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Research into composite tubular construction for offshore jacket structures
Research into Composite Tubular Construction for Offshore Jacket Structures Dr Colin J. Billington HEAD OF STRUCTURES DIVISION, WlMPEY LABORATORIES LIMITED
SYNOPSlS This paper describes recent developments in the use of steel and concrete composite construction for offshore structures. The developments concern filling steel tubular members with cementitious materials to improve static atut fatigue strength of mentbers and joints. Applications are' described jbr jacket to pile contlections, ttodaljoittts attd repair situations. INTRODUCTION Most ott~horc structures for the production of hydrocarbons are of the traditional welded steel tubular type with piled foundations. Other structural forms (concrete gravity, steel gravity, tethered semi-submersible) have had limited use, but none has achieved the universal technical and economic acceptability of the piled steel structure. As requirements develop for production from deeper waters, various more radical approaches such as tethered buoyant platforms (TBPs) or other compliant systems have been suggested but, so far at least, the economics and outstanding technical problems have largely dictated the continuing use of piled steel structures (e.g. Shell's Cognac platform which stands in 312 metres water depth). It has been predicted (I) that TBPs could provide a cheaper alternative at depths greater than 200 metres, although above these depths payload requirements (particularly the limitation on the number of risers put at approximately 15) mean that in many cases the fixed steel platform will prove to be the most economic. One of the principal conclusions from the Platforms Working Party at the recent Symposium on EEC-funded projects in Luxembourg ~2) was that while the development of new structural forms must continue t o be encouraged, their widespread practical application is unlikely to be imminent and that therefore techniques for economising on traditional structural forms should be pursued to encourage the more rapid development of hydrocarbons, particularly for the exploitation of marginal finds. In other words, there is a requirement to
fill the gap between present technology and future unproven techniques. Furthermore, the need to extend the limits of application of fixed structures was also emphasised. Once the decision is taken to develop a field, the profitability of the operation will depend very much on the development programme being kept to schedule. A disappointing feature of novel concepts is that they invariably take longer to complete than originally envisaged and consequently the oil companies" reluctance to adopt new methods is understandable. This then is further encouragement for the re-examination of existing methods. The principal design and construction problems associated with tubular steel structures concern the nodal joints, the design of which is controlled by static punching shear strength and fatigue considerations. These design problems are exacerbated in deeper water where both static and environmental toadings are greater and massive welding, post weld heat treatment and inspection problems result. Furthermore, the installation of steel platforms in deeper water is complicated by transportation limitations on the size and weight of structures which therefore may have to be installed in sections with the problems of underwater mating and connection. However, a feasibility study carried out by Wimpey has shown that the use of steel and concrete composite construction could alleviate the joint design problem. Composite construction in this context is the filling of annuli between concentric steel cylinders or the filling of complete steel members with a cementitious material (concrete, mortar or grout) (see Figure 1). In this way the materials are used efficiently in applications most suited to their mechanical properties. The eementitious material is contained and therefore greater strength and ductility is achieved. The steel tubular is the containment medium and therefore is predominantly subjected to hoop tension, and the cementitious filling minimises any tendency for buckling of the steel shell and reduces the need for internal or external stiffening.
Research into composite tubular construction for offshore jacket structures The technique is of course already widely used for the attachment of tubular steel piles to the main legs of the jacket and the grouting process could easily be extended to other structural members previously left unfilled. (Research and design guidance on grouted jacket to pile connections has until recently lagged behind their use offshore. The results of a recent major research programme are discussed in a later section of this paper.)
19
(iv) Resistance to external hydrostatic pressure would be greatly enhanced and the need for internal or external stiffening would be considerably reduced. (v) In the event of a ship collision or other impact (e.g. earthquake) the energy absorption characteristics of a composite structure, and hence its resistance to collapse, would be enormously increased over a traditional steel structure. The remainder of this paper describes research already carried out at Wimpey Laboratories into composite components for steel jacket structures. This has concentrated on two areas - annular connections (for connection of jacket to piles or members end-to-end) and nodal connections.
PILE TO JACKET CONNECTIONS
SLEEVE
POSSIBLE FILLING OF BRACE
GROUT ANNULUS
PILE
~ ~ ' ~
Figure 1
"~'~] ",
~
POSSIBLEFILLING OF TUBULAR PILE
Possible uses of composite construction for tubular joints
The use of composite construction could lead to improvements in the following areas: (i) Punching shear and bending capacities of nodal joints would be greatly enhanced. (ii) Stress concentration factors at nodal joints would be considerably reduced offering much improved fatigue life. (iii) Axial load capacity of composite members (e.g. legs or braces) would be much greater than that of the steel section alone (the compression strength of concrete filled tubular columns can be greater than the sum of the individual squash strengths of the concrete and steel components).
The steel tubular piles are driven (or placed into predrilled holes) through the legs of the structure or through tubular sleeves attached to the lower part of the structure. The annulus between pile and jacket leg or sleeve is then filled with cement grout. In shallow water normally a single pile is placed through each leg of the structure and often extends to the top of the leg so that the deck structure is welded directly to the pile. The jacket resists the effects of wave loading, stabilises the piles and stiffens the complete structure and the grout transmits lateral forces between the piles and jacket but is not normally required to transmit vertical loads from the deck superstructure to the piles. Bond stresses at the steel/grout interfaces are low and debonding is unlikely to prejudice the overall structural stability. In deeper water, however, the piles do not extend to the surface and are either grouped in clusters around the main legs of the structure or in some cases are evenly distributed around the base of the structure (skirt piles). The grouted connection forms the only structural connection between the jacket and its foundations and the grout is required to transmit forces arising both from the dead weight of the jacket, deck and superstructure, and from environmental loading. In addition, with the development of offshore oil finds in deeper water, the capacity of offshore construction plant fot handling and driving piles has increased, resulting in a trend towards foundations consisting of smaller numbers of larger diameter less radially stiffconnections, in view of the increasing importance of grouted connections and this trend towards larger diameters, a large research effort has been directed into defining the principal parameters which affect the strength of grouted connections and establishing design procedures which describe the effects of these parameters. Testing programmes for several oil companies (J) showed that the main parameters which affect strength are: (I) Radial stiffness (i.e. diameter/thickness) of pile
(O/t)p
20
Journal of Constructional Steel Research: Vol. 1. No. 1: September 1980
(2) Radial stiffness of sleeve (o/t), (3) Radial stiffness of grout (D/f ). (4 Grout compressive strength (5) Length :diameter ratio (LID,) (6) Shear connector axial spacing ratio (D,/s) (7) Shear connector height ratio (h/D,) _ _ I- ~ (8) Surtace roughness (C,). Therefore the U.K. Department of Energy has funded a major programme of research at Wimpey Laboratories to investigate the effect of these parameters on the strength of the connection. The strength of axiallyloaded grouted connections is generally defined in terms of an equivalent bond strength which is obtained by dividing the ultimate capacity of the connection by the total surface area of the pile/grout interface.
F nv
FJL’d
bond strength fbu is given by
f
F,” =bu 1.105
50 [J”
1 O.’
(3)
where /bu and f,, are in N/mm*. Equation 3 gives the multiple of the API RP2At6’ recommended ultimate (i.e. six times design stress) bond stress at a cube stress of 50 N/mm*. The Department of Energy programme has confirmed the above relationships.
Effect of length : diameter ratio Two series of tests have been carried out to investigate the relationship between bond strength and L/D, for
/
= 21 I2
i IO
-.
06
. 04
02
0 0
Figure ?
I
2
3
4
6
5
7
6
9
IO
II
I2 ‘L/d
Eflecl of LID pnrcrmerer on bond srrenRlh
Effect of grout strength and tubular stiff nesses Reference 3 showed that for both plain pipe connections and connections with weld beads the ultimate bond strength Jbu is related to the grout compressive strength fcu (measured by compressive strength on 3 inch grout cubes) by: fb” =
(1)
f t;’
Reference-3 also showed a linear relationship -/;I" and a stiffness factor given by K = 2
[(D/r),k’ s
Equation common
+ [(D/f),
the range I.0 < L/D, < 12.0. The results are shown in Figure 2 in which the results are presented as the ratio C, between Fbu at the LID, considered and Fbu at L/D, = 2.0. This shows a slight reduction in average bond strength with increasing L/D, although the maximum reduction is 20 per cent of the value at L/D, = 2.0.
+ (D/l)&-’
between
(2)
I can be used to normalise test results to a compressive strength. A non-dimensionalised
Effect of shear connector geometry The tests have considered complete hoop orcontinuous spiral weld bead or welded bar shear connectors with the same height and spacing on both pile and sleeve surfaces. Both height and spacing have been considered separately and linear relationships have been established between bond strength and the ratios D,ls and h/D,.
Research into composite tubular construction for offshore jacket structures
Generalised formula for ultimate bond
COMPOSITE NODES
strength Using the relationships established above it is possible to arrive at a general formula for grout bond strength. The results of both plain pipe specimens and those with shear connectors are shown in Figure 3 by plotting the parameter Fbu/KC L against h/s. This allows for all the parameters listed earlier. This figure demonstrates a number of important points:
Where piles pass through the main legs of structures the annulus between the pile and leg member is usually filled with a cement grout. In some cases the centre of the pile is also filled. In either case a composite section results which improves the strength of nodal joints where bracing members meet the main leg. In the past, this increase in strength has generally not been allowed for in design, principally because of the lack of data on the strength of composite nodes, but also due to doubts about the efficacy of the grouted annulus. However, using newly developed radioactive density measuring techniques, m the completeness of the grouting can be
(i) There is a linear relationship between Fb,/KC t and h/s. (ii) There is a greater scatter in proportion to the mean for plain pipe connections (h/s = 0).
F~
21
600-
KCL
500 -
400
F,., KC,
-
300 -
J
1
76 Cs ~ 9 4 6 8 h/S
f
200 i
lO0 -
I
00
Figure 3
0.01
0.02
0.03
h/S
0.04
Relationship between Fb./KC L and his
(iii) The contributions to bond strength due to plain pipe bond and shear connectors are additive. From the above a generalised grout bond formula which is a mean to the test results is given by: Fbu = KCL(76C , + 9468h/s)
(4)
However this formula does not reflect the large scatter in results particularly for plain pipe connections and therefore a characteristic ultimate bond strength (i.e. the strength above which 95 per cent of test results fall) should be used for design. Some 400 tests have been carried out at Wimpey Laboratories and a study of these shows that the characteristic value is 75 per cent of the mean. A working party has been set up by the Department of Energy to review the test data and it is hoped that this will result in a new comprehensive design approach.
established and a major research project, partly funded by the EEC, in progress at Wimpey Laboratories, is providing the design data. To illustrate the potential strength improvement, the results from a testing programme {~1carried out for BP in connection with an appraisal of three- and four-legged jacket structures in the Abu Dhabi Marine Areas (ADMA) in the Arabian Gulf are described below. The testing programme was devised to (a) check the ultimate strength of ungrouted T-shaped joints representing typical ADMA geometry and compare this to the API design clauses ~6~and to (b) measure the increase in strength due to the presence of the grouted pile. A total of ten specimens were tested with three different loading conditions. Five ungrouted specimens were tested, one subject to axial tension, two to axial compression and two to in-plane bending loads to
22 Journal of Constructional Steel Research: Vol. 1, No. 1: September 1980 ten specimens and where applicable first crack loads are also given. The ultimate strength implied by API RP2A as defined in the following sections is also given in Table 1 for each specimen. The ratios of recorded ultimate, first yield and first crack loads to the API RP2A ultimate strength are also given.
obtain data for direct comparison with the API design clauses. Five grouted specimens were tested with the same loading conditions to determine the enhancement in strength due to the grouted pile. The dimensions of the tubulars from which the specimens were fabricated are consistent with a I : 1.6 scale representation of typical brace to main leg (chord) joints. Outline details of the specimen geometry are given in Figure 4.
(a) Axial tension For the ungrouted specimen (AI 1) the recorded first yield and ultimate loads were 8.0 ton and 47.2 ton respectively. The mechanism of failure was formation of a crack through the chord wall. This crack propagated, from the toe of the joint and was first
Recorded ultimate loads and failure modes A summary of the test results is given in Table 1. The first yield and ultimate loads are given for each of the
144in
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•
III
:~No 90TO~ ~ I RAMS
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~ ,~.SPHERICAL i
6625m BRACE
OD
SEATINGS
8Oin PIN CENTRES
CHORD OOIn OO
30,n PILE 12 5in 0 O GROUTED SIMECIMENS
•
-i- TON., SECTION 'A- ~,'
ELEVATION
ELEVATION
( a ) COMPRESSION
Figure 4
Table !
(b) I N - P L A N E
BENDING
Detailsofjoint specimens
Tubular T-joints. Recorded first yield, first crack and ultimate loads Recorded loads
Specimen reference AI1 AI2" AI3 A 14" A 15
Loading
Tension
Compression
A 16 ° AI7 A 18" A 19 A 20 °
First yield Py (ton)i"
First crack Pc (ton)
Ultimate P~ (ton)
API design ultimate load P~ (ton)
P~
P,
Pc
8.0 23.0
37.0 68.0
47.2 81.8
23.0 23.0
2.05 3.56
0.16 0.28
0.78 0.83
26.0 • 166.9 24,1 170.5
23.0 23.0 23.0 23.0
1.13 7.26 1.05 7.41
0.35 0,72 0.33 0.59
4.42 8.18 4.30 7.27
0.38 0.33 0.39 0.33
9.0 120.0 8.0 100.0
In-plane bending 20 in lever arm
" Grouted specimens f Imperial ton (2240 Ib)
2.8 4.4 2.8 4.0
6.9 I 1.0 6.7 12.0
7.3 13.5 7.1 12.0
1.65 1.65 1.65 1.65
0.95 0.82 0.94 0.67
Research into composite tubular construction for offshore jacket structures observed at a load of 37.0 ton. Considerable distortion of the chord cross-section was observed. The corresponding grouted specimen ( A I 2 ) g a v e first yield and ultimate loads of 23.0 ton and 8.18 ton. The mechanism of failure was similar to A 11 with a crack forming at the toe of joint, this being first observed at a load of 68.0 ton. No distortion of the chord cross-section was noted except in the immediate vicinity of the joint. Thus, the first yield load, first crack load and the ultimate strength were enhanced by 185 per cent, 84 per cent and 73 per cent respectively due to the presence of the grouted pile.
23
specimens loaded in tension, considerable distortion of the chord cross-section occurred. For the two grouted specimens ( A I 4 and A I 6 ) a punching shear failure did not occur because the grouted pile prevented distortion of the chord wall. The maximum loads reached for the two specimens were 166.9 ton and 170.5 ton at which loads the braces collapsed in compression. Thus, an enhancement in ultimate load by a factor of at least 6.7 was recorded. At these ultimate loads the mean brace stress was approaching the ultimate strength of the material. A large enhancement of the load at which yield
J
J
120
12 in
IOO
21n
;L o 80
KEY
60
UNGROUTED TENSION == GROUTED
TENSION
UNGROUTED COMPRESSION o
40
GROUTED
201
lf.a
o¢ 0
I
I 0 05
I
I 0 iO
I
RELATIVE
FigureJ
I
0 a5
O 20 DISPLACEMENT
......[
0 25
COMPRESSION •
J
0 30
(IN)
Relationship between load and relatice displacement between brace and chord
Graphs of load against relative displacement between brace and chord are given in Figure 5 and show clearly the enhancement in ultimate and first yield loads and also in the stiffness of the specimen due to grouting.
occurred in the chord wall was also recorded with an increase from 9.0 ton and 8.0 ton for the two ungrouted specimens to 120.0 ton and 100.0 ton for the two grouted specimens. The load-displacement graphs for the compression specimens are given in Figure 2.
(b) Axial compression Two ungrouted specimens (AI3 and A I5) were tested in compression and failed at loads of 26.0 ton and 24. i ton. The failure mode was a classic compressive punching shear collapse of the chord wall with considerable plastic deformation of the chord wall at the chord/brace intersection. As for ungrouted
(c)
In-plane bending
The two ungrouted specimens (A17 and A I9) failed at loads of 6.9 ton and 6.7 ton applied at a 20 in lever arm to the outside of the chord wall. The first yield load was 2.8 ton for both specimens. During the ultimate load test large plastic deformations of the chord wall
24
Journal of Constructional Steel Research: Fol. 1, No. 1: September 1980
occurred. Inward displacements of the chord wall were recorded on the compression side of the joint and outward displacements on the tension side. At the maximum load a permanent relative rotation of approximately 5 ° between the longitudinal axes of the chord and brace had occurred. The centre of this rotation was close to the neutral axis of the brace. After these large plastic deformations, a small crack was noted in the chord wall on the tension side of the joint. The failure loads of the two grouted specimens (AI8 and A20) were 13.5 ton and 12.0 ton respectively with corresponding first yield loads of 4.4 ton and 4.0 ton. These loads represent average increases of 87 per cent and 50 per cent in ultimate and first yield loads respectively. The mechanism of failure was modified by the grouted pile. As for the axial compression loading case, inward deformation of the chord wall was prevented and the centre of rotation at failure was displaced by half the brace diameter from the position recorded for ungrouted specimens to the compression face of the brace. The mechanism of failure was therefore that of yield followed by extensive cracking of the chord wall which propagated from the tension side of the joint.
be expected, but also that for the same applied cyclic load range a significant enhancement in the fatigue life of the joint may be expected. The fatigue life of a joint is controlled by the maximum "hot spot' stress range, and the design fatigue life for a given 'hot spot" stress range can be determined from a S - N curve such as the American Welding Society (7) X-curve which can be expressed as logl0 N = 10.2 - 4.38 logt0 S
(5)
when N is the fatigue life in cycles and S is the applied stress range in ton/in z. For this particular curve, S must be the.'hot spot" stress as determined either by
150~ 12
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50 E tNSO IN
IZ5- - ~ O i 0~
\ COM5p5I(1ON \
\
,:;f
o
o
(oI *AAX~MUUPRN I CP I AL3'TE t;I SSES
Stress distribution around the joint From the strain results and the known material properties, the elastic distributions of principal stresses on the outer surface of the chord and brace were determined at gauges positioned 0.375 in from the weld. The distributions of maximum and minimum principal stresses and maximum shear stress are given in Figure 5 for axial tension and compression. The stresses are expressed for axial specimens as a multiple of the mean brace stress J;, and for in-plane bending specimens as a multiple of the maximum positive brace stress f~ as calculated from simple beam theory. The outer surface principal stress distributions given in Figure 6 as multiples of the brace stresses f , (axial) show that in all cases the magnitudes of the stresses are reduced by the presence of a grouted pile. The most significant reductions are in the chord wall for axial loading where the maximum recorded principal stresses are reduced from 13. If, and 13.5f, for tension and compression respectively to 5.3f= and 3.0f, respectively. The reductions due to grouting are 60 per cent and 78 per cent respectively. The corresponding reductions in the maximum brace stresses are 50 per cent and 58 per cent respectively. The reductions again demonstrate that there are differences in behaviour between grouted specimens subjected to axial tension and compression loading. There is good agreement between the magnitude of stresses in ungrouted tension and compression specimens with maximum stresses of 13. If= and 13.5f, respectively, i.e. a difference of only 3 per cent. The reduced magnitude of stresses indicates not only that the increase in first yield and ultimate loads should
:7 '2"~ ~o ~
~r
COUem(ssJow ~ ~ ~'-...
..
-2~, . NK(~ (} i CO
eAAC( . . . . .
1
i i ~---\x\r(NSION
~c I u~l~u~
5HEAR STRESSES
Figure 6 Stress distribution around thejoint
experimental means or by detailed analysis (e.g. finite element methods). Although the actual "stress concentration factors' cannot be accurately determined from the test results, a reduction of the same order as that of the maximum recorded principal stress might be expected, i.e. 60 per cent for tension loading and 78 per cent for compression loading. These reductions would result in a 55-fold and 760-fold increase in design fatigue lives respectively, based on the AWS X-curve. Alternatively, significantly larger load ranges may be allowed for the same design fatigue life.
C O M P O S I T E REPAIR M E T H O D S The above results indicate significant improvements in strength due to composite action at nodes and also
Research into composite tubular construction for offshore jacket structures
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¢. O N
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PULt.F- 0 0 v £ ~ . NL~," Z o~.NT
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Details of typical annular connection i¢a,'2 ~
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r Figure8 Possible repair method.for a T node provide design information for grouted annular connections. There is an increasing requirement for strengthening or repairing structures where both annular connections and composite action at nodes can provide the most economic and technically satisfactory solution. Annular grouted connections were used in the West Sole platform strengthening. Such connections must be designed for combinations of bending and axial loads and therefore a further testing programme at Wimpey Laboratories has been investigating the strength of annular connections subject to combined loading. The type of connection is shown in Figure 7 where a sleeve is placed over the tubulars to be connected (or over damaged section) and the annulus filled with grout. A repair method for a T node is shown in Figure 8. This consists of two or more sections bolted together around the damaged node and again the annuli are grout filled.
The testing programme has so far concentrated on annular connections which have been tested in axial tension and compression, pure bending and combined axial and bending loads. Tests have been carried out with and without shear connectors on the brace surface. The pure bending tests showed that there was no bond failure within the yield capacity of the brace. The axial load tests showed that tensile brace loads produced lower bond failure loads than compressive loads. Tensile ultimate bond strengths were in reasonable agreement with those predicted from the generalised grout bond formula derived from tests on jacket:pile connections. The results of combined axial load and bending tests are shown in Figure 9 where non-dimensionalised ultimate bond strength is plotted against the ratio of extreme brace fibre bending stress (fb) to mean axial stress (f,). This shows an 18 per cent reduction in bond strength from f b / f , = 0 to
26 Journal of Constructional Steel Research: 1Iol. 1, No. 1: September 1980
1.6
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1.2
~0
O-g
f,/f,
O.tl,
Figure 9
o
o~
i'"5
I'-O
t'-o
Interaction curve for combined bending and tension specimens
f h / f . = 1.0 and an increasing strength as the bending component increases beyond this. The next phase of the work will consider nodal repairs.
2
CONCLUDING
3
REMARKS
This paper has demonstrated the benefits obtained when composite members-and connections are used in steel tubular offshore structures. Composite construction can be used in design or for strengthening and repairing existing structures. A new formula derived from a large number of test results is proposed to describe the strength of annular connections subjected to axial loading. The work described in this paper forms the background to a feasibility study carried out by Wimpey which has shown that significant savings can be made in the cost of piled steel tubular offshore jacket structures. As a result a major research programme is now in progress to produce comprehensive design guidance for composite offshore structures.
4 5
6
7
European Conlmunitics. Symposium on New Technologies for Exploration and Exploitation of Oil and Gas Resources. Luxembonrg, April 1979. COMMISSION OF TIlE EUROPEAN COMMUNITIES. SympQsium of New Technologies for Exploration and Exploitation of Oil and Gas Resources. Lu~embourg, April 1979. BnA, INGTON, C. J. and LEWIS, G. It. G. 'The strength of large diameter grouted connections." Paper O T C 3083 of Offshore Technology Conference. Texas, May 1978. "Radiation cuts risks in offshore monitoring.' New Civil Engineer, 4 October 1979. TEBBE'I'I', I. E., BECKETT, C. D. a n d BILLINGTON, C. J. 'The punching shear strength of tubular joints reinforced with a grouted pile.' Paper O T C 3463 of Offshore Technology Conference, Texas. April 1979. AMERICAN PETROLEUM INSTITUTE. 'Recommended practice for planning, designing and constructing fixed offshore platforms." API RP2A (7th Edition), Texas, January 1976. AMERICAN WELDING SOCIETY. Structural Welding Code, AWS DI 1-75, 1975.
In this paper test results have been presented in Imperial units. The following conversions may be helpful: 1 tonf = 9.964 kN l inch -----25.4 mm
REFERENCES l
LANG, J. R. A. and HEDLEY, C. J. 'BP development of tethered buoyant platform production system.' Commission of the
Contributions discussing this paper should be received by the Editor before I January 1981.
SYNOPSlS This paper describes recent developments in the use of steel and concrete composite construction for offshore structures. The developments concern filling steel tubular members with cementitious materials to improve static atut fatigue strength of mentbers and joints. Applications are' described jbr jacket to pile contlections, ttodaljoittts attd repair situations. INTRODUCTION Most ott~horc structures for the production of hydrocarbons are of the traditional welded steel tubular type with piled foundations. Other structural forms (concrete gravity, steel gravity, tethered semi-submersible) have had limited use, but none has achieved the universal technical and economic acceptability of the piled steel structure. As requirements develop for production from deeper waters, various more radical approaches such as tethered buoyant platforms (TBPs) or other compliant systems have been suggested but, so far at least, the economics and outstanding technical problems have largely dictated the continuing use of piled steel structures (e.g. Shell's Cognac platform which stands in 312 metres water depth). It has been predicted (I) that TBPs could provide a cheaper alternative at depths greater than 200 metres, although above these depths payload requirements (particularly the limitation on the number of risers put at approximately 15) mean that in many cases the fixed steel platform will prove to be the most economic. One of the principal conclusions from the Platforms Working Party at the recent Symposium on EEC-funded projects in Luxembourg ~2) was that while the development of new structural forms must continue t o be encouraged, their widespread practical application is unlikely to be imminent and that therefore techniques for economising on traditional structural forms should be pursued to encourage the more rapid development of hydrocarbons, particularly for the exploitation of marginal finds. In other words, there is a requirement to
fill the gap between present technology and future unproven techniques. Furthermore, the need to extend the limits of application of fixed structures was also emphasised. Once the decision is taken to develop a field, the profitability of the operation will depend very much on the development programme being kept to schedule. A disappointing feature of novel concepts is that they invariably take longer to complete than originally envisaged and consequently the oil companies" reluctance to adopt new methods is understandable. This then is further encouragement for the re-examination of existing methods. The principal design and construction problems associated with tubular steel structures concern the nodal joints, the design of which is controlled by static punching shear strength and fatigue considerations. These design problems are exacerbated in deeper water where both static and environmental toadings are greater and massive welding, post weld heat treatment and inspection problems result. Furthermore, the installation of steel platforms in deeper water is complicated by transportation limitations on the size and weight of structures which therefore may have to be installed in sections with the problems of underwater mating and connection. However, a feasibility study carried out by Wimpey has shown that the use of steel and concrete composite construction could alleviate the joint design problem. Composite construction in this context is the filling of annuli between concentric steel cylinders or the filling of complete steel members with a cementitious material (concrete, mortar or grout) (see Figure 1). In this way the materials are used efficiently in applications most suited to their mechanical properties. The eementitious material is contained and therefore greater strength and ductility is achieved. The steel tubular is the containment medium and therefore is predominantly subjected to hoop tension, and the cementitious filling minimises any tendency for buckling of the steel shell and reduces the need for internal or external stiffening.
Research into composite tubular construction for offshore jacket structures The technique is of course already widely used for the attachment of tubular steel piles to the main legs of the jacket and the grouting process could easily be extended to other structural members previously left unfilled. (Research and design guidance on grouted jacket to pile connections has until recently lagged behind their use offshore. The results of a recent major research programme are discussed in a later section of this paper.)
19
(iv) Resistance to external hydrostatic pressure would be greatly enhanced and the need for internal or external stiffening would be considerably reduced. (v) In the event of a ship collision or other impact (e.g. earthquake) the energy absorption characteristics of a composite structure, and hence its resistance to collapse, would be enormously increased over a traditional steel structure. The remainder of this paper describes research already carried out at Wimpey Laboratories into composite components for steel jacket structures. This has concentrated on two areas - annular connections (for connection of jacket to piles or members end-to-end) and nodal connections.
PILE TO JACKET CONNECTIONS
SLEEVE
POSSIBLE FILLING OF BRACE
GROUT ANNULUS
PILE
~ ~ ' ~
Figure 1
"~'~] ",
~
POSSIBLEFILLING OF TUBULAR PILE
Possible uses of composite construction for tubular joints
The use of composite construction could lead to improvements in the following areas: (i) Punching shear and bending capacities of nodal joints would be greatly enhanced. (ii) Stress concentration factors at nodal joints would be considerably reduced offering much improved fatigue life. (iii) Axial load capacity of composite members (e.g. legs or braces) would be much greater than that of the steel section alone (the compression strength of concrete filled tubular columns can be greater than the sum of the individual squash strengths of the concrete and steel components).
The steel tubular piles are driven (or placed into predrilled holes) through the legs of the structure or through tubular sleeves attached to the lower part of the structure. The annulus between pile and jacket leg or sleeve is then filled with cement grout. In shallow water normally a single pile is placed through each leg of the structure and often extends to the top of the leg so that the deck structure is welded directly to the pile. The jacket resists the effects of wave loading, stabilises the piles and stiffens the complete structure and the grout transmits lateral forces between the piles and jacket but is not normally required to transmit vertical loads from the deck superstructure to the piles. Bond stresses at the steel/grout interfaces are low and debonding is unlikely to prejudice the overall structural stability. In deeper water, however, the piles do not extend to the surface and are either grouped in clusters around the main legs of the structure or in some cases are evenly distributed around the base of the structure (skirt piles). The grouted connection forms the only structural connection between the jacket and its foundations and the grout is required to transmit forces arising both from the dead weight of the jacket, deck and superstructure, and from environmental loading. In addition, with the development of offshore oil finds in deeper water, the capacity of offshore construction plant fot handling and driving piles has increased, resulting in a trend towards foundations consisting of smaller numbers of larger diameter less radially stiffconnections, in view of the increasing importance of grouted connections and this trend towards larger diameters, a large research effort has been directed into defining the principal parameters which affect the strength of grouted connections and establishing design procedures which describe the effects of these parameters. Testing programmes for several oil companies (J) showed that the main parameters which affect strength are: (I) Radial stiffness (i.e. diameter/thickness) of pile
(O/t)p
20
Journal of Constructional Steel Research: Vol. 1. No. 1: September 1980
(2) Radial stiffness of sleeve (o/t), (3) Radial stiffness of grout (D/f ). (4 Grout compressive strength (5) Length :diameter ratio (LID,) (6) Shear connector axial spacing ratio (D,/s) (7) Shear connector height ratio (h/D,) _ _ I- ~ (8) Surtace roughness (C,). Therefore the U.K. Department of Energy has funded a major programme of research at Wimpey Laboratories to investigate the effect of these parameters on the strength of the connection. The strength of axiallyloaded grouted connections is generally defined in terms of an equivalent bond strength which is obtained by dividing the ultimate capacity of the connection by the total surface area of the pile/grout interface.
F nv
FJL’d
bond strength fbu is given by
f
F,” =bu 1.105
50 [J”
1 O.’
(3)
where /bu and f,, are in N/mm*. Equation 3 gives the multiple of the API RP2At6’ recommended ultimate (i.e. six times design stress) bond stress at a cube stress of 50 N/mm*. The Department of Energy programme has confirmed the above relationships.
Effect of length : diameter ratio Two series of tests have been carried out to investigate the relationship between bond strength and L/D, for
/
= 21 I2
i IO
-.
06
. 04
02
0 0
Figure ?
I
2
3
4
6
5
7
6
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Eflecl of LID pnrcrmerer on bond srrenRlh
Effect of grout strength and tubular stiff nesses Reference 3 showed that for both plain pipe connections and connections with weld beads the ultimate bond strength Jbu is related to the grout compressive strength fcu (measured by compressive strength on 3 inch grout cubes) by: fb” =
(1)
f t;’
Reference-3 also showed a linear relationship -/;I" and a stiffness factor given by K = 2
[(D/r),k’ s
Equation common
+ [(D/f),
the range I.0 < L/D, < 12.0. The results are shown in Figure 2 in which the results are presented as the ratio C, between Fbu at the LID, considered and Fbu at L/D, = 2.0. This shows a slight reduction in average bond strength with increasing L/D, although the maximum reduction is 20 per cent of the value at L/D, = 2.0.
+ (D/l)&-’
between
(2)
I can be used to normalise test results to a compressive strength. A non-dimensionalised
Effect of shear connector geometry The tests have considered complete hoop orcontinuous spiral weld bead or welded bar shear connectors with the same height and spacing on both pile and sleeve surfaces. Both height and spacing have been considered separately and linear relationships have been established between bond strength and the ratios D,ls and h/D,.
Research into composite tubular construction for offshore jacket structures
Generalised formula for ultimate bond
COMPOSITE NODES
strength Using the relationships established above it is possible to arrive at a general formula for grout bond strength. The results of both plain pipe specimens and those with shear connectors are shown in Figure 3 by plotting the parameter Fbu/KC L against h/s. This allows for all the parameters listed earlier. This figure demonstrates a number of important points:
Where piles pass through the main legs of structures the annulus between the pile and leg member is usually filled with a cement grout. In some cases the centre of the pile is also filled. In either case a composite section results which improves the strength of nodal joints where bracing members meet the main leg. In the past, this increase in strength has generally not been allowed for in design, principally because of the lack of data on the strength of composite nodes, but also due to doubts about the efficacy of the grouted annulus. However, using newly developed radioactive density measuring techniques, m the completeness of the grouting can be
(i) There is a linear relationship between Fb,/KC t and h/s. (ii) There is a greater scatter in proportion to the mean for plain pipe connections (h/s = 0).
F~
21
600-
KCL
500 -
400
F,., KC,
-
300 -
J
1
76 Cs ~ 9 4 6 8 h/S
f
200 i
lO0 -
I
00
Figure 3
0.01
0.02
0.03
h/S
0.04
Relationship between Fb./KC L and his
(iii) The contributions to bond strength due to plain pipe bond and shear connectors are additive. From the above a generalised grout bond formula which is a mean to the test results is given by: Fbu = KCL(76C , + 9468h/s)
(4)
However this formula does not reflect the large scatter in results particularly for plain pipe connections and therefore a characteristic ultimate bond strength (i.e. the strength above which 95 per cent of test results fall) should be used for design. Some 400 tests have been carried out at Wimpey Laboratories and a study of these shows that the characteristic value is 75 per cent of the mean. A working party has been set up by the Department of Energy to review the test data and it is hoped that this will result in a new comprehensive design approach.
established and a major research project, partly funded by the EEC, in progress at Wimpey Laboratories, is providing the design data. To illustrate the potential strength improvement, the results from a testing programme {~1carried out for BP in connection with an appraisal of three- and four-legged jacket structures in the Abu Dhabi Marine Areas (ADMA) in the Arabian Gulf are described below. The testing programme was devised to (a) check the ultimate strength of ungrouted T-shaped joints representing typical ADMA geometry and compare this to the API design clauses ~6~and to (b) measure the increase in strength due to the presence of the grouted pile. A total of ten specimens were tested with three different loading conditions. Five ungrouted specimens were tested, one subject to axial tension, two to axial compression and two to in-plane bending loads to
22 Journal of Constructional Steel Research: Vol. 1, No. 1: September 1980 ten specimens and where applicable first crack loads are also given. The ultimate strength implied by API RP2A as defined in the following sections is also given in Table 1 for each specimen. The ratios of recorded ultimate, first yield and first crack loads to the API RP2A ultimate strength are also given.
obtain data for direct comparison with the API design clauses. Five grouted specimens were tested with the same loading conditions to determine the enhancement in strength due to the grouted pile. The dimensions of the tubulars from which the specimens were fabricated are consistent with a I : 1.6 scale representation of typical brace to main leg (chord) joints. Outline details of the specimen geometry are given in Figure 4.
(a) Axial tension For the ungrouted specimen (AI 1) the recorded first yield and ultimate loads were 8.0 ton and 47.2 ton respectively. The mechanism of failure was formation of a crack through the chord wall. This crack propagated, from the toe of the joint and was first
Recorded ultimate loads and failure modes A summary of the test results is given in Table 1. The first yield and ultimate loads are given for each of the
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ELEVATION
ELEVATION
( a ) COMPRESSION
Figure 4
Table !
(b) I N - P L A N E
BENDING
Detailsofjoint specimens
Tubular T-joints. Recorded first yield, first crack and ultimate loads Recorded loads
Specimen reference AI1 AI2" AI3 A 14" A 15
Loading
Tension
Compression
A 16 ° AI7 A 18" A 19 A 20 °
First yield Py (ton)i"
First crack Pc (ton)
Ultimate P~ (ton)
API design ultimate load P~ (ton)
P~
P,
Pc
8.0 23.0
37.0 68.0
47.2 81.8
23.0 23.0
2.05 3.56
0.16 0.28
0.78 0.83
26.0 • 166.9 24,1 170.5
23.0 23.0 23.0 23.0
1.13 7.26 1.05 7.41
0.35 0,72 0.33 0.59
4.42 8.18 4.30 7.27
0.38 0.33 0.39 0.33
9.0 120.0 8.0 100.0
In-plane bending 20 in lever arm
" Grouted specimens f Imperial ton (2240 Ib)
2.8 4.4 2.8 4.0
6.9 I 1.0 6.7 12.0
7.3 13.5 7.1 12.0
1.65 1.65 1.65 1.65
0.95 0.82 0.94 0.67
Research into composite tubular construction for offshore jacket structures observed at a load of 37.0 ton. Considerable distortion of the chord cross-section was observed. The corresponding grouted specimen ( A I 2 ) g a v e first yield and ultimate loads of 23.0 ton and 8.18 ton. The mechanism of failure was similar to A 11 with a crack forming at the toe of joint, this being first observed at a load of 68.0 ton. No distortion of the chord cross-section was noted except in the immediate vicinity of the joint. Thus, the first yield load, first crack load and the ultimate strength were enhanced by 185 per cent, 84 per cent and 73 per cent respectively due to the presence of the grouted pile.
23
specimens loaded in tension, considerable distortion of the chord cross-section occurred. For the two grouted specimens ( A I 4 and A I 6 ) a punching shear failure did not occur because the grouted pile prevented distortion of the chord wall. The maximum loads reached for the two specimens were 166.9 ton and 170.5 ton at which loads the braces collapsed in compression. Thus, an enhancement in ultimate load by a factor of at least 6.7 was recorded. At these ultimate loads the mean brace stress was approaching the ultimate strength of the material. A large enhancement of the load at which yield
J
J
120
12 in
IOO
21n
;L o 80
KEY
60
UNGROUTED TENSION == GROUTED
TENSION
UNGROUTED COMPRESSION o
40
GROUTED
201
lf.a
o¢ 0
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I
I 0 iO
I
RELATIVE
FigureJ
I
0 a5
O 20 DISPLACEMENT
......[
0 25
COMPRESSION •
J
0 30
(IN)
Relationship between load and relatice displacement between brace and chord
Graphs of load against relative displacement between brace and chord are given in Figure 5 and show clearly the enhancement in ultimate and first yield loads and also in the stiffness of the specimen due to grouting.
occurred in the chord wall was also recorded with an increase from 9.0 ton and 8.0 ton for the two ungrouted specimens to 120.0 ton and 100.0 ton for the two grouted specimens. The load-displacement graphs for the compression specimens are given in Figure 2.
(b) Axial compression Two ungrouted specimens (AI3 and A I5) were tested in compression and failed at loads of 26.0 ton and 24. i ton. The failure mode was a classic compressive punching shear collapse of the chord wall with considerable plastic deformation of the chord wall at the chord/brace intersection. As for ungrouted
(c)
In-plane bending
The two ungrouted specimens (A17 and A I9) failed at loads of 6.9 ton and 6.7 ton applied at a 20 in lever arm to the outside of the chord wall. The first yield load was 2.8 ton for both specimens. During the ultimate load test large plastic deformations of the chord wall
24
Journal of Constructional Steel Research: Fol. 1, No. 1: September 1980
occurred. Inward displacements of the chord wall were recorded on the compression side of the joint and outward displacements on the tension side. At the maximum load a permanent relative rotation of approximately 5 ° between the longitudinal axes of the chord and brace had occurred. The centre of this rotation was close to the neutral axis of the brace. After these large plastic deformations, a small crack was noted in the chord wall on the tension side of the joint. The failure loads of the two grouted specimens (AI8 and A20) were 13.5 ton and 12.0 ton respectively with corresponding first yield loads of 4.4 ton and 4.0 ton. These loads represent average increases of 87 per cent and 50 per cent in ultimate and first yield loads respectively. The mechanism of failure was modified by the grouted pile. As for the axial compression loading case, inward deformation of the chord wall was prevented and the centre of rotation at failure was displaced by half the brace diameter from the position recorded for ungrouted specimens to the compression face of the brace. The mechanism of failure was therefore that of yield followed by extensive cracking of the chord wall which propagated from the tension side of the joint.
be expected, but also that for the same applied cyclic load range a significant enhancement in the fatigue life of the joint may be expected. The fatigue life of a joint is controlled by the maximum "hot spot' stress range, and the design fatigue life for a given 'hot spot" stress range can be determined from a S - N curve such as the American Welding Society (7) X-curve which can be expressed as logl0 N = 10.2 - 4.38 logt0 S
(5)
when N is the fatigue life in cycles and S is the applied stress range in ton/in z. For this particular curve, S must be the.'hot spot" stress as determined either by
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Stress distribution around the joint From the strain results and the known material properties, the elastic distributions of principal stresses on the outer surface of the chord and brace were determined at gauges positioned 0.375 in from the weld. The distributions of maximum and minimum principal stresses and maximum shear stress are given in Figure 5 for axial tension and compression. The stresses are expressed for axial specimens as a multiple of the mean brace stress J;, and for in-plane bending specimens as a multiple of the maximum positive brace stress f~ as calculated from simple beam theory. The outer surface principal stress distributions given in Figure 6 as multiples of the brace stresses f , (axial) show that in all cases the magnitudes of the stresses are reduced by the presence of a grouted pile. The most significant reductions are in the chord wall for axial loading where the maximum recorded principal stresses are reduced from 13. If, and 13.5f, for tension and compression respectively to 5.3f= and 3.0f, respectively. The reductions due to grouting are 60 per cent and 78 per cent respectively. The corresponding reductions in the maximum brace stresses are 50 per cent and 58 per cent respectively. The reductions again demonstrate that there are differences in behaviour between grouted specimens subjected to axial tension and compression loading. There is good agreement between the magnitude of stresses in ungrouted tension and compression specimens with maximum stresses of 13. If= and 13.5f, respectively, i.e. a difference of only 3 per cent. The reduced magnitude of stresses indicates not only that the increase in first yield and ultimate loads should
:7 '2"~ ~o ~
~r
COUem(ssJow ~ ~ ~'-...
..
-2~, . NK(~ (} i CO
eAAC( . . . . .
1
i i ~---\x\r(NSION
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Figure 6 Stress distribution around thejoint
experimental means or by detailed analysis (e.g. finite element methods). Although the actual "stress concentration factors' cannot be accurately determined from the test results, a reduction of the same order as that of the maximum recorded principal stress might be expected, i.e. 60 per cent for tension loading and 78 per cent for compression loading. These reductions would result in a 55-fold and 760-fold increase in design fatigue lives respectively, based on the AWS X-curve. Alternatively, significantly larger load ranges may be allowed for the same design fatigue life.
C O M P O S I T E REPAIR M E T H O D S The above results indicate significant improvements in strength due to composite action at nodes and also
Research into composite tubular construction for offshore jacket structures
~i, t-lt.,,.li,¢.
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PULt.F- 0 0 v £ ~ . NL~," Z o~.NT
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Details of typical annular connection i¢a,'2 ~
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r Figure8 Possible repair method.for a T node provide design information for grouted annular connections. There is an increasing requirement for strengthening or repairing structures where both annular connections and composite action at nodes can provide the most economic and technically satisfactory solution. Annular grouted connections were used in the West Sole platform strengthening. Such connections must be designed for combinations of bending and axial loads and therefore a further testing programme at Wimpey Laboratories has been investigating the strength of annular connections subject to combined loading. The type of connection is shown in Figure 7 where a sleeve is placed over the tubulars to be connected (or over damaged section) and the annulus filled with grout. A repair method for a T node is shown in Figure 8. This consists of two or more sections bolted together around the damaged node and again the annuli are grout filled.
The testing programme has so far concentrated on annular connections which have been tested in axial tension and compression, pure bending and combined axial and bending loads. Tests have been carried out with and without shear connectors on the brace surface. The pure bending tests showed that there was no bond failure within the yield capacity of the brace. The axial load tests showed that tensile brace loads produced lower bond failure loads than compressive loads. Tensile ultimate bond strengths were in reasonable agreement with those predicted from the generalised grout bond formula derived from tests on jacket:pile connections. The results of combined axial load and bending tests are shown in Figure 9 where non-dimensionalised ultimate bond strength is plotted against the ratio of extreme brace fibre bending stress (fb) to mean axial stress (f,). This shows an 18 per cent reduction in bond strength from f b / f , = 0 to
26 Journal of Constructional Steel Research: 1Iol. 1, No. 1: September 1980
1.6
I-4
1.2
~0
O-g
f,/f,
O.tl,
Figure 9
o
o~
i'"5
I'-O
t'-o
Interaction curve for combined bending and tension specimens
f h / f . = 1.0 and an increasing strength as the bending component increases beyond this. The next phase of the work will consider nodal repairs.
2
CONCLUDING
3
REMARKS
This paper has demonstrated the benefits obtained when composite members-and connections are used in steel tubular offshore structures. Composite construction can be used in design or for strengthening and repairing existing structures. A new formula derived from a large number of test results is proposed to describe the strength of annular connections subjected to axial loading. The work described in this paper forms the background to a feasibility study carried out by Wimpey which has shown that significant savings can be made in the cost of piled steel tubular offshore jacket structures. As a result a major research programme is now in progress to produce comprehensive design guidance for composite offshore structures.
4 5
6
7
European Conlmunitics. Symposium on New Technologies for Exploration and Exploitation of Oil and Gas Resources. Luxembonrg, April 1979. COMMISSION OF TIlE EUROPEAN COMMUNITIES. SympQsium of New Technologies for Exploration and Exploitation of Oil and Gas Resources. Lu~embourg, April 1979. BnA, INGTON, C. J. and LEWIS, G. It. G. 'The strength of large diameter grouted connections." Paper O T C 3083 of Offshore Technology Conference. Texas, May 1978. "Radiation cuts risks in offshore monitoring.' New Civil Engineer, 4 October 1979. TEBBE'I'I', I. E., BECKETT, C. D. a n d BILLINGTON, C. J. 'The punching shear strength of tubular joints reinforced with a grouted pile.' Paper O T C 3463 of Offshore Technology Conference, Texas. April 1979. AMERICAN PETROLEUM INSTITUTE. 'Recommended practice for planning, designing and constructing fixed offshore platforms." API RP2A (7th Edition), Texas, January 1976. AMERICAN WELDING SOCIETY. Structural Welding Code, AWS DI 1-75, 1975.
In this paper test results have been presented in Imperial units. The following conversions may be helpful: 1 tonf = 9.964 kN l inch -----25.4 mm
REFERENCES l
LANG, J. R. A. and HEDLEY, C. J. 'BP development of tethered buoyant platform production system.' Commission of the
Contributions discussing this paper should be received by the Editor before I January 1981.