bose and hughes 1995

Verifying the performance of standard ductile connections for semi-continuous steel frames B. Bose, BSc(Eng), PhD, CEng, MIStructE, and A . F. Hughes, MA, CEng, MICE, MIStructE

W It is proposed to revive the concept of semi-continuous design of steel frames. A set of predesigned standard ductile connections has been developed by the Steel

believe that the recognized theoretical Construction Institute. The authors

benefits of what was known traditionally as semi-rigid design can be unlocked by the application of plastic global analysis to frames with the standard connections. The properties of representative standard connections have been verified by full-

in this paper. scale destructive tests which are reported

Semi-continuous design is an umbrella term Introduction

which has been adopted by Eurocode 3.’ It embraces both semi-rigid (elastic) and partial strength (plastic) approaches to the analysis of moment resisting frames. Previous codes such as BS 5950’ have obscured the distinction between these approaches which are fundamen- tally different even though the connections themselves may well be both semi-rigid and partial strength.

are to save money andtor reduce member size. 2. The motives for semi-continuous design

If the frame is unbraced, moment connections are unavoidable but these can be significantly less elaborate, and less costly, than with the continuous (‘rigid ’) approach. If the frame is braced, the semi-continuous approach competes with simple construction; it offers lighter and shallower beams in return for marginally more elaborate connections.

struction are well known’ and have been recog- 3. The advantages of semi-continuous con-

nized for the greater part of this century, but, for one reason or another, design practice has tended to overlook them. Perhaps the main reason is the enduring attraction of simple construction, which is easy for the designer (one member is considered a t a time). suits the way the industry is organized (the fabricator is responsible for the connections) and results in economical structures. Although design standards outwardly encourage what has traditionally been labelled ‘semi-rigid construction’.

as unacceptably involved. the profession has perceived the design process

can and should change. The benefits of semi- 4. The authors contend that this situation

continuous design can be unlocked by the combination of

(a ) plastic global analysis ( b ) standardized ductile connections.

which are capable of acting as plastic hinges. In Eurocode 3 terms, they possess ‘rotation capacity’. ‘Ductile’is not an official Eurocode 3 term but aptly describes the property which a partial strength connection must possess if it i s to perform satisfactorily in a plastically analysed frame.

context is that it offers the designer control of 6. The great virtue of plastic design in this

the bending moment diagram. Determinacy is second only to that of simple construction. By contrast, elastic analysis (the semi-rigid approach) suffers from the practical difficulty of predicting connection stiffness. It is, in the authors’ view, not a realistic (or even safe) method for the ‘strength design’ of practical structures.

can be calculated with relative ease, but there is 7. The moment resistance of a connection

a procedural difficulty to overcome. The con-

things remaining equal) on the beam depth, nection moment resistance depends (other

while the beam size required is inversely dependent on the connection moment. The two need to be chosen simultaneously. Standardized connections, whose moment resistance is tabulated by beam size, eliminate the trial and error which would otherwise be necessary. These predesigned connections can be relied on to be ductile without further calculation.

‘strength design’ is not enough. Connection flexibility exerts an important influence on serviceability (lateral deflection) and also on sta- bility, although the latter is relatively unlikely to govern design of low-to-medium-rise frames. For the serviceability calculation, the designer may employ the semi-rigid elastic analysis which was mentioned, somewhat dismissively, above. This requires a ‘best guess’ of the rota-

further advantage of standardized connections tional stiffness of the connections, and one

ought, in principle, to be available. is that more reliable semi-empirical predictions

5. Ductile connections are connections

8. In the case of an unbraced frame,

Proc. lnstn Civ. Engrs Structs & Bldgs. 1995,110, Nov., 441 -451

Structural and Building Board Structural Panel Paper 1089 7

closes IdJanuary 1996 Written discussion

B. Bose, Lecturer,

Abertay Dundee University of

A . F. HuRhes, Associate. Ove Amp &Partners -

441

BOSE AND HUGHES

9. A range of standard ductile connections has been developed at the Steel Construction Institute, in collaboration with Warwick Uni- versity and the SCI/BCSA Connections Group which acts as a focus for the UK industry’s activity in this area. These have been designed with reference to Eurocode 3 Annex J, which is itself informed by many years’ Dutch testing and experience. Nevertheless, it was judged desirable to submit some representative connections from the range to experimental verification of their performance, especially in the crucial respect of ductility. The tests, which are reported and interpreted here, were carried out by the Structures Research Group of the University of Abertay D ~ n d e e . ~ . ~

Design of ductile connections

the often competing demands for

0 strength (moment resistance) 0 rigidity (rotational stiffness) 0 ductility (rotation capacity) 0 economy (freedom from stiffeners or other

10. Practical connection design must juggle

welded attachments).

11. For the present purposes, ductility cannot be compromised. Economy is also para- mount; it would be all too easy, by increasing the workmanship content at the connections, to negate the advantages which motivated the choice of semi-continuous construction in the first place. Strength must be maximized within these constraints. Similarly, rigidity, which is often, in unbraced frames, a controlling influence on the design, should not be sacrificed any more than necessary.

12. Eurocode 3 Annex J gives design rules for bolted end plate connections, both flush and extended. Other styles of connection may also perform in a ductile manner but end plate connections are, by common consent, the most suitable candidates. The rules, although simple in principle, are rather complicated to apply, which is another argument in favour of the standardized approach.

13. Essentially, for the connection to be ductile, it is required that its resistance should be governed by one of a small number of components which can be relied on to deform plastically, in a benign way. Other components which could fail in relatively brittle fashion, notably the bolts and the welds, must not be allowed to become the weak link in the chain.

14. There are three ‘ benign ’ failure mecha- nisms.

( U ) bending in the end plate around the tension

( b ) bending in the column flange (c) shearing in the column web panel.

bolts

15. The first of these is the basis for the design of the standard ductile connections since

it is the only one which is always available. (Column flange thickness is usually determined in advance of connection design; web panel shear may be cancelled out by two mutually opposing beams on either side of the column.)

16. For any given connection geometry, there is a limit to the acceptable end plate thickness relative to the bolt size and strength. Because prying action inescapably accompanies thin end plates, effective bolt tension is limited to around 70% of resistance in pure tension. This is why ductile connections are relatively inefficient in strength terms, and could never be promoted as all-purpose standard connections.

17. It is advantageous for stiffness, and no disadvantage in other respects, to adopt the most compact’ practical geometry, keeping the bolts close to the webs, beam flange and stiffeners, if any. For the same reason, there is a preference for larger and stronger bolts: M24 8.8 bolts are preferred, although M20 8.8 bolts (the ‘industry standard’ bolts for simple construction) may suffice in some cases.

Standard ductile connections 18. Connection standardization has been

identified as an important component in the campaign to deliver efficiency gains at all stages of the steelwork design and production process. The arguments are well rehearsed6 and widely accepted. They apply as much to moment connections as to any other. For the special case of ductile moment connections, the merits of standardization go beyond this. They simplify the design process at two levels.

19. Firstly, frame design is facilitated because the moment resistance of the standard connections is known (from tables) in advance. The designer is in a position to make an informed choice of member size and connection type, together. Much trial and error is elimi- nated, or at least made manageable. One of the virtues of the semi-continuous approach is the freedom it gives the designer to trade connection capability against member capability, so it can hardly be expected to match the directness of ‘simple construction’. What can be claimed is that standard connections make it a painless practical alternative for the designer.

20. The second level is that of the detailing of the connection itself. Standard connections are predesigned-there is little or nothing for the user to check. This is an obvious advantage. Manual design of a ductile moment connection is not to be undertaken lightly; software for code-checking is available or on the way, but programs that actually design an ‘optimized’ connection are a more distant pro- spect.

21. The range of standard ductile connection details developed at the Steel Construction Institute was based originally around five bolt

442

SEMI-CONTINUOUS DESIGN OF STEEL FRAMES

W1 Tic "ttt" Optional extra row for shear

(250)

W3 (3 +) 7r 4T

W5 (5 +) "?r 4w

W2 (2 +) ".lr 4w

I

Optional extra row for shear

I l l

1

I W4 (4 +)

configurations, with either flush or extended end plates in two standard widths. Bolts are M24 or M20, 8.8. Fig. 1 illustrates the range for M24 bolts; the geometry is the same for M20 bolts, except that end plate thickness is one size down, e.g. 12 mm in place of 15 mm.

22. The standard connections have been designed with reference to Eurocode 3 Annex J, but the rules have been adapted for use with

Notes

2. Optional extra bottom bolt rows for shear 1. '+' denotes the wider (250) end plate.

3. Flange to end plate weld size to be in the

4. Web to end plate weld 2 X 8 FW as

5. End plates S275 6. M24 8.8 bolts as standard. 7. Prefix W indicates connections suitable

for the wind-moment method, the bottom half of the detail mirroring the top. This referencing system has been changed in the Design Guide.'

are shown on 1, 2 and 4.

range 10-12 mm visible fillet.

standard.

BS 5950. The ' strength design ' follows the procedure adopted in the SCI/BCSA Connections Group's Design Guide for Moment Connec- tions.' This is somewhat conservative relative to Eurocode 3 in its bolt tension values.

Eurocode 3 Annex J prescribes that a connection whose end plate deforms in ' Mode 1 ' (essentially, in double bending) may be regard-

23. As far as rotation capacity is concerned,

Fig. 1 . Range of standard details for M24 bolts (all dimensions in mm)

443

BOSE AND HUGHES

ed as ductile. Unfortunately, this is rather restrictive; it would, if applied to the standard details, limit the end plate thickness to no more than half the bolt diameter, i.e. 10 mm or 12 mm. While such connections would undoubtedly be admirably ductile, their rotation capacity would have been achieved at sig- nificant cost in both strength and stiffness. It was felt in the Connections Group that slightly thicker end plates would remain adequately ductile and represent a more acceptable balance between the competing demands made on the connection. It should be emphasized that this view is not in defiance of Eurocode 3, which rules in ' Mode 1 ' connections but does not rule out those which are not. Indeed, current proposals' for the much-needed revision of Annex J contain a reformulated design rule which will admit the geometry of these standard details, in which end plate thickness is 60% of bolt diameter.

Verification of standard details 24. Although ductile connections have,

knowingly or otherwise, been used for many years in frames designed to the ' wind-moment ' m e t h ~ d , ~ there is a degree of novelty attached to the concept in most designers' minds. There is also a certain suspicion, not always justified, of Eurocode 3.

25. If the hearts and minds of these designers are to be won over to the idea of semi-continuous design, and if standard connections are to be promoted as the ' tool kit' for its application, it is reasonable to expect some demonstration that they perform as claimed. This is one motive for the set of tests reported here. Another is the deliberate overstepping of the Mode 1/Mode 2 boundary by a small margin-best justified by experimental evidence. It is also of interest to compare strength predictions according to the Design guide's' 'hybrid EC 3/BS 5950' approach with test results. Finally, the rotational stiffness of the standard details is of interest, as reliable quan- tification of their flexibility is a prerequisite for rigorous serviceability checking of unbraced frames.

26. Before the present series of tests was commissioned, an attempt was made to dis- cover published test data for similar connections. Although much research effort has gone into so-called semi-rigid approaches of recent years, including the establishment of a com- puterized data bank of test results," disap- pointingly little of this work was found to be directly relevant. It may be no exaggeration to say that many investigators have concluded tests (on grounds of 'excessive end plate deformation ' perhaps) at rotations which were just beginning to be of interest. Dutch results reported by Zoetemeijer" and others did contain comparable connections, but in no case

did the materials and geometry match those now proposed as standards.

27. The scope of the test programme at Dundee in autumn 19934 was dictated by the funding available. Tests were chosen to represent typical combinations of member size and connection style, but some emphasis was given to deeper beams. This was in recognition that (other things remaining equal) the amount of yielding demanded at the tension and compression zones is proportional to the lever arm. A further series of tests, including three non- standard details (with end plate thickness variations), was undertaken in 1994.'

28. All the tests were on cruciform specimens, loaded symmetrically. This represents a typical beam-to-column connection under gravity moment. The web of the column is not subject to shear, which means that it cannot enhance rotation capacity by this means. However, another consequence of the symmetry of the test is that this component is a. non- contributor of flexibility. In other words, the same connection, loaded differently, would behave less rigidly. Appropriate compensation must be made before rotational stiffnesses measured in symmetrical tests are used in serviceability calculations for unbraced frames.

29. Adding an extra bolt row to a configuration which is already ductile ought to improve matters in this respect, since the effective length of equivalent T-stub per bolt row is reduced. For this reason, the tests concentrated on details with not more than one bolt row below the flange, i.e. W1, W2 and W4. These are respectively

W1 : Flush end plate with single tensile bolt

W2: Extended end plate with single bolt row

W4: Extended end plate with two tensile bolt

row

(in extension)

rows.

Initially, only connections using M24 8.8 bolts were tested, since this bolt size will normally be the first choice. Two connections using M20 8.8 bolts were included in the second series of tests.

30. End plates were 15 mm thick with M24 bolts, and 12 mm thick with M20 bolts, with a few exceptions. For detail W2, a 20 mm thick end plate is specified with M24 bolts. This detail has bolts in the extension only and the effective length of equivalent T-stub is limited by the breadth of the end plate. A ductile connection should therefore be achievable with a thicker end plate. As with the other standard details, the 20 mm thickness is a little greater than that which would correspond to the divid- ing line between the two modes of deformation.

31. The other test which featured a 20 mm thick end plate is test 12, whose geometry corresponds with standard detail W5 in all other

444


respects. The test was included to obtain some

and stiffness is gained a s the thickness of the indication of the rate at which ductility is lost

end plate is stepped up. For the same reason, one of the tests using M20 bolts (test 10) was given a 15 mm thick end plate, one size up from the standard for M20 bolts in detail W4. This provides a comparison with test 4, which differs only in having M24 bolts. Test 11 repeats test 3 (standard detail W2). this time with an end plate one size down, i.e. the same thickness as in the other standard details with M24 bolts.

32. The remaining nine tests were on specimens which correspond to the proposed standard details, and eight of them used M24 bolts. Detail W1 was tested three times, W2 was

explore fully the influence of the other vari- tested once and W4 was tested five times. To

ables involved-such as beam depth, column size and column mass-would require resources well beyond those available. Choices had to be made, and one was that all columns would be 254UC. Indeed, nearly all were 254UC89. One (test 6) was 254UC73, and two 254UC132 tests were included as second series tests 7 and 8 (otherwise identical to tests 4 and

range 406 X 178UB to 762 X 267UB; the size 5 respectively). The beam size was varied in the

most often chosen was 457 X 191UB74.

of end plates which are S275 (i.e. grade 43 in the old system), even if the members they connect are S355. It cannot be overemphasized that substitution of S355 end plates (or other excessively overstrength material) would set at risk the ductility of the connection. For the tests, all beams and columns were also S275. Measured material strengths are shown in Table 1. For the most part the material was found to be no more than ‘averagely’ overstrength, but an exception is the 12 mm thick end plate (used in test 9).

Testing of the connections Test programme

34. In order to verify the performance of the standard connections. a programme of tests

connections was planned, which covered the involving nine standard and three non-standard

following variables (a) various beam depths

(c) various column sizes ( 6 ) various connection details

(d) two bolt sizes (M20 and M24 bolts, both 8.8 fully threaded).

33. The standard detail are based on the use

35. Destructive tests of standard cruciform beam-to-column connections were performed in the Heavy Structures Laboratory of the Uni- versity of Abertay Dundee. Details of the test specimens are given in Fig. 2 and Tables 1 and 2.

A+ 25 mm top plate

,on both sides 12 mm sMfenen

normal commercial channels and no attempt was made to control material or workmanship,

test specimens

except that punched (not drilled) holes were specified. Bolts were normally tightened (i.e. not preloaded).

37. Weld contraction tends to induce con- vexity in the end plate (viewed end-on). although the effect can be minimized by good fabrication technique. The second series specimens came from a different fabricator, and this curvature was particularly noticeable. It is unlikely that this ‘initial imperfection’ has much effect on the ultimate resistance of the connection, but it may exert some influence on the ductility and rotational stiffness.

38. The loading and instrumentation set up

The load was applied by a 1000 kN capacity for the full-scale tests is shown in Figs 3 and 4.

hydraulic jack and monitored by a load cell of similar capacity. Joint rotation is defined a s the change of angle between column and beam centre-lines. Rotation was measured separately on each side of the column by two independent Fig. 3. General view means, a pair of dial gauges and a pair of dis- of test set-up

36. The specimens were procured through Fig. 2. Geometry of

Table 1 . Material properties

* Premature failure by thread stripping.

placement transducers. The load cell and the displacement transducers were connected to a datalogger. A steel bar of square section was connected rigidly to the column web at position A, and a steel angle section was attached to each of the beam webs at B, as shown in Fig. 5. Point A was located in the column at the inter- section of column and beam centre-lines, while point B was positioned on the beam centre-line very close to the welded end plate. Two dial gauges mounted on magnetic stands were set up on the beam attachment at 300 mm distance apart, with their pointers resting on the column attachment. A pair of displacement transducers on magnetic stands was supported on a rigid steel frame, their pointers resting on the beam attachment at 300 mm distance apart. The steel frame was fixed to the laboratory floor at a

short distance in front of the test specimen, a s shown in Fig. 4. Similar arrangements were made on either side of the column to measure the joint rotation. At any applied load, the difference between the two dial gauge readings or two displacement transducer readings in mm divided by 300 mm represented the rotation of the connection. The moment acting at the connection corresponding to an applied load is calculated by multiplying the support reaction (half of the applied load) by the distance between the roller support and the face of the column flange (2400 mm for tests 5 and 8, and 1200 mm for others).

Test results

tested connections are given in the two Consul- 39. Detailed test results for the twelve


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BOSE AND HUGHES

Ductility

dard details, seven demonstrated respectable 42. Of the nine tests on the proposed stan-

rotation capacity. Two tests, namely tests 5 and 8, failed to do so: test 5 could be judged borderline, but test 8 could not. Test 11, non-standard because of a thinner end plate, performed satisfactorily. The remaining two non-standard tests, both with end plates one size too thick, displayed-not unexpectedly-low rotation capacity.

Before attention is focused on the individual 43. The ductile/non-ductile distinction.

test results, it is appropriate to discuss where the borderline between ductile and non-ductile should be drawn. The rotation capacity required in practice will vary with the circumstances in which the connection finds itself. Imposition of a black/white distinction at some arbitrary shade of grey is, like the rigid/semi-

Fig. 4 . Close up uierv tancy Reports4.’ prepared for the Steel Con- rigid distinction, for convenience of design of test set-up struction Institute. A summary of important rather than an absolute structural imperative. It

test results is reported in Table 2. The moment is worth remarking that Euro,.ode 3 is silent on resistance, rotation capacity and failure mode the subject of what 9 rotation capac. of each connection are given. All the tests were ity in to its outspokenness conducted until failure, which occurred for a (relative to BS 5950 anyway) on the numerical variety of reasons: column web buckling, bolt treatment of the ‘rigid’ limit, stripping, bolt fracture and end plate fracture along weld toe. As mentioned PrevioUslY, two tion is that it should be capable of rotating, a s a

44. The requirement for a ductile connec-

sets of readings were recorded to compute con- plastic hinge, as much as the rest of the Stmc- nection rotation at each side of the column. It ture requires and That will depend on was found that rotations computed from dial the elastic behaviour of the beam, which could, gauge and displacement transducer readings in turn, be influenced by the presence of a were almost identical, which proved the effec- plastic hinge at its other end. It is rea. tiveness Of the procedure and instrumen- sonable to rely on the safety factor as a defence tation set up. Figs 6-9 present the against even partial plastification of the beam moment-rotation curves for the four beam at midspan. depths used in the tests. The characteristics of 45, Although i t would, in principle, be pos- the standard details W1. W2 and W4 are i h - sible to track the behaviour of a structure con. trated separately in Figs 10-12. A combined taining ductile connections using advanced plot of moment against rotation for all twelve elastic.plastic this is of limited practi. tests is also presented a s Fig. 13.

40. Standard tensile tests were carried out subject to and repeated load. The semi. cal application because most structures are

to determine the tensile strengths of column flange, column web, end Plates and bolts. The it has shaken down, the response to this load

continuous frame must be designed so that once

results are summarized in Table 1. variation will be,elastic. The question is there-

Interpretation of test results 41. The test results are discussed under the 46, practical purposes, however, it can

headings of ductility, strength and stiffness. be assumed that the worst-case outcome is that the middle of the operating range-the bending moment envelope due to variable load-is a moment little greater than zero. The connection experiences negative moment at one extreme. and quite possibly positive moment of almost equal magnitude at the other. This is the result of a wind-moment design, which in a sense is one end of the spectrum of possibilities offered by semi-continuous design. The opposite end of that spectrum is continuous design, and,

fore what form the frame shakes down into, and the answer is load history dependent.

Fig. 5. Locations at broadly speaking, the ductility demands made which arms are on the connections are lessened as it is - connected I

448 approached.


47. One approach to quantifying the required rotation capacity is to work back- wards from the serviceability criterion which will be applied in design. This does, of course, presume that proper allowance will be made for connection rotation in the deflection calculation for the beam.

48. It can readily be calculated that a deflection of span/250 under uniformly distributed load corresponds, in a simply supported beam, to an end rotation of 0.0128 rad. This would vary only slightly under third point loading, or any other likely pattern of point loading from secondary beams. If, at the same time, the frame is unbraced, a horizontal deflection of height1400 corresponds to (no more than) 0.0025 rad. Even i f , somewhat unrealistically, the connection is expected to absorb a rotation equal to the sum of these extreme-case figures, and a ‘ comfort factor ’ (let us not call it a load factor in these circumstances) of say 1.6 is applied, the rotation capacity demanded is a t most 0.025 rad. That would be adequate for a connection to a simply supported beam, and connections which are more ‘performing’ (in terms of moment resistance) will tend to require less rotation capacity.

49. This fairly simple approach should be compared with expert opinion. The figure of 0.03 rad, proposed by Surtees and quoted fre- quently by other authors, is reasonably comparable. Lawson,I3 for the slightly different case of a composite beam, has proposed 0.025(1-0.5 MJM,) where M , is the connection design moment. This is based on a span/depth ratio of 30.

50. It seems reasonable to conclude this discussion by narrowing down the ‘grey area ’ to 0.02-0.03 rad. Any connection achieving 0.03 can confidently be regarded a s ductile. Any connection not achieving 0.02 ought not to be promoted as such.

51. Review of test results. The rotation capacities achieved by the proposed standard details are presented in descending order in Table 3. The most obvious conclusion to draw is that all the connections with beams up to and including 686UB are admirably ductile, whereas those with 762UB fail to perform acceptably. It was expected that ductility would decline more or less in proportion to depth, with no change in the deformability of the end plate and column flange. The abrupt- ness of the fall-off in ductility comes a s a sur- prise.

is reduced because the deeper section is also a thicker one, and the bolt-to-weld distances are less. If this is the case, it should be possible to regain ductility (at slight cost to the standardization) by ‘easing’ the bolt spacings (e.g. to 100 mm cross centres and 50/70 mm from top of flange instead of 90 mm and

52. One hypothesis is that the deformability

Table 3. Rotation capacities of standard connections

Test Column Detail

Test 4

254UC89 W2 Test 3 254UC73 W4 Test 6 254UC89 W1 Test 2 254UC89 W1 Test 9 254UC132 W4 Test 7 254UC89 W1 Test 1 254UC89 W4

Test 5 W4 254UC89 Test 8 W4 254UC132

Beam

457UB 406UB 457UB 406UB 686UB 457UB 457UB

762UB 762UB

Bolts Rotation at failure:

rad

M24 0,061 M24 0.050 M24

0,031 M24 0,033 M24 0.034 M24 0.036 M20 0.039

M24 0.019 M24 0.009

40/60 mm respectively). However, further testing would be necessary to confirm this, and in the interim it is prudent to restrict the standard details to beams less than 700 mm deep.

53. Without the two results for 762UB, it would be rather difficult, on the evidence of these tests, to declare confidently that rotation capacity is inversely proportional to depth. There is not much difference in the performance of the middle band of tests, and there is no obvious explanation for the strikingly superior performance of tests 4 and 1.

54. The influence of web buckling. One possible ‘ wild card ’ is column web buckling, which terminated tests 4, 5 and 6. These include the best a s well as the second worst, so its effect must be somewhat random. It has in the past been suggestedI4 that web buckling is a ductile mode of failure, but these tests prove that it must be regarded as one to avoid. Ductility means deformation without loss of strength; rotation alone is not enough.

55. Comparison of tests 6, 4 and 7 (all W4 with 457UB; only the column varies) shows the influence of web buckling. Ductility is greatest with 254UC89, whose 17.3 mm thick flange clearly contributes useful deformation. It is reduced by a third, although still very respectable, with 254UC132 (25.3 mm flange). But with 254UC73 (14.2 mm flange), the ductility is the lowest of the three, because web buckling supervenes. Test 10, with M20 bolts but otherwise identical to test 4 (i.e. with a too-thick end plate for its bolt size), can be added to the comparison; its behaviour is emphatically non- ductile.

56. Test 12, with a ‘ too-thick ’ end plate, displayed the lowest rotation capacity of all and is surely conclusive proof that web buckling is not a ductile failure mode.

57. Finally and somewhat contrarily, comparison of tests 5 and 8 (both with 762UB, identical except in column mass) might suggest that web buckling is a benign effect since the lighter column doubles the rotation capacity. The more likely explanation is that the positive effect of

449

BOSE AND HUGHES

Fig. 6. Moment-rotation curves for 406UB

Fig. 7. Moment-rotation curves for 45 7UB


450

O Y 1 I 0 0.02 0.04 0

Rotation: rad

Rotation: rad

Rotation: rad

16

0.06 0.08

2

0.03 0 )4

500

400

E 300 5

g L

0,

z 200

100

0 0.01 0.02 0.03

Rotation: rad

I

0.04 0


0- 0

v 0.01


l

l I

0.02 0.03 0 Rotation: rad

Fig, 10. Moment-rotation curves for standard detail W1

Fig. 11, Moment-rotation curves f o r standard detail W 2 -

451

BOSE AND HUGHES

l7 for detail W 4 Rotation: rad

the thinner flange outweighs the negative effect of the thinner web.

58, M24 details versus M20 details. Com- parison of tests 1 and 9, both with 406UB and identical except in bolt size and corresponding (standard) end plate thickness, suggests that the combination of M24/15 mm thick is a slightly more ‘ductile’ one than M20/12 mm thick. The known tendency for threads to strip with the BS 3692 M20 nut dimensions may be partly responsible, as may the significantly overstrength 12 mm end plate of the test 9 specimen (see Table 1).

59. Detail W2. Detail W2, with 20 mm thick end plate, seemed to justify itself by performing almost as well, in terms of rotation capacity, in test 3 as a similar detail with 15 mm end plate in test 11 which was done for

comparison. Nevertheless, the 17.3 mm column flange may have contributed significantly. (Perhaps the ‘ ideal ’ end plate thickness for the configuration of detail W2 is 18 mm, with M24 bolts. Unfortunately this is not a standard thickness in the UK.)

Strength 60. Without exception, the connections

achieved moment resistances strikingly in excess of those predicted by the design method of reference 7. The conservatism of the method is in one sense comforting, but it should be remembered that the design of a ductile connection aims to ensure that the end plate is a ‘ weak link ’ relative to other components. If the ‘ weak link ’ is twice as strong as predicted, what if (for example) the column web now governs?

Fig. 13. Moment-rotation curves for all tests Rotation: rad


61. Column web influence. The impression gained from this series of tests is that web buckling is very ready to govern, whereas web crushing was never observed. (Web tension was not observed either, but this would never be expected to govern when the compression zone is unstiffened.) According to the design method, crushing and buckling resistances tend to be fairly similar. If the webs had been restrained by beams in the other direction, as so often occurs in practice, a ‘crushing’ (or ‘crippling’?) failure might have been observed.

62. (As an aside, it may be remarked that all the webs which buckled did so in the ‘ non- sway’ mode, despite the shortness of the specimen columns and the absence of the usual ‘ constructional restraints ’. Perhaps the ‘ sway ’ mode-illustrated in Eurocode 3’s Figure J.2.4-just does not occur with H-sections.)

of the tests, the predicted web buckling resistance varies slightly with stiff bearing length and therefore with beam flange thickness. Typically, for a 14.5 mm thick flange (e.g. 457 X 191UB74) it is 625 kN. (This compares with a web crushing resistance of 620 kN.) The observed web buckling failures for this column section correspond to a compression of approximately 900 kN. The one such failure for 254UC73 (test 6) at approximately 625 kN can be compared with the predicted 465 kN (crushing: 470 kN). The difference is partly explained by the column sections, and in particular their webs, exceeding the ‘guaranteed minimum’ yield strength (see Table 1). Some further observations are as follows.

63. For the 254UC89 which features in most

( a ) The calculation method (based on BS 5950) seems to give safe predictions of web buckling resistance. Eurocode 3’s method- which takes the effective area t J (h2 + sf) instead of t (h + ss), where h is the column depth and S, is the stiff bearing length-is somewhat more conservative.

symmetrical test set up. ( b ) Web shear failure was precluded by the

64. M24 details uemus M20 details. With regard to the tests which failed in the tensile zone, those using M20 bolts will be considered first. Tests 9 and 10 both failed by thread stripping, although in the case of test 9 (standard detail Wl), this took place only after satisfactory ductility had been displayed.

65. Although strengths were well in excess of predictions, the M20 bolt assemblies do seem handicapped, in comparison with M24, by their thread stripping tendency (although M24 are not immune; threads stripped in test 1 and in one row of test 8). With 70% of the area, the smaller bolt achieves only 54% of the performance in detail W1 (comparing tests 9 and 1). In test 10, identical in every other respect to

test 4, the M20 bolts achieve 60% of the moment resistance of the larger bolts.

66. Two conclusions can be drawn.

( a ) The M24 8.8 bolt assembly specified to BS 3692 enjoys a hidden tensile strength advantage over its M20 counterpart.

( b ) It would be advantageous for the industry not to delay the move to IS0 4014, etc. (BSEN 2401416 etc.), with amended nut dimensions designed to give more, and more uniform, resistance to thread stripping.

67. Modes of failure. It is reasonable not to attach too much importance to the mode of failure-even if it involves the bolts- provided that it occurs only after the gross deformation which constitutes ductile behaviour. Something has to come apart in the end and it may well be the bolt, commonly after subjection to very visible bending strain. However, some observations are noteworthy.

68. Firstly, it was comforting that at no point did any weld fail, or show signs of failure, and this in spite of the deliberate provision of 10 mm fillet welds to the flanges throughout. Such welds are ‘ understrength ’ relative to the flanges of all the beams tested except the 406UB. Weld failure is even more brittle than bolt failure and greatly to be feared in ductile connections. The Design Guide’ is conservative in demanding that flange welds should be ‘full strength’, even if this is (less conservatively) interpreted to mean: ‘ make the throat thickness half the flange thickness on each side ’.

69. Another slightly adventurous choice was to specify punched (as opposed to drilled) bolt holes. In spite of the extremes of deformation of bolts and end plate, no i l l effects could be seen.4 There seems no reason to impose the restrictions of BS 5950 clause 5.3.7 on these connections.

70. The end plate fracture, which was observed in four tests, occurred not at the flange weld but very close to it, clearly in the heat-affected zone; only in test 3 was it the sole cause of failure. It was possible to repeat this test with the beams inverted; a similar result w a s ~ b t a i n e d . ~

71. Detail W2. Although the ductility achieved in test 3 was adequate, this rather brittle looking failure seemed to cast a shadow over detail W2. Test 11 of the second series was made identical except for a 15 mm thick end plate (as in all the other details using M24 bolts). The result was: a reduced moment resistance, albeit less than the [ 1 - (15/20)2 = 3 44% reduction that might have been anticipated; only slightly greater rotation capacity; and a failure which still involved end plate fracture! It is difficult to draw conclusions. In any event, it was decided

453

BOSE AND HUGHES

for other reasons not to promote detail W2 as one of the standard details featured in the Design Guide.

72. Detail W1. Detail W1 with M24 bolts was tested twice. In test 1, the beam was 406UB and the end plate 200 wide. In test 2, the beam was 686UB and the end plate 250 wide. Both tests demonstrated over twice the predicted moment resistance. This over-performance exceeded that of any other detail. The force in the bolts a t failure must have been in the region of 680-700 kN, plus any prying force which was present. Their actual tensile stress of little (if any) less than 1000 MPa compares with BS 5950’s 450 MPa or with the Design Guide’s 560 MPa. The bolts are of course stronger than their ‘guaranteed minimum’, as are the other components involved, but testing showed that the margin of overstrength was not extraordinary-generally in the region 10-20% (see Table 1).

73. Clearly, the design method on which the predictions are based is very conservative, perhaps excessively so. The tests on detail W4 paint a similar picture, although the overprediction is distinctly less. Before detailed comparisons for W4 are made, some explanation should be offered for this.

74. It is generally found that for detail W1 the effective length of equivalent T-stub is determined by the circular yield pattern (I,,, = 2 m ) , irrespective of end plate breadth. Although this pattern is undeniably a valid mechanism within the terms of the simple plastic theory being applied, it is

(a ) free from any prying effect and therefore

( b ) because of its uniquely symmetrical incompatible with the T-stub simplification

geometry, capable of developing substantial in-plane (membrane) forces to resist deformation.

75. For whatever reason, circular yielding seems not to happen; none of the test specimens displayed anything which could be said to resemble this pattern. There is a case for ignor- ing it completely, at least with plate and flange thicknesses exceeding half the bolt size a s in these standard details.

using the extended a-chart of the proposed revision’ to Annex J of Eurocode 3 and allow- ing l,,, to exceed 2 rrm, the predicted performance of test 2 improves by some 8%. Test 1, with its narrower end plate, is relatively little changed.

76. If the moment resistance is recalculated

77. Detail W4. For detail W4, three tests were performed with 457UB and varying column mass. Test 7, with 254UC132, achieved only marginally greater moment resistance than test 4 with 254UC89. This was in spite of the latter’s web buckling under a compression

in the region of 900 kN. Test 6, with 254UC73 (14.2 mm thick flange), achieved approximately 68% of the resistance of the other two. Web buckling was responsible.

78. Since the 15 mm thick end plate has, in test 7 at least, outplayed the bolts, it is tempt- ing to ask what, if any, advantage a 20 mm or 25 mm thick end plate would possess in terms of strength. (Its stiffness would, of course, be superior.) Unfortunately, test 12 does not answer this question, since web buckling governs with the result that its three bolt rows and 20 mm thick end plate transmit only 31 % more moment than test 2 with a single bolt row.

79. Summary. Broadly speaking, it is in the cases where web buckling governs that the calculation method appears to give reasonable predictions of moment resistance. Where the tension zone governs, it is systematically over- conservative with the effect that

(a ) actual moment resistances for the standard details can be expected to exceed those tabulated by 60-120%

buckling are liable to be governing even when the calculation reckons they are not.

80. Although the tests give confidence that reasonable ductility remains in these cases where web buckling is governing, but only just, it would be desirable from several points of view for the overprediction of moment resistance to be reduced to a margin which more appropriately reflects the variabilities and uncertainties involved

81. Apart from the circular yield pattern already mentioned, some features of the design method which are candidates for attention include the following.

(a) The bolt design tensile stress of 560 MPa

( b ) other component resistances such as web

has been conservatively assessed by a committee conscious of the fact that it represents a substantial increase on BS 5950’s 450 MPa. It is based on a ‘guaranteed minimum ’ ultimate tensile strength of 785 MPa, while 8.8 bolts to the lastest standards are specified as 828 MPa. Moreover, the 10% allowed for thread stripping should, in principle, be omittable once new standard dimensions become current.

( b ) The end plate and column flange also have access to reserves of strength not acknowledged by the design method. Strain- hardening and in-plane forces (which develop a s the plate bending idealization is left behind) may both be considerable.

(c) The bolt head and nut dimensions may have a non-negligible effect on the yield pattern. (The current proposed revisiona of Annex J contains a not altogether satisfactory way to take account of this; as i t


affects only the Mode 1 expression, it is unlikely to alter the calculation for these standard details.)

Rigidity

rotational stiffness. The present standard details make no claim to be Rigid ‘ with a capital R I , i.e. to qualify for use in continuous hyperstatic frames analysed elastically.

82. Rigidity is used here as shorthand for

83. Comparison with ‘Rigid ’ connections. Nevertheless, it is of interest to compare the test results with code criteria. These vary between and within codes. British Standard 5950 calls for rotational stiffness equal to that of the beam. Depending on conditions at the far end of the beam, that might mean 2 B / L , 3 B / L , 4 B / L , 6 B / L . . . ( B is the bending stiffness, and L is the length of the connected beam); the most generous interpretation (i.e. the easiest to satisfy) is 2 B/L. Eurocode 3, although not guilty of vagueness, is considerably more demanding. For unbraced frames, 25 BjL is the target; for braced ones, a more relaxed view is taken but 8 B / L remains greater than anything BS 5950 can be construed to require.

when the beam is 10 m long but fail to do so if the beam is 8 m long. In this sense, ‘ there is no such thing as a Rigid connection ’. To make sensible comparisons, we have to presume a likely span for the beam size in question, say 25 times its depth. Thus a 457 deep beam will be taken as spanning 11 m. Also, 457 deep beams come in a range of masses per unit length, within which bending stiffness may vary by a factor of 2 (whereas the connection stiffness would alter only marginally, a s a consequence of flange and web thickness change). The ‘ target ’ stiffnesses listed in Table 4 for the four beam sizes used in the tests are therefore not at all precise, but they give some indication of what the code writers may be looking for.

85. The three lines which correspond to these stiffnesses have been plotted on to the test characteristics (Figures 6-9). It can be seen that all the connections tested would qualify as Rigid, if the ‘generous interpretation ’ of BS 5950 is correct. Surprisingly perhaps, most would also satisfy or come close to satisfying the Eurocode 3 ‘ braced ’ requirement. None, not even those of tests 10 and 12 with their ‘too thick ’ end plates, would satisfy Eurocode 3’s ‘ unbraced ’ requirement. The two tests with 762 deep beams came the closest. It must be borne in mind, moreover, that in an unbraced frame the web panel shear would contribute flexibility additional to that observed in the tests.

86. Frames which feature ductile connections will be designed plastically, for which purpose the Rigid/semi-rigid distinction is strictly irrelevant. Nevertheless, it is

84. A connection could well qualify as Rigid

Table 4. Target rotational stiffnesses f o r Rigid joints

Beam designation

406 UB60 457 UB74 686 UB125 762 UB147

4ssumed length

L: m

10 11 17 19

Bending stiffness

B: MNm2

45 69

244 349

Rotational stiffness required to earn the ’ Rigid ’ appellation

BS 5950 (Figure 6.9.8) ‘ Generous Eurocode 3

interpretation Unbraced Braced of clause 6.1.4’

.

2BiL: MNm/rad MNm/rad MNmirad 25 BIL: 8 B / L :

9

459 147 37 359 115 29 157 50 13 112 36

gratifying, from the point of view of the stabil- ity and serviceability of these frames, that the standard details do not lag too far behind in this respect.

87. Prediction of rotational stiffness. It would be of value to designers (for serviceability calculations) to be able to quantify the rotational stiffness of the standard details by calculation or in semi-empirical fashion. It is the initial, or unloading/reloading, stiffness which is of interest.

88. The calculation method of Eurocode 3 Annex J does not give accurate predictions, but its draft revision’ promises to be more satisfactory. It is based on summing component flexibilities (such a s flange, web in tension/compression/shcar, end plate and bolt), each of which is evaluated from a given formula. A comparison with the present series of tests has been made. The results of this study are reported separately;” they are, as can be seen in Table 5 , still extremely variable. While there seems to be a systematic overprediction of connection stiffness (by 25- 180%), care is needed in drawing conclusions from a set of tests concentrating on a particular family of connections with thin end plates. They are not representative of the connection population at large, although it can be said that they typify the connections for which accurate

Table 5. Rotational stiffnesses of connections with 457 UB

Test Detail T Test 3 Test 4 Test 6 Test 7 Test 10 Test 11

W2 W4 W4 (light column) W4 (heavy column) as test 4 but M20 bolts (same plate) as test 3 but thinner plate

* Reference 8.

Rotational stiffness: MNm/rad

Actual

60 100 65 75 60 35

Predicted*

113 128 105 177 122 99

455

BOSE AND HUGHES

stiffness prediction is most keenly desired. It remains to be seen whether or not there is scope to improve the new formulae by calibration of the component stiffness coefficients.

89. Stiffness comparisons f o r 457 UB. Interesting comparisons can be made between the various connections to 457UB (see Table 5 and Fig. 7). The tabulated stiffnesses are approximate; they are judged to represent a ‘best fit’ straight line through the initial portion of the characteristic for the test concerned. If the two sides vary, an average is taken.4,5

test 4 should be so much stiffer than test 7 (identical but for a heavier column), or test 10 (identical but for smaller bolts). Perhaps test 4’s plot is deceptive and the stiffness has been overestimated: 70 MNm/rad would be consis- tent with the others and would not be implausi- ble. If we make this adjustment, the variations become relatively small and in line with expec- tations.

90. Results seem erratic. It is not clear why

91. M24 details versus M20 details. Between tests 1 and 9, both with 406UB (see Fig. 6), the bolt size changes and with it the end plate. The preferred M24 detail is almost twice a s stiff as the M20 equivalent.

92. Summary. The biggest single influence on rigidity is the depth-the lever arm-of the connection. Other things remaining equal, the rotational stiffness would be expected to be proportional to the square of the distance between the centres of tension and compression. The results fit this pattern, although there seems to be a striking gain in performance of the 762UB specimens by comparison with the 686UB ones. This, of course, is the obverse of the ductility degradation phenomenon. It is similarly difficult to explain. Something dispro- portionate seems to be going on and it may possibly be a consequence of the larger cruciform specimens used for tests 5 and 8 (because of the limited capacity of the test rig). Their effect is to reduce the ratio of shear force to bending moment at the beam end. This was not expected to alter the M - 4 characteristic of the connection significantly, and it would be another ‘ wild card ’ if it were found to do so.

Conclusions

have been shown to provide a well-balanced performance with beam depths up to 700 mm. Satisfactory ductility is achieved without undue sacrifice of strength or stiffness.

94. Above a beam depth of 700 mm, ductility deteriorates and the standard detail may require modification. Further work would be needed to confirm whether and how this should be done.

93. The SCI’s standard ductile connections

95. The strength calculation method of the Design Guide’ has been shown to be conservative-perhaps excessively so-when applied to connections with the geometry of these standard details.

Acknowledgements 96. Development of the standard ductile

connections has been a team effort. Many indi- viduals inside and outside the Steel Construc- tion Institute have contributed ideas, expertise and insights. Special mention is due to Dr R. M. Lawson, Dr G. W. Owens, Mr G. K. Raven and Mr J. C. Taylor at the Institute, to Dr D. Ander- son and Mr N. D. Brown at Warwick University (the home of the Wind-Moment method), to Professor D. A. Nethercot at Nottingham Uni- versity, to Dr R. Cunningham of Cunningham Associates, and to the members of the SCI/BCSA Connections Group, especially Task Group leaders Mr P. Gannon and Dr D. B. Moore. Encouragement and support from Mr E. V. Girardier at the Steel Construction Industry Federation are gratefully acknowledged.

97. The ongoing investigation into end plate connections at the University of Abertay Dundee is supported by grants from the Institu- tion of Civil Engineers’ Research and Develop- ment Enabling Fund and the Department of Environment. Financial support for the test programme was provided by the British Steel Market Development Fund, and assistance with the test specimens by the Steel Construction Industry Federation. The assistance given by Mr Z. M. Wang and Mr. G. K. Youngson at the University of Abertay Dundee during the tests and in the comparative analysis is acknowledged.

References 1. BRITISH STANDARDS INSTITUTION. Eurocode 3:

Design of steel structures, Part 1.1, General rules and rules f o r buildings. BSI, London, 1993, DD ENV 1993-1-1 : 1992.

2. BRITISH STANDARDS INSTITUTION. Structural use of steelwork in building, Part l , Code of practice f o r design in simple and continuous construction: hot rolled sections. BSI, London, 1985 5950: Part 1.

3. ANDERSON D., COLSON A. and JASPART J. P. Connec- tions and frame design for economy. New Steel Construction, 1993, Oct., 30-33.

4. BOSE B. Tests to verify the performance of standard ductile connections. Dundee Institute of Technology, 1993, Dec., Consultancy Repcrt C1/93.

5. BOSE B. Additional tests of standardised ductile connections. University of Abertay Dundee, 1994, Dec., Consultancy Report C1/94.

6. FEWSTER S. M. C., GIRARDIER E. V. and OWENS G. W. Economic design and the importance of standardised connections. Proc. First World Conf. on Constructional Steel Design, Acapulco, 1992.

7. SCI/BCSA Connections Group. Moment connections (Design guide in the series: ‘Joints in steel construction ’). The Steel Construction Institute, Ascot, 1995.

456


8. EUROPEAN COMMITTEE FOR STANDARDIZATION. Euro- code 3: Part 1.1. Revised Annex J: Joints in build- ingframes. European Committee for Standardization (CEN), Brussels, 1994, ENV 1993-1-1 : 1992/prA2.

9. ANDERSON D., READING S. J. and KAVIANPOUR K. Wind-moment design for unbracedframes. The Steel Construction Institute, Ascot, 1991.

connections-a review of test data and their application to the evaluation ofjoint behaviour on the performance ofsteelframes. CIRIA Project Record 338, London, 1985.

11. ZOETEMEIJER P. A design method f o r the tension side of statically loaded, bolted beam-to-column connections. Heron, Delft, 1974, 20, No. 1.

12. SURTEES J. 0. and MANN A. P. End plate connections in plastically designed structures. Conf, Joints in Structures, Institution of Structural Engineers, University of Sheffield, 1970, July, 1.

13. LAWSON R. M. Design of composite connections in

10. NETHERCOT D. A. Steel beam to column

composite frames-interim guidance for end plate type connections. The Steel Construction Institute, Ascot, 1994, Report No. SCI-RT-330.

14. EUROCOIJE 3 EDITORIAL GROUP. Eurocode No. 3 : Design of steel structures, Part 1, General rules and rules f o r buildings. Background documenta- tion, Chapter 6, Document 6.09, Beam to column connections. Prepared for the Commission of the European Communities. 1989, Mar.

15. BRITISH STANDARDS INSTITUTION. Specification fo r I S 0 metric precision hexagon bolts, screws and nuts. BSI, London, 1967, BS 3692.

16. BRITISH STANDARDS INSTITUTION. Hexagon head bolts, screws and nuts-product grades A , B and C. BSI, London, 1992, BS EN 24014,24015, 24016, 24017, 24018,24032,24033,24034.

appraisal of the design rules in Eurocode 3 for bolted end plate joints by comparison with experimental results. (To be published in Proc. Instn Civ. Engrs Structs C? Bldgs.)

17. BOSE B., YOUNGSON G. K. and WANG Z. M. An

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bose and hughes 1995

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semirigid design

semicontinuous approach

semirigid construction

advantages of semicontinuous

semirigid approach

benefits of semi

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