810' -N

Results ofthe fatigue investigations on RHS X-type joints under load spectrum

(b/b

_o

= 0.5 and tolt = 1.0) 600 <100 200 100 80 [N/mm 2J Im ~ = 3.5 I~ <10

Elli1J

(Jo - SF< (Ju 60 R ~ 0.1 20 3 10 2 100.,00.4.0

~

Chord: Web member: 90.90.3.0

-c{P-Matarlal: St 37

~

1 1 1 '-... I'---. p = 1/3 I

1 I

I

_I

I I

~

I

~

I~ _... l'-.. ' , i

"

_'

,

1'--... ~ "'-...' , . I

_.

"---I , I~I I I ~ ~

r-::::

r32.5 N/mm 21

I

I

N <1 6 8105 2

Results ofthe fatigue investigations on RHS X-type (b/b_o = 0.9 and tolt = 2.0)

under laad spectrum P = 1/3 .

Fig. 8 Fig. 9 Cl P [kN J j ".oj-- - - t - - - - j -I r j t + -37. 27

-t----t----+-11 _ _ lOm" = 70 1 N/mm21

I

~I---,

1

~:~~~~_-_

_{E--I-_-=3}

A

~

~

correspon~s

10 a 10 lal elongatron

31.78 kN-:" of Ihe lesl speCimens of 0.6 rnrn

\

25.0t- -24.02

-l

i

r

chord member

~1016

,56 ~ web member ~ 510 ' 4.0 ~p

I

I

vrsrble crack

IR =+ 0,11 rnrtralror (-15% end of lesl (ÓI, 2,lmml

20.0

'---==---~:__--~--+----:-:":.

I

t__

, __ =:__----::::':-:-:-< 1

-20000 40000 60000 .

!

8000₁ 100000 120000

70S161 C 814731 c laad cycles

Reduction ofthe upper loads (simultaneous increase ofthe lower load) when testing a hollow section joint upon a constant oil transporting amount (gain) with a testing machine without internal con trol system in dependence on the crack or load cycle number.

In the figure, each stage corresponds to an elongation ofthe test specimen by 0.1 mm

Go [N' mm-ZJ Maximum stress as nomlnal stress of tension diagonal

500 f---.---,----,--,-,--r---r----r-,....,---rI- - ,Ir--TI- r -1r

-l - , - -1---'-1- ' -1-'--1 _

400 1---+---+--+-+-+--+--+---1-+-+-- eh 0 rd' 100.100.4,0

300 t -- t --t-+-t-t----i'======='''"rt-t-- dl ago n a I' 50. 50.4,0 / ' _

r-:--1 North Sea lIJ materlal' St 37

'

~'

22'

B

Wind -..; -200 f--__ f----j- p' l ~ ----J _ _ : : - - :-r-. - -

-lp,

t

I---;::::::::r:::~;::::~

-

~

--

---

--100 f----+-I S _ N _ 1I nel ~r- --=:~:::--r-~

____

g. 2'>.4_

I

m ,

5_2{

--~j--

---r---t~f::::::::::::t:::j--:-r-I

:-r-r-a-

0

:-r-r-__

r--':::F=::r-r-50 f - - - - j - - + - _ I-- r-- r-

r-r-40 _ r - - - ' - - - ' ---'- " - -. -r- --t--, --t--, - '.

~

s,

l

~r---

r-r-20- ~ , au R • 01 1

I

10L--~-~-'-~-~-~~~~-~-~~~~-~-~~~~~ 10' 2 4 6 810' 2 4 6 810' 2 4 6 810' 2.10' 4 6 810'

Results ofthe fatigue investigations under load spectra ofRHS K-typejoints

(b/b

_o

= O.S; tolt = 1.0)

Fig. 10

Fig. 11

Test specimen after [ailure; K-typejoint (b/bO = 0.5)

\0' I---.,---.---~ 1O'r---t----~~---L-~ 10' f---,/L---I~ .... _!_lL----__1 "

....

\0' X"JOlOt 00.8/1.4 •

I

Pallngr.n -Miner-Rul.

\0' \0' ",g N .... t H: mean utufS • -Wind .-p=~ .. -p = t .-WG D1-North Su • -l8F -North 50. • -laplit!

Graphical illustration ofthe Miner resuits

A23 lel Neale, 10 ,

I)~

.

, 10

/' J.:

* ~

.

.,

/4.

:V~

10 10 10' x -joint 0 05/1.4

I

Patmgren -Hiner- Rul!

'" 11' ~

..

111' leoN ... , H = Iftun ulues • - Wind .-p=~ "-p = t • - WG III-North S •• • -l8F -North 50. • -taplace

1

'

'

,

I

~

I

~

___ -I ...

I

I L -_ _ _ _ -LI ___ ... ~~. I • 2.106 • 2106 t

~

I

I

I' I ' 5106 •

Fig. 12 Different cut -offlimits ofS-N-lines for the application ofthe Palmgren-Miner-rule

-5 120 r

Steels. Al-and TI- alloys

VI <-( room - temperature. ClJ '= 100 r _{without corroslon)} L

Ö _{block program tests}

<-80 r ClJ total 348 calculations .D E ::::J _{. tests in Karlsruhe} c:

-

60

>-c:::J

X - jOint 0 u c: ClJ

c:=J

K - joint 0 ~ 40 ClJ

-

<-CD

K - JOint 0 u. •.. - *> ... --: 20

~~

N

O ~-=~~~LL~~~~~~~~~lJ-__ 0,01 0,02 0,05 0,1 0,2 05 1.0 2,0 5,0 10,0 ni Cumulatlve damage S = [~

Fig. 13 Failure frequency; presentation according to [3] supplemented by tests carried out in Karlsruhe

Damage accumulation

Fonn of a load spectrum Damage accumulation non considering the series

Collective All test results with the crack on the

(All types offailure) , inside ofthe tube

for all R-values and R~ 0,1

S _nun . _Sproposed S . _{nun Sproposed}

2/3 0.27 0.3 0.27 0.5

1/3 0.34 0.5 2.11 3.0

Laplace 1.87

---

1.87 2.5

North Sea (LBF-IABG) 0.48 1.0 3.18 4.0

North Sea (LBFWG III) 0.56 1.0 1.59 4.0

Wind 0.33 1.0 1.47 2.8

Table 2 Cumulative damages in dependence on the type of collective Smin is the lowest value detennined during the tests in Karlsruhe

Sproposed is the sum of damages to be used for the caJculation. The poor single results due to weid defects and roughnesses ofthe specimens were not considered

North Sea I

SR' 2/3·1Q1 t---+_r----""oç--~ ...

SR .'o7t---+----~k

laad cycles

Fig. 14 Determlnailon of the laad factors I schemalic )

~ - -

-Definition ofthe Fullness1) Load increment coefficient Proposed ratio value2)

oad spectrum for the design

Min. Max.

Block collective) V collective fcollmin f coll.max. fcoll.design

2/3 6.586'107 1.31 1.68 1.3

1/3 4.395'107 1.82 2.98 1.8

lNonh sea LBF 2.330'107 3.15 5.60 3.0

lNonh sea WGIII 2.164'107 3.40 6.80 3.4

Wind 2.201'107 2.87 4.80 2.8

1) determined for a maximum stress range of Sr = 100 N/mm2 and a collective range of106 cycles

2) due to the low number of the test data, one starts from the lowest value, while the minimum value is rounded offto the lower side (presently, fcoll,design is similar to the increase factor fE)'

Table 3 _{V collective' fcollective-design for different load spectra}

FATIGUE BEHA VIOUR OF RECTANGULAR HOLLOW SECTION JOINTS MADE OF HIGH-STRENGTH STEELS

F. Mang, Ö. Bucak and K. Stauff

Versuchsanstalt fur Stahl, Holz und Steine Universität Karlsruhe

1. Introduction

The reason for applying high-strength rolled sections is th at effective dimensions and cross sections can be economically produced through a prefabrication in the rolling mill and thus, the favourable propenies ofhigh-strength steels can be utilized for structures.

In the last years, the structural steel market increasingly tends to products with bigger wall thicknesses and higher strength. Hi-tech areas such as the offshore industry and high building construction in addition require a high toughness as weil as improved working propenies of these steels.

With the publication of the German standard DASt-RI 011 (February 1979) "Application of High-Strength Weldable Fine-Grained Structural Steels StE 460 and StE 690 for Steel Structures" [2], the increasing application of high-strength weldable fine-grained refined structural steels has been taken into account. In the meantime, they have developed a broad field of application such as in the construction of tanks, pipelines, vehicles, cranes and steel framed structures. This development is essentially connected with the progress in steel production as weil as in welding technology. The previously mentioned DASt-standard is presently revised.

lts economical application matches with the utilization of higher allowable stresses compared to those of "conventional" structural Steels «St 37 and St 52). In this case, those structural members come off favorably that are subjected to tensile stress or those structures that are loaded under fatigue stress, having a lower fullness ratio. In addition, high strength steels under fatigue stresses, for which pans of the stress show higher values than the allowable stresses, are advantageous for the static measurement, for example through a too high mean stress and a low load cycle nu mb er.

The latest editions of Eurocode 3 on steel structures do not indicate any regulations for structures made ofhigh-strength steels.

Therefore, the results of the experimental investigations on hig!. c:-ength steels are provided in the scope ofthis paper. In addition, they are compared with the results of conventional steels.

In [4] and [5], structural shapes are indicated forwhich the same reduction factors are valid for the dimension as they have been previously worked out for the conventional steels St 37 and St 52. These structural shapes are bending stiffened corner plates for which the failure occurred on this positions through by-passing the forces to the more stiffened corners and plastic buckling resulting from this, as wel! as through T-joints under moment stress. With this, the torsional stiffness ofthe joints have been compared with those of conventional steels. In this context, they resulted in a good accordance with the results of conventional steels. It is planned to continue these investigations with high-strength steels under static load.

An

EGKS-research program on the fatigue behaviour ofthe most imponant notch cases as wel! as on X- and K-typejoints made ofthe high-strength steels StE 460 and StE 600TM is presernly performed in Karlsruhe [6]. In the fOllowing, the first results are provided.

2. Principle investigations on high-strength steels

Test series with sections made ofthe high-strength steels StE 460 and StE 690, partly showed a parallel dislocation ofthe corresponding S-N-line (Wöhler) as it comes out of fig. 1 for a 100% overlapped K-joint made of rectangular hollow sections.

Since the number of test series with StE 460 and StE 690 was low and there we re contradictions for the results with regard to the course ofthe S-N-line in a high load cycle range compared to the results of small test specimens, these steels have been excluded from standardization work up to

now.

Based on the draft for Eurocode 3, the stress range (SR = 6. a) is taken as a basis for future

diagrams. From fig. 2, the justification for the application of the 6. a-concept for welded structures made ofhigh-strength steels can be derived.

3. Designoflongitudinal attachments (ribs) on hol!ow sections made ofhigh-strength steels.

The advantages of high-strength steels such as small cross section surface and resulting lower dead weight as wel! as their notch sensitivity have been known for a long time. For some time it has been recommended to design the structure suitably for materiaIs. One of these recommendations is the design oflongitudinal attachments. The simp lest way is to shear off or saw off the attachments from the flat material, and to weId them to the required position by means offillet weIds. In Eurocode 3 [1], this notch case has been classified into category 50, 71 or 80, depending on the length of the attachment. With an attachment length of 200 mm (corresponding to detail category 50 according to Eurocode 3) for high-strength steels, the experimental investigations carried out in Karlsruhe resulted in a value of about 80 N/mm2 for a survival probability of97 .5% (fig. 3).

lfthe abrupt stiffness. increase on the same structural details is avoided by the fact that it received a continuo us transition with a radius r = 40 mm, and if the transition has been worked off by grinding after wel ding of the attachment by ~eans of a double high-strength weId, better values ofthe stress ranges are obtained for the 2·10 load cycle. The results of these investigations can be taken from figure 4. In the same figure, they are recorded comparatively with the results ofthe rectangular attachments.

Thro~gh the comparison ofthe values for a survival probability of50%, an increase ofthe values 2·10 load cycles resulted in a factor around 1.4; where for the load cycle numbers, an increase factor ofabout 3.0 was obtained. These different factors are to be put down to the relatively low slope ofthe S-N-line. The detail categories for longitudinal stiffeners of 45, 71 or. 90 (r) 150 mmJ as weIl as the slope of the S-N-line with 3.0 within the load cycle range smaller than 2·10 indicated in Eurocode 3, do not correspond to reality.

From previous experimental investigations on structural members it can be derived that a slope of 4.0 or 5.0 better corresponds to reality compared to the slope of3.0 which has been mainly ascertained for smal! test specimens.

Sin ce, through working off of the stiffener end in the joint area, only a small stiffener length remained, further experimental investigations have been carried out with a rectangular stiffener

1= 120 mm for clarifying the influence ofthe length ofthe attachments. In figure 5, the results [or Pü = 50% are recorded comparatively.

Figures 6 and 7 show the test specimens after failure, where the smooth transit ion radius r initially formed by machine or gas cutting ofthe gusset plate before wel ding, and subsequently grinding ofthe weId area parallel to the direction ofthe arrow.

From figures 8 and 9, the hardness distributions on these test specimens are evident. With values of HV 10, hardness peaks usual in practice are the result which points at a practice related production ofthe test specimens.

4. Bun welded joints on hollow sections made ofhigh-strength steels

The next group of notch cases are butt welded hollow sections. Tenacito 75 (E-I0018-6 according to AS TM) for the material StE 600 TM and Tenacito 60 (E-8018-6 according to AS TM) for the material StE 460 TM have been choses as electrodes. The edge preparation was V-shaped with an included angle of a = 600, and a gap of approximately 1.5 to 2.0 mmo The results ofthe first investigations can be taken from fig. 10. As expected, the bearable stresses are much higher compared to the classes ofEurocode 3. Unacceptable lacks offusion were stated for 3 test specimens. They have been subjected to a fatigue test in order 10 determine the decrease of load cycles. These 3 test results are also plotted in the giagram of fig. 10. The evaluation results in a decrease of the stress range of 72% for 2·10 load cycles where the bearable cycles decreased to ab out one fifth.

Tests are being continued.

5. Hollow sections with transverse attachments

Some hollow sections are provided with a transverse rib with the dimensions 100 x 60 x 6,0. The weldings are carried out with rutile electrodes as weil as with Tenacito 75 (alkaline electrodes). A common evaluation of all test data available can be taken from fig. 11. It becomes evident that there is no big difference between specimens welded with alkaline electrodes and those rutile electrodes. In contrast, the test specimens made ofthe material Q StE 460 TM are in the lower part ofthe scatter area. The slope ofthe S-N-lines is flatter with a value of 4.5 compared to the indications made in the Eurocode 3. The value for 2·106 10ad cycles is farly higherwith 123.7 for Pü = 97.5% than th at ofEurocode 3 forthe mild steels St 37 and St 52 (category 80).

Fig. 12 shows specimens af ter failure. 6. Conclusions

In this paper, the first results of experimental investigations made on the high-strength steels StE 460 TM and StE 600 TM are presented and compared with those of Eurocode 3. In addition, it has been shown that an increase of fatigue strength values can be gained by an additional treatment ofthe ribs appropriate for the material involved.

References

[1 ] Eurocode 3

[2] DASt-Ri 011

"Design of Steel Structures", Part 1-9 Nov. 1989

Feb. 1990

Sept. 1989 Eurocode 3, Annex

"Hochfeste, schweiBgeeignete Feinkombaustähle StE 460 and StE 690, Anwendung flir Stahlbauten" (High-Strength, Weldable Grain Refined Steels StE 460 and StE 690, Application for Steel Structures)

Stahlbau Verlag Köln, Feb. 1979

[3] [4] [5] [6]

[7]

DIN 18 808 Mang,F. Bucak, Ö. Mang,F. Bucak, Ö. KlingIer, I. Mang.F. BucakÖ. N.N.

"Stablbauten, Tragwerke aus Hohlprofilen unter vorwiegend ruhender Beanspruchung" (Steel Structures ofHollow Sections under Predominantly Static Load) Beuth Verlag Köln, Oct 1984

"Behaviour ofStructures ofHigh-Strength Steels" RILEM-Workshop on "Needs in Testing Metals" Naples, May 1990

Wöhlerlinien-Katalog fûr Hohlprofilverbindungen, Studiengesellschaft fûr Anwendungstechnik von Eisen Stabl e.V., Düsseldorf, Juni 1987 und Januar 1988 "Untersuchungen an Verbindungen von offenen und geschlossenen Profilen aus hochfesten Stählen", S. 11 und S. 71

Studiengesellschaft fûr Anwendungstechnik von Eisen Stahl e.V., Düsseldorf, Dez. 1977 und 1981

Fatigue Behaviour ofHollow Section Joints made of High-Strength Steels

EGKS-Forschungsprogramm Nr. 7210-SA 117 (EGKS-Research Program No. 7210-SA 117) Contractor: Fa. Kloeckner-Mannstaedt

- - ---~---

~---Fig. 1

Fig. 2

Go IN mm-2J Maximum strrss as nomlnol strpss of trnslon dlogonol

200~~.--_,-r_,,--,_-,__,_,_.-_,--,__.---_.

I

150 100 90 80 70 60 50 40 I

I

GO~

S. ~-~--+.S~t~E~69~O~~--~-~-~-+---+---+--+ Gu St 52 ~ ... R: 0,1 I ... " " I I --

r--~

(hord dlogonols 100.100.8 100. 100. 8 ~ ~

J

41,0 N/mm I

I

32,5 N/mm I ~ 1

1

! I

1

I

30

r---

... ""-

"

I

' '

20~~1_-~

I

~I_~

I

~

I

J

_~I~

I

~

I

~

I

~~

I

~--~~~~~28

,

-'N~/i-ml~

I

'

~

l

~

1

~

ovrrlop

to l 6 8 10' 6 8 tol 6 8 10' 2' 10' 6 8 to'

Test series with the samejoint geometry but with different materials

500

400

S;. [N mm-2j Stress range - normmol stress

, I

,.

--

---

- -'._-- -200

'~~

70

=_~_

~~

~

==~

~-

=======

-

=

-

-=-~~

.

~-~

-

~

~

60 f--- - -50

---

--- -- - -40 - -.---- - -30 - - - - -- +---+---20 RHS 80. 80.6,5 St E 460 • R:. Ol • R,· 1,0 • R:. 0,5 ., 55:22 Nimm:2

•

43,52 Nlmm 2 / ' 3626Nlmm1 ,0,~0~l---6--8-'~0'--~--~--~6~8-'0~'--~---6~8-,~0'--~2~'0~'---6~8-'0~

Justification of the application of the Cl 0 -concept of welded structures made of

high-strength steels

- -S.. [N/mm2 ] 600 ~-.--'-'-rr--'--'-'-''---~L=2Q ~---+ + ... " oos·' 'I --'---L---,I 9~

r:::

-='"~:l_t-"~"..__+__+_+_l

t

5ii2

t

-

.

400

-rt

m - 3.8 ~ " " " t'... " RHS 100 x 100 x 4,0 IongltudlN11 .t1ach. 200 x 60 x 6,0 130.6 N/mm 2 ₁ 100 80 R = 0.1 N 4 6 8106 2 4 6 810 7

Fig. 3 Results of the experimental investigations on hollow sections with longitudinal ribs (attachments); material StE 600TM; Pü

=

2.5%; 50% and 97.5%

400 1

1

~

'I

;--

I

tl

Îl-I

'I

i

i I - I' L

,--~-H

-I-

I--I 1 ! : 11 200

-

l

--r-

~i

i

.

_

;

;-

~I

I-;--I

'l

'

~

~r---

r

\135.1 N/mm

2J

___

iJ_I

__

~

_

~

~

195.6

N/mm 2 100 80 ;---j--:-I- H - -I---II- j--60 R = 0 1 1 N 103 2 4 6 810" 2 4 6 810"' 2 4 6 8106 2 4 6 8107

Fig. 4 Comparison of the results of fatigue tests (Pü = 50%) on test specimens with longitudinal ribs (attachments) stiffness; different design ofthe longitudinal stiffeners;

material StE 600 TM

A32

[N/mm2 ) 600 'ZQ 80

i

I

....

rm = 3.6

'"

008-Q

_t

c:::J I 9~ -<00..

t

- l

~

_~

AHS 100 x 100 x 4,0 longItudInai 8ttaCh. 200 x 60 x 8,0 ___ e

;--

_~

120 x 80 x 8,O_ .X

N

•

~ "-,~ e~

~

_~

~

e It

~~

EmzJ

(Jo f103.6 N/mm21 - SR 1 - -. (Ju 400 200 100 R = 0.1 N I 60 3 10 2 4 6 810" 2 4 6 810 e 2 4 6 8106 2 4 6 8107

Fig. 5 Comparison ofthe results of tests on specimens with a longitudinal rib (attachment) in

dep enden ce on the stiffener length

Fig. 6 Test specimen with rectangular stiffener 1 = 200 mm after fallure; material StE 600TM

Fig. 7 Test specimen with longitudinal rib (grained) af ter failure

The smooth transition radius r is initially formed by machine or gas cutting of the gusset plate before wel ding, and subsequently grinding ofthe weId area parallel to the direction ofthe arrow.

RHP 100, 100,4.0 LongitudinaJ rib 200 x 60 x 6.0

Electrode Tenacito 75

[·10018-6 (ASME)

Fig. 8 Macro section with hardness distribution, longitudinal attachments, material Q StE 600 TM

1

;.:i!!'. r, :O{l\-L!1

, 'n~:: ,IJ1:1-: rit"-I ·HI Olm

r .•• !I \ t\.l~ I r.:Il<H:IlLl -~ I .,)!':~.'" t \~\lEI

r) "':1 ""U(, 1 \1

Fig. 9 Macro section with hardness distribution, longitudinal attachments with

radius r = 40 mm, material Q StE 600 TM

[N/mm2] I

r

₃₀₀

-i

ZOOr300 , I

_I

_I

I;

I

_z:;

I 0

1000 800 X ... buttw.ld

600

•

... buttw.ld wlth laek of fu.lon

400 RHS 100 x 100 x 4.0 I--lJm = 6.2

I

StE 600 TU ... I ... ... X ... r-....

!

I

I

~~

I

><~

~

I

I

_! ... r---

f"-...

I

'r-.... r- I° "~ I

>

cro~

118!'.2 N/mm 2 1 - s"' r---...r--.. 1---1"-... cru 200 R = 0.1 N 4 6 810'" 2 _{4 6 810} 6 ₂

Fig.l0 Results ofthe experimental investigations on butt weldedhollow sections

s .. [N/mm2_] 1000 -ZQ.Q...

i i

... I I

_B~

t

2,

_t

lZOO 800 600 _{, ,} transverse attachments RHS 100 x 100 x 4,0 ,

-Il

m = 4.6 J~ , , attachment8 _{StE 600 TM} 100 x 60 x 6,0 , "- , ,

""-N

, , _, , 1'- , , , 400 , .~ , , fl!. ' ,

~

, ,

~

, , ,

~

, , i' , 189.7 N/mm 2 (jo _'-L

~

- S .. , , 158.1 N/mm 2 , _, (ju _, R = 0,1 131.7 N/mm 2 200 N 4 6 810'" 2

Fig. 11 Results of the experimental investigations on hollow sections with transverse attachments

Fig. 12 Test specimen with transverse attachment after failure