International Symposium

On Tubular Structures

Delft, The Netherlands June 26,27 & 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

On Tubular Structures

Delft, The Netherlands June 26,27 & 28, 1991

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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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lc3il .) 'r 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 Njmm 2 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.

A 1

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.

A 2

On loading the dome to 60 % of the design load, unacceptably large deflexions and local dimples developed in the region of the pendenti­ves. 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.

A 3

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, permit­ting 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.

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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 stiff­jointed. 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

A 5

[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

A 6

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photo 2.

photo 3.

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figure 1.

A 8

fig. 2. Bolted connection tube node

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figure 5.

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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

All

-- -- - - - - -

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 S-N-lines for different load collectives has the same slope as the S-N­lines 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.

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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 fE = 3.4 fE = 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

A13

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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· ·f=C 1 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.

A14

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)

A15

For a collective P = 0/3 the fullness VO/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

A16

- - - - - -- -

[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 Berück­sichtigung des Dauerfestigkeitsabfalls mit Fortschreiten der Schädigung LBF Darmstadt, TM 50170

[10] Lipp, W.: Zur Lebensdauerabschätzung mit dem Blockprogramm und Zufalls­lastenversuch 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 Betriebs­festigkeitsverhaltens aufgro13e Rohrknoten fUr die Offshore Technik 5. LBFKolloquiumin Darmstadt, March 8-9,1988 Bericht Nr. TB-180 (1988)

A17

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~p

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~ 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 082x4,0 St 37-2 +0,1

0101,6x8,0 082x4,0 -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

AlS

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; dldo = 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/bo = 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 ~

w a O.6

r'" _r ~ ~~ (hord I('\lorf~ (hord: .101>15.6

~ ~p~ -48G I", web member: • 5'"4.0 /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 (dldo= 0.5 and tolt = 1.4)

A20

Fig. 6

Fig. 7

300

200

150

100

70

50

40

30

20

10

f-

[jo I N· mm·1] (M aximum stress as nominal stress of the bracmg member)

) ) ] I I rw=3.omml

11 chord 100.100 .0.0

"b • •• b" ".;0. ~ lp, ~/

mater,al : St 37 +- ' _

r-... ,

I --I'--r--.....

lp, 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--... --!....

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/bo = 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'---. I 1 I I I p = 1/3 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/bo = 0.9 and tolt = 2.0) under laad spectrum P = 1/3 .

A21

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

~:~~~~_-_ A E--I-_-=3

~ ~ correspon~s 10 a 10 lal elongatron 31.78 kN -:" of Ihe lesl speCimens of 0.6 rnrn

\ 25.0t---

24.02 -

l ir chord member ~1016 ,56 ~ web member ~ 510 ' 4.0

~p

I I vrsrble crack

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

20.0 '------==-----~:__--~--+----:-:":.I t__ , __ =:__----::::':-:-:-< 1 -20000 40000 60000 . ! 80001 100000 120000

70S161 C 81 47 31 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-1 r--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_

Im

, 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--

,,- '. ~ 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/bo = O.S; tolt = 1.0)

A22

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 • 2.10 6 • 2106

t

~ I

I

I ' I '

510 6 •

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 withou t 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

A24

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 . 'o7 t---------+----~k

laad cycles

Fig. 14 Determlnailon of the laad factors I schemalic )

A25

~ - - -

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

A26

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.

A27

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.

A28

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 2·10610ad 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" (High­Strength, Weldable Grain Refined Steels StE 460 and StE 690, Application for Steel Structures) Stahlbau Verlag Köln, Feb. 1979

A29

[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

A30

- - ---------------------~---~---

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 S. I GO~

~-~--+.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 _~I~I~I ~I~~I ~--~~~~~28,-'N~/i-ml~I' ~l~1~ J

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-2 j 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

/ ' 3626Nlmm 1

,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 high­strength steels

A31

- -

S.. [N/ mm 2 ]

600 ~-.--'-'-rr--'--'-'-''---------~L=2Q~---------

+ + ....... " oos·' 'I --'-----L-----,I 9~

r:::-='"~:l_t-"~"..__+__+_+_l t 5ii2 t -. 400 -rtm - 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 ._ ; ;-~II-;--

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 81 06

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/mm 2 )

600 'ZQ

80

i I .... rm = 3 .6

'" 008-Q

I 9~ t c:::J -<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 6 810 7

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

A33

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 (AS ME)

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

A34

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

, 'n~:: ,IJ1:1-: rit"- I ·HI Olm

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

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 I 0 z:;

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 21

- s"' r---...r--.. 1---1"-.......... cru

200

R = 0.1

N

4 6 810'" 2 6

4 6 810 2

Fig.l0 Results ofthe experimental investigations on butt weldedhollow sections

A35

s .. [N/ mm 2 ]

1000 -ZQ.Q...

i i ...... I I B~ t 2, t lZOO

800

600 transverse attachments , , RHS 100 x 100 x 4,0 , - Ilm = 4.6 J~

, attachment8 100 x 60 x 6,0 , StE 600 TM , "- ,

, ""-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

A36

BEHAVIOR OF HOLLOW SECTION COLUMN-TO-H BEAM CONNECTIONS IN MOMENT RESISTING STEEL FRAMES

UNDER HORIZONT AL LOADING

Yoshinori Matsubara, Hiroshi Kanatani, Mototsugu Tabuchi and Teruyasu Kamba

Department of Architecture, Faculty of Engineering, Kobe University, Rokkodai, N ada, Kobe, Japan

Summary

This paper describes the effect of the corner radius of cold forming SHS columns on the

behavior of joint panels in moment resisting steel frames under lateral force. To investigate

the behavior of joint panels, tests on sub-assemblages consisting of as-pressed or stress-relieved

SHS columns with various extern al corner radii are carried out.

The external corner radii of SHSs used in tests are R=O, 2T, 4T and 8T.

From these studies, empirical formulae for predicting strengths of joint panels are proposed.

1. Introduction

In moment resisting steel frames under lateral force, high shear stress occurs in joint panels of

column-to-beam connections.

The behavior of joint panels with square hollow section (SHS) column-to-H beam connections

have been reported by the authors [1], [2]. In those papers, the design formulae obtained

from test results of SHS columns with R:::;: 2T (R:external corner radius, T:wall thickness of

SHS).

However, the behavior of joint panels of cold forming SHS will be affected by the size of the

corner radius, because cross sectional properties and mechanical properties are varied with the

corner radius.

In this paper, experirnental investigations on sub-assemblages of moment resisting frames

consisting of as-pressed or stress-relieved SHS columns with various external corner radii are

presented.

The external corner radii of SHSs used in tests are R=2T, 4T and 8T. A test with a box

section built-up by four steel plates which is considered as R=O is also carried out as the

extreme case.

2. Test

Specimen Test specimens and test setup are shown in Fig.I. Beams are rolled H section H-400 X 200 X

8 X 13 and columns are cold-rolled RHS 0-300 X 300 X 12. Four types of SHS are used as

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+0 I , I

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

I i. 1 900 , 1900 • i

I Designation of Specimen

, ~ , FI- - - - - - - - - - - - I - - - - - - - - - - - -13 , I ,

IBxt2~Fh, .. ;IT ~:Stress 8ehevmg

DIT , -r,- '-t- . 25 25

Fig.1. Test specimen and test setup.

D-300mm

DF-300mm 8 8 I ,

DF-252 .8 OE-204 , I

BX25RO

OF : Width ol FIa! pan

joint panels shown in Fig.2.

BX25R2SR

tr=12

BX25R4 BX25R4SR

t:r=12

Fig.2. Sectional properties of SHS.

8 ,oF= 1 08-B-.

Table I shows cross sectional details of the SHSs used in the joint panels, that is, outer radius

of a corner, R, cross sectional area, A, and ratio of width of flat portion 10 external width of

SHS, DF/D. In Table I, manufacturing proccss of the SHSs are also shown. Ratios of the

outer radius of a corner to wall thickness, Rrr, are 0, 2,4 and 8. Here, Rrr=ü rneans SHS

built-up by four steel plates. Tbe SHS with R{f=2 is cold-rolled tube and the ones with

R!f=4 and 8 are made by cold-pressing. Three SHSs with "SR" in specimen designation are

subjected to heat treatment. Tbe heating condition for stress-relieving was keeping the SHSs

at 620'C for two hours and then cooling gradually in the furnace.

External width (=300mrn) and wall thickness (=12mrn) of the joint panels are common to all

specimens.

Cross sectional area of the joint panels, A, becomes smaller as R{f increases. Tbe NARa ratios oftbejoint panels with R{f=2, 4 and 8 are 0.97, 0.94 and 0.87 respectively, wh ere ARa

is tbe cross sectional area of the SHS with Rrr=O.

Tbe SHSs in specimens BX25RO, BX25R4, BX25R8 and BX25R8SR are made of tbe same

steel plate.

Built-up and cold-pressed SHSs are produced by sub-merged are welding.

Tbe connections are stiffened by a couple of through diaphragms at tbe beam flange levels.

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Table 1 Details of SHS.

Specimen Size Manufacturing

process SR RIT A(an~) DF/D BX25RO Built-uD 0 138.2 1.00 BX25R4 Cold-Pressed - 4 129.6 0.68 BX25R8 0-300 X300 X 12 Cold-Pressed 8 119.7 0.36 BX25R2SR Cold-Rolled

Stress-2 134.5 0.84

BX25R4SR Cold-Pressed 4 129.6 0.68 BX25R8SR Cold-Pressed

relieving 8 119.7 0.36

R: Extemal corner radus T: wan thickness A: Cross Sectbnal area OF: Wdth of flat part 0: Extemal wK:Ith

The beam and the colurrm are proportioned to avoid premature failure of the members prior

to shear failure in the joint panels.

Test procedure and rneasurernent The test setup is shown in Fig.l. Every specimen was applied unsymmetricalloads

monotonously at the ends of the beams. Arrangements of displacement transducers and

strain rosette gauges to mcasure shear distortions of joint panels are shown in Fig.3.

D-300.,.,

DF=300""" B DF-204 B , I I , R DF=108 R

I ,

A BX25RO BX25R4

t:. t:. t:. t:. t:. 3060,60,60,60,30 ,51,51,51,51, ~

Fig.3 Arrangement of transducer displacement and strain gauges. (A-A' section)

Mechanica} properties of SHS Mechanica! properties of the SHSs are listed in Table 2.

Grade of SHS for BX25R2SR is STKR400 and that for the others is SS400 according to JIS.

Tensile test pieces along longitudinal direction of the SHSs were cut out from flat parts and

corner parts. T abIe 2 Material properties of SHS

Y : Yleld ratio t: u : Uniform elongation EL: Fracture ebngaten • : Meam values 0/ the corner part ~ained trom twelve test pieces (point 1-4)

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- --~--~-

Number 5 and No.12 test pieces according to JIS Z 2201 are coupons with whole thickness and

NO.14 test piece is small size coupon shown in Fig.4.

Figures 5(a) and 5(b) show locations of the corner part cut out the NO.14 test pieces. In specimens R4SR, R8 and R8SR, the test pieces were taken from four points (1-4) in the

corner shown in Fig.5(b). At each corner point, the test pieces are cut out from three points

along thickness direction, that is, outside (0), middle (M) and inside (I) of wall thickness.

F 1

= j 3.5 7

(a) BX25R2SR and BX25R4. (b) BX25R4SR, BX25R8 W-1-4

Fig.4 JIS NO.14 Test Piece. and BX25R8SR.

Fig.5 Locations of JIS NO.14 Test Pieces.

Figure 6 shows stress-strain rclationships of the flat part (by NO.5 test piece) and the corner

parts (by NO.12 test piece) of specimens BX25R4 and BX25R8. The (J-E curve of the flat

part showed clear yield point and plateau. On the other hand, the (J-E curves of the corners

showed a kind of round house type and gradually yielding-plateau-strain hardening type

without a sharp knee. The 0.2% off-sct yield stress was adopted for SHS which shows round

house type (J-E curve.

--- Flat part 21----+ ----- Corner part (BX25R4)

----- Corner part (BX25R8)

L-__ ~ ____ ~ ____ ~ ____ ~~(%)

o 5 10 15 20 Fig.6 Stress-strain relationships .

cr y (1fI cm2)

3.5

4.5

4.0

RIT = al 0 as-pressinQ I • stress-reli8\ling

0 L~

0 .I1.

• 3.0 I

• - I~

2 3 4

Fig.7 Distribution of yield stress

at the corner part.

Figure 7 shows the yield stress ay at each point of corner part (1-4) and flat part (F) of SHSs

with R!f=8, wh ere ay is mean value of three points along thickness direction (0, M and n. At any points along circumferential direction the corner, the yield stresses of stress-relieved SHS are lower than those of as-pressed SHS.

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Table 3 shows the effect of stress-relieving on SHSs with R!f=4 and 8. On the corner part

the cryAP/crySR ratios of SHSs with R!f=4 and 8 are 1.19 and 1.17 respectively.

Tab1e 4 shows the ratios of corner part to flat part in yield stress, cryc/cryf, and in tensi1e

strength, cruc/cruf. The cryc/cryf ratio of the as-pressed SHSs with R{f=4 and 8 is about 1.3.

And the ratio of the stress-relieved SHS decrease to about l.I.

However, the cryc/cryf ratio of stress-relieved SHS with R!f=2 are 1.23.

Table 4 Comparison between Table 3 Effect of stress-relieving on material properties. the comer part and the flat part.

Rat Corner ayd a.,-i (Jur!(Juf

R2SR 1.23 1.15

RfT :::A . .......... L.9.9L. ..... J.:.9? ....... J:J.~.L. ..... J . .J§ "Rif =8 0.99: 1.01 1.17' 1.04

R4 1.27 1.15 R4SR 1.07 1.01 R8 1.29 1.06 flsSFf ···········fö·g ··········fÖ·2

3. TEST RESULTS

Load-deflection curves are shown in Fig.8, where load M is represented in bcnding moment at

column surface. The deflection e is equal to o/Lo, wh ere 0 is an overall deflection

measured at Ioading points of the beam ends and Lo is a beam length.

Every specimen shows excessive shear deformation of joint panel and load-deflection curves

are very stabIe. Tests were terminated without fmding the maximum load carrying

capacities of the specimens when the deflection e exceeds 0.I3rad.

50 M (!fm) 50 M (!fm)

BX25RO 1 -- .. -... -.-- - BX25R2SR ' -------. BX25R4 I - - - - BX25R4SR - '-BX25R8 1-

IX2~r-~8 -" L-~ __ -L __ -L __ ~ __ ~~ __ -J y(xl0 2rad )

10 12 14 16 10 12 u. e (x l O·2rad.)

Fig.8 M- () curves. Fig.9 M- 'Y curves.

Figure 9 shows the load versus shear deformation relationships of joint panels. The

behaviors of the M-y curves are very similar to those of the M-e curves because the shear

deformation is dominant in this test.

Table 5 shows the test results.

Here, My = Yield load (the load when stiffness of M-e curve becomes one third of the

initial stiffness)

Mu = Maximum load (the load when the shear deformation g becomes 0.11 rad.)

Mmax = Maximum strength (the maximum load observed in test)

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Table 5 Test resuls and calculated values Specimen M~lfml : MJlfml : MmaJttn) MulMy Mpyltfm) : MpJtfm) My/~y ! MjMpu Mpy<tfm): MpJtfml M./~y : Mu/~u BX25RO 22.8! 44.1! 44.1 1.94 23.6: 41.0 0.96: 1.08 23.6 41.0 0.96: 1.08 BX25R2SR 28.0! 47.T 47.7 1.70 26.0: 41.2 1.08: 1.16 - - - -BX25R4 22.6: 45.2 ~ 45.7 2.00 23.6: 41.0 0.96 ~ 1.10 - - - -BX25R4SR 20.9 ~ 40.9: 42.8 1.95 23.4' 40.5 0.89 ~ 1.01 22.1 : 37.7 0.95 1.08 BX25R8 24.5' 46.5: 46.7 1.90 23.6: 41.0 1.04: 1.13 - - - -BX25R8SR 20.1 : 39$ 42.0 1.98 24. 0 ~ 40.7 0.84 ~ 0.98 20.7 35.3 0.9T 1.13

My : Yeld lead Mu : Maximum lead Mmax : Maximum strength Mpy : Shearyield strength of joint panel caadaled with eq.(1) Mpu : Shear ultimata strength of joint panel cab.Jlated with eq.(2) Mpy : Shearyield strength of joint panel cab.Jlated with eq.(3) Mpu : Shear ultimata strength of joint panel cab.Jlated with eq.(4)

4. DISCUS SION

Shear stress in joint panels Figures lO(a)~ 10(c) show shear stress distributions along circumferential direction in joint

panels (see Fig.3), where thick lines and fine lines show the experimental results and analytical

results based on the theory of shear flow respcctively . The shear stresses in these figures are

represented by components of the stresses against shear force of joint panel. The

experimental results agree with the analytical results. It can be observed that the corner

parts carry a certain amount of the shear force and the ratio of the shear force canied by the

corner part to that of the whole joint panel increases with the increase of Rff ratio.

::rl:~:~' mml :::rl(·~~/;;:lf.-$-f-+-~·~~~ m -I 0.00- . L:± ·J. . u (cm) 0.00' ! _ - - ~ ! • U (cm)

.30 ·20 ·10 0 10 · 20 30 -30 ·20 -10 0 10 20 30

I --: Analysis I _ : Experiment :::ll'IB~lu 1=1

0.0930-. .,.20'--....... 1 0--0~-1:'::O-~20~-='30

Fig.10 Shear stress distributions.

Figures 11(a) and 11(b) show the average shear stress versus the shear defonnation

relationships in joint panels. The ordinate and the abscissa of these figures are indicated by

the nondirnensional quantity 't/'ty and y/yY respectively, where 'ty (=cry/F 3) and yy (='ty/G) are

shear yield stress and shear yield strain of material at flat part of the SHS respectively.

In all specimens, yielding occurs at 't/'ty<1 .0. At maximum loads, the 't/'ty ratio and the Y/YY

ratio are exceed 1.5 and 50 respectively.

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Figure 11(a) shows the effect of the cross sectional properties of the corners in SHS on t-r curves, because all specimens in this figure are the built-up or the stress-relieved SHSs which

have little effect of work-hardening. The 't/ty ratio decrease with the increase in RIf ratio

at the same rfyy ratio, except for roll-fonned stress-relieved SHS (BX25R2SR).

Figure 11 (b) is the comparison of the 't-r curves of as-pressed SHS with those of stress-

relieved SHS. This figure shows the effect of the mechanical properties of the corners in

SHS. The general yieldings of the joint panels with stress-relieved SHSs occur earlier than

those of the joint panels with as-pressed SHSs. After yielding, the t/ty ratios of stress-

relieved SHSs are larger than those of as-pressed SHSs at the same r/'YY ratio.

1.5 .. ~ s---

:2 ::.-;:-- -~-~ ;.::--:...:::-.

.:::-.-.

~, .. ~.~ :.--. :/.

~ -8X25RO ';/ ·········8X25R2SR ii - - - - 8X25R4SR

.5 _.- 8X25R8SR

I 1 10 20 30 ..0 50 60

(a) Effect of sectional properties. Y/Yy

----8X25R8 -8X25R4 ----·8X25R4SR

.5 It----t---t -.- 8X25R8SR --+---1 I I

! C 10 20 JO 40 50 y/ . 60

(b)Effect of mechanica I properties ~ at the corner parts.

Fig.11 Nondimensional .. - y relationships.

Strength of joint panels In the Previous paper [1, 2], the authors proposed the equations for predicting shear yield

strength and shear ultimate strength of joint panels expressed in eq.(l) and eq.(2), where the

strengths are represented in the bending moment at column surface. These equations predict

the strength of joint panels used rolled-fonned SHS with RIf=2.

Mpy=.8.~~ Vp 9 ~ 1-A-/1 2 (1 )

M _~~Vp pu- Y3 l-A-/12

Vp=2 DaDe T

Here, Da =Distance between beam flange

De =Distance bet ween column flange

T =Wall thickness of joint panel

A =Dc/l

/1 =DB/h

DY = Yield stress of flat part of joint panel

au =Tensile strength of flat part of joint panel

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(2)

I

Equations (1) and (2) give tbe same yield strength MPy and tbe ultimate strength MPu for tbe

connections witb tbe same material properties on flat part of SHS and tbe size of SHS (DXT)

even if tbe outer radius of tbe corner changes.

The comparison of observed yield load My and maximum load Mu witb tbe predicted values

by eq.(l) and eq.(2) are shown in Table 5.

Yield strength of specimens with tbe built-up SHS (BX25RO), tbe as-rolled stress-relieved SHS

(BX25R2SR) and tbe as-pressed SHSs (BX25R4 and BX25R8) can he predicted by eq.(1).

On the otber hand, the My!Mpy ratio of stress-relieved SHS with R{f=4 and 8 are 0.89 and

0.84, respectively. This means that eq.(l) may ovcrestimate the yield strength of tbe joint

panel with stress-relieved SHS.

Figure l2(a) shows the relation between My!MPY and nondimcnsional cross sectional area of

SHS NARO. Figure 12(b) is the relation between Mu!Mpu and A/ARO.

My/Mpy Mu/Mpu 1.1 r---.---~---, 1.2 ,...---,----,----,

o

o

1.0 ............... ·· ····· ·· ·······1···············

0.9 ............... ···············!"···t:r·······

0:

1.1 r----+--~-___I

1.0 ....... ...... ... ..... .... ..... ~ ... . t:l ..... .. . 0 :

0 .8 L..-_--'-__ "'--_-' 1 0.9 '-----'-~-'--~-' 0 .7 0 .8 0.7 0.8 0.9 AI Ap,o 1.0

(a) Yield load. (b) Maxirrum load.

Fig.12 Relationships between strength and cross sectional area of SHS.

Except for a specimen with RIT=2 (rolled SHS), it is obvious that the My/MPY ratio and the

Mu/MPu ratio decrease proportionally with the A/ARo ratio and these inclinations are about

1.0. This implies that it is necessary to consider thc dccrease of the cross sectional area of

SHS with the increase in RIT ratio for estimating the strcngth of the joint panel with stress­

relieved SHS which have almost equal mechanical properties in the flat part and the corner

part. Therefore, it was tricd to estimate the strengths of joint panels with stress-relieved

SHSs by eq.(3) and eq.(4) substituting A· DD/2 as panel volume Vp in eq.(l) and eq.(2).

Kr -.8. ay ~ Vp PY- g Y31-À-jl 2

-,:;r- - au ~ Vp PU - Y3 1-À-jl 2

Vp= A .08/2

(3)

(4)

For the specimens BX25R4SR and BX25R8SR, the calculated results Mpy and Mpu by eq.(3)

and eq.(4) are shown in Table 5. The My!MPY ratios are 0.95, 0.97 and the Mu!MPU ratios

are 1.08, 1.13, respectively. These values are comparabie to those of BX25RO which ~e 0.96 for My/MPY and 1.08 for MufMpu.

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The strengths of specimens with as-pressed SHS are estimated by eq.(1) and eq.(2) as stated

before, because the effect of the decrease in cross sectional area of SHS by increasing in the

corner radius can be offset by increase of material strength at the corner part subjected to

work-hardening.

5. Conclusion

Based on these experimental investigations, the following conclusions for the effect of the

corner radius of SHS on the behavior of joint panels of SHS could be drawn.

1) The corner parts of the joint panels carry the shear stresses and the ratio of the shear force

carried in the corner parts to that of the whole joint panel increases with the increase of

Rff ratio.

2) In the built-up and the stress-relieved SHSs which have little effect of work-hardening the

My/Mpy ratio and Mu/Mpu ratio decrease proportionally with the AI ARO ratio and these

inclinations are about 1.0.

3) Strengths of the joint panels with the as-pressed SHSs are larger than those of the stress­

relieved ones because of work-hardening of the corner parts.

4) Strengths of the joint panels with little effect of work-hardening in the corner parts can he

estimated by eq.(3) and eq.(4) which consider the sectional properties. On the other

hand, the joint panels with the effect of work-hardening in the corner parts can be

estimated by eq.(l) and eq.(2) because the effect of work-hardening at the corner parts

canceled the decrease in cross sectional area of SHS.

Reference [1] Tabuchi, M., Kanatani, H. and Kamba, T. (1988). "Behavior of Tubular Column to H

Beam Connections under Seismic Loading". Proceedings of 9th WCEE, August 1988.

[2] Kanatani, H., Tabuchi, M. and Kamba, T. (1990). "State of the Art of Tubular Column­

to-H beam Connection in Moment Resisting Frames Subjected to Earthquake Loading".

IIW Doc. XV-E-90-159, July 1990.

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