A11102 L 5 5 5 3 4

NATL INST OF STANDARDS & TECH R.I.C.

All 102655534

Yokel, Felix Y/Mat foundations for offsh QC100 .U56 NO.86-341 9 1987 V19 C.1 NBS-P

) IMUOdS 86-3419

Mat Foundations for Offshore

Structures in Arctic Regions

Felix Y. Yokel

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

Center for Building Technology

Gaithersburg, MD 20899

Robert G. Bea

PMB Systems Engineering, Inc. 500 Sansome Street, Suite 400

San Francisco, CA 94111

May 1986

Issued February 1987

Prepared for: U.S. Department of the Interior

Minerals Management Service

Technology Assessment and Research Branch

647 National Center

Reston, Va 22091

1986

100

.1156

-3419

NBSIR 86-3419

MAT FOUNDATIONS FOR OFFSHORE

STRUCTURES IN ARCTIC REGIONS

Felix Y. Yokel

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

Center for Building Technology

Gaithersburg, MD 20899

NBS

RESEARCH

INFORMATION CENTER

mi

4 ^

Robert G. Bea

PMB Systems Engineering, Inc. 500 Sansome Street, Suite 400

San Francisco, CA 941 1

May 1986

Issued February 1987

Prepared for: U.S. Department of the Interior

Minerals Management Service

Technology Assessment and Research Branch

647 National Center

Reston, Va 22091

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS. Ernest Ambler. Director

ABSTRACT

Design limit states for offshore gravity structures in the

Alaskan Beaufort and Eastern Chukchi continental shelves are

discussed. The report contains a description of geological

conditions, design loads, and type of structures used. Three

foundation types are considered: foundations for artificial

islands; foundations for caisson retained island with sand cores;

and rigid foundations for various types of gravity structures

which are positioned on the ocean floor with a minimum of

preparation. Design limit states for these foundations are

identified and the required reliability against the occurence of

these .limit states is discussed. Our ability to determine foundation resistance is assessed.

Key Words: artificial islands; geotechnical engineering; ice

forces; mat foundations; ocean engineering; offshore platforms;

oil production; sand and gravel berms; soil ex p 1 or at 1 on ; soil

test i ng

.

i i i

TABLE OF CONTENTS

ABSTRACT i i i LIST OF FIGURES v

1.

INTRODUCTION 1

2.

SUBSURFACE CONDITIONS

2. 1 General ..................................

2.2 Geological Character i st i cs of Unconsol i dated Deposits

2.2. 1 Surficial Geology

2.2.2 Evidence of Instability

2.2.3 Seismicity and Faults

2.2.4 Phenomena Associated with the Arctic Region . 2.3 Engineering Character i st i cs of Unconsolidated Deposits

2.3.1 Soil Types Encountered

2.3.2 Strength Char acter i st i cs of Soils ......... 2.3.3 Summary of Engineering Character i sti cs ....

3.

DESIGN LOADS

3. 1 General

3.2

Loads Acting on Structures

4.

DESIGN LIMIT STATES

4. 1 General

4.2 Requirements for Safety and Servi ceabi 1 i ty ........

4. 3 Structural Concepts . 4.3.1 Background Information

4.3.2 Gravity Structures Used or Considered ...... 4.4 Design Limit States for Foundations

4.4.1 Definition of Limit States ................. 4.4.2 Practical Limits to Load and Resistance

4.4.3 Design Strategies

4.4.4 Failure Probabi 1 i t i es Assoc, with Limit States

4.4.5 Design Limit States for Artificial Islands

4.4.6 Design Limit States for Caisson Retained Isl.

4.4.7 Design Limit States for Rigid Mat Foundations.

4.4.8 Design Limit States for Sand and Gravel Berms.

4.5 Resistance to Loads and Di spl acements

4.5.1 General

4.5.2 Character i zati on of Soils

4.5.3 Augmentation of Soil Resistance 1 4.6 Reliability of Predictions 1

5. SUMMARY 133

6. NEEDED STUDIES AND RESEARCH 137

7. REFERENCES 139

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LIST OF FIGURES

Figure 1

.

Fi gure 2.

F i gure 3.

Fi gure 4.

F i gure 5

.

F i gure 6

.

F i gure 7.

Figure 8.

F i gure 9.

Figure 10.

Figure 1 1

.

Figure 12.

Figure 13.

Figure 14.

Figure 15.

F i gure 16.

Fi gure 17.

Figure 18.

Seismic Re-flection Profiles Taken by U.S.G.S. in

The Alaskan Beaufort and Eastern Chukchi Seas ... Thickness of Holocene Marine Sediments on the

Alaskan Beaufort and Chukchi Shelves

Youthful Landslide Terranes

Data on Scour and Erosion

Recorded Earthquake Epicenters . Areas of Shallow Gas Deposits and Hydrated Gas

Schematic Sketch of Ice Regime in Early Spring

Ice Zonation in the Alaskan Beaufort and Eastern

Chukchi Seas

Idealized Sketch of Ice Gouge

Ice Gouge Intensities in the Alaskan Beaufort and

Eastern Chukchi Seas

Distribution of Bottom Sediments on the Beaufort

Sea Shelf

Sand and Gravel Resources on the Beaufort Sea

Shelf

Soil Profiles in the Prudhoe Bay Area

Idealized Schematic Profile of Unconsolidated

Deposits in the Beaufort Sea Shelf

Normalized Shear Strength and Excess Pore Water

Pressures vs. Shear Strain Obtained in Undrained

Consolidated Triaxial Tests of Beaufort Sea Silts

Undrained Shear Strength Profile for Beaufort Sea

Silts

Over consol i dat i on Ratio Profile for Beaufort Sea

Silts

Water Content vs. Undrained Shear Strength for

Beaufort Sea Silts
Figure 19. Pro-file of Expected Range of (Jndrained Shear

Strengths for Normally Consolidated Alaska OCS

Silts 31

Figure 20. Soil Profiles in Harrison Bay 32

Figure 21. Provisional Map of Subsea Permafrost Distribution 34

Figure 22. Current Patterns Recorded in the Beaufort and

Eastern Chukchi Seas 39

Figure 23. Typical Strength Character i st i cs of Sea Ice .... 43

Figure 24. Correlations Between Ice Pressures and Contact

Area 44

Figure 25. Global Ice Forces as a Function of Mean Recurrence

Interval 46

Figure 26. Estimated Global Ice Forces Against a 120-180m

Wide Exploration Structure as a Function of

Water Depth 47

Figure 27. Estimated Global Ice Forces Against a 120- 180m

Wide Production Structure as a Function of

Water Depth 48

Figure 28. Winter Season Ice Pressure Record for the

Canadian Beaufort Sea 49

Figure 29. Calculated Force-Time History for Large Ice

Feature Impact 50

Figure 30. Typical Winter Ice Loading Histogram 51

Figure 31. Risk for Selected Engineering Projects 56

Figure 32. Probability of Failure Obtained from Cost Benefit

Analysis 57

Figure 33. Typical Shallow Dredged Island Cross Sections .. 62

Figure 34. Tar suit Island 63

Figure 35. The Gulf Mobil Arctic Caisson (MAC) 64

Figure 36. Exxon's Super CIDS 65

Figure 37. The BWACS Floating Caisson 67

Figure 38. The Arctic Cone Exploration Structure (ACES) ... 68

vi i i
Figure 39. Adjustable Monocone Rig 69

Figure 40. The Sohio Arctic Mobile Structure (SAMS) ........ 70

Figure 41. Load Resistance o-f Mat Foundation .............. 73

Figure 42. Instrumentation Used to Monitor the Performance o-f SAMS 76

Figure 43. Settlement Control 1 Planned -for SAMS ........... Figure 44. Potential Failure Modes o-f Fill Islands 83

Figure 45. Failure Modes Identified -for Tarsuit Island .... 86

Figure 46. Failure Modes -for Caisson Retained Islands 88

Figure 47. Failure Modes -for Gravity Structures 92

Figure 48. Contact Pressures at the Ridges o-f Ice Gouges .. 93

Figure 49. Contact Pressures on Mat Foundations in the

North Sea 94

Figure 50. Data on CPT-SPT Correlations 102

Figure 51. Data on Correlations Between SPT Blow Count and

Li que-f acti on Potential of Sands and Silty Sands . 103-

Figure 52. Predicted and Back Calculated Skirt Resistance . 104

Figure 53. Deformation Modes Associated With Shear Failures . 108

Figure 54. Effect of Deformation Modes on Shear Strength .. 109

Figure 55. Effect of Consolidation Under the Weight of the

Structure on Shear Strength Ill

Figure 56. Effect of Apparent Over consol i dat i on on Shear

Strength Profiles 112

Figure 57. Effect of Cyclic Load on Monotonic Shear Strength

in Undrained Conditions 114

Figure 58. Undrained Cyclic Shear Strength of North Sea

Clays 115

Figure 59. Shake Table Data on Cyclic Mobility 117

Figure 60. Stress Controlled Cyclic Triaxial Test of

Reconstituted Sand Sample 119
Figure 61. Observed Threshold Strain 120

Figure 62. Correlations Between Cyclic Strain and Porewater

Pressure Bui d up 121

Figure 63. Correlation Between Shearwave Velocity and

Liquefaction Potential of Sands 122

Figure 64. Steady State Strength 124

Figure 65. Definition of the State Parameter 127

Figure 66. Correlation Between the State Parameter and the

Angle of Shearing Resistance 128
1. INTRODUCTION

Offshore structures for oil and gas exploration in Arctic

regions, and particularly in the Beaufort Sea, must cope with

unique environmental conditions. The configuration, design and

construction of these structures and their foundations are to a

large degree dictated by these conditions. The most critical

environmental effects are; the short period of time available for

construction (about 30 days of open water conditions); the

logistic difficulties associated with the transport of structural

materials and components to the site; the scarcity of nearby sand

and gravel deposits at most sites; the large lateral ice forces

acting on the structures; and the unique soil conditions prevailing at most sites presently under consi derat i on

The transfer of lateral loads from offshore structures to the

supporting soil on the sea floor is a design constraint which has

a major effect on the cost and configuration of these structures.

Foundation types presently under consi derat i on are large mat

foundations. The structural systems comprising these foundations

can either be constructed on artificial sand or gravel berms or

placed directly on the seafloor with a minimum of preparation.

The load capacity of such mat foundations can be augmented by

adding skirts or large diameter piles (spuds), by various methods

of soil improvement which increase the shear strength of the

supporting soil, and by underfilling methods which improve the

contact between the mats and the supporting soil. Alternately,

lateral loads or required safety margins can be reduced by

various design and operations strategies.

The purpose of this report is to define critical design limit

states for mat foundations in the arctic, discuss design strategies used to prevent occurrence of these limit states, and assess

our capability to determine the reliability against foundation

failures associated with these limit states.

1
Chapter 2 of the report deals

discusses design loads; design

discussed in chapter 4; and

suggests needed research.

with soil conditions; chapter 3

limit states and strategies are

chapter 5 contains a summary and
2. SUBURFACE CONDITIONS

2. 1 General

The currently available data base -for continental shelf areas in

the U.S. Beau-fort and Eastern Chukchi seas is derived -from multichannel seismic re-flection profiles (see figure 1) and other

geotechnical exploration data taken by U.S.G.S. and interpreted

in Reference C 1 D , and from proprietary soil exploration studies

carried out in conjunction with specific projects. Some of the

available information does not lend itself to accurate interpretation and is likely to be re-interpreted as more geotechnical

information becomes available.

The National Petroleum Council estimated C21 that " proven

technology and sufficient expertise for advanced design work is

available for the industry to proceed confidently with operations

in water as deep as 650 feet (200m) in the Southern Bering Sea

and to about 200 feet (60m) in the more severely ice covered

areas of the Northern Bering, Chukchi, and Beaufort Seas". Consequently, this report is primarily concerned with conditions to a

water depth of about 60m (the 60m isobath). In the U.S. Beaufort

Sea this covers most of the area landward of the continental

shelf break.

Since this report deals with the stability and load-deflection character i sties of structural foundations, the discussion of

suburf ace char ac t er i st i cs is confined to the unconsolidated

deposits which affect structural performance. It is, however

recognized, that structural performance may also be affected by

drilling problems encountered in much deeper seated formations

(reservoir subsidence). Such problems are not within the scope of

this report.
Fi gure 1 „ Seismic

A1 ask an

Reflection Profiles Taken by U.S.G.S. in

Beaufort and Northeastern Chukchi Seas [1J.

4

the
2.2 Geological Character i st i cs of Unconsol i dated Deposits

2.2.1 Surficial Geology

The continental shel-f of the U.S. Beaufort Sea is underlain by

the Quaternary deposits of the "Gubik Formation", which consists

of shallow water marine and terrestrial sediments, containing

sands, gravels, silts and clays. These deposits vary in thickness

from 10 to 200m. The Gubik formation is generally covered by much

looser Holocene deposits which are 5 to 45m thick and typically

tend to form a wedge which thickens in the offshore direction

and reaches its maximum thickness near the shelf break. This

wedge of Holocene deposits tends to be thicker in the Eastern

half of the Beaufort shelf. The Holocene deposits in the shallow

parts of the Eastern Chukchi Sea appear to be in most places less

than 5m thick but to increase locally to about 12m.

Figure 2 shows contours of the thickness of the most poorly

consolidated marine sediments. Thewe contours were based on USGS

high resolution profiles, with some of the information corroborated by drilling and diving samples. Along the coast, the

Holocene layer appears to be thin or absent because of erosion

effects. Also, there is relatively little sediment accumulation

in deltas and offshore from rivers.

2.2.2 Evidence of Instability

In the middle and outer shelf there is some evidence of development of very low angle bedding plane slides. This may be taken as

an indication that the shear strength of the deposits, or of some

discrete strata within these deposits is probably low. High

resolution seismic profiles spaced 15 to 50 km apart along the

entire Western part of the shelf show evidence of instability

terranes seaward of the 50 to 65m isobath along the shelf

break. These include tabular sheets up to 38 km long and typi5
Figure 2. Thickness of Holocene Marine Sediments on the Alasl an Beaufort and Chukchi shelves [ID
cally 20 to 230 m thick which move seaward along slip planes

which dip only 0.5 tO 1.5 degrees, and thus must include layers

of very low shear strength. Figure 3 shows a preliminary mapping

of these youthful landslide terranes.

Along the Beaufort coast, coastal erosion and scour are widespread and result in the seaward migration of barrier islands.

Large scour craters (strudel scours), 15 to 25m in diameter, are

formed near the mouth of rivers during the spring flooding of

fast ice areas. Erosion and scour data are shown in figure 4.

2.2.3 Seismicity and faults

Most of the area under discussion has been historically aseismic,

however a single earthquake has been located 200 miles north of

the Coleville River delta, and there is a zone of concentrated

seismic activity in the vicinity of Barter Island. Figure 5 shows

a U.S.G.S mapping of epicenters for earthquakes of Magnitude 3 or

greater. The largest of these recorded earthquakes had a Magnitude of 5.3.

Young faults are abundant in the Camden Bay area. They are

thought Cl] to be associated with an area of Holocene uplift.

Available data suggest that at least some of these faults are

active. Some additional faults which displace Pleistocene

deposits were also recorded. It is suggested Cl] that the

midshelf faults, which would be of concern in conjunction with

offshore structures, are either quiescent or have a very long

recurrence interval.

2.2.4 Phenomena Associated With the Arctic Region

( 1 ) Relict Permafrost

Quaternary deposits below the Holocene deposits were frozen

during the last glacial sea level lowstand to a depth

estimated to exceed 300m Cl, 4, 5], Much of this permafrost

7
F i gure 3. Youth -f ul Landslide T err anes [13.

8
Figure 4. Data on Scour and Erosion L 1

3
* Recorded earthquake epicenters

Figure U m Recorded Earthquake Epicenters C1D.
melted after exposure to saline water above the freezing

point- However, a significant portion of the deposits is

still fully or partially ice bonded- The distribution of this

"relict" permafrost was not extensively explored. However, it

is reasonable to assume that there is no relict permafrost in

the offshore Holocene deposits which are geologically younger

than the Pleistocene permafrost. Areas on the inner shelf

that were character i zed C43 exhibit ice bonded permafrost at

variable depths which may be several hundred meters, overlain

by partially bonded permafrost- Nearshore, permafrost is also

likely to be present in areas of rapid coastal

erosion- Subsea permafrost recorded in the Chukchi Sea is

discussed in Reference [53.

The presence of relict permafrost beneath other soil deposits

poses important constraints to the design of structural

foundations. These include settlements and loss of bearing

strength associated with the thawing of these

deposits. Thawing can be caused by changes in the thermal

regime of the subsoil induced by the presence of the offshore

structure and by production or exploration wells. Very

large settlements can result from the effect of oil

production or exploration, even if the relict permafrost

is located at a considerable depth.

( 2 ) Shallow Gas and Gas Hydrates.

Shallow gas deposits, either of thermogenic origin (originating from natural gas deposits at greater depth, or of

biogenic origin (decomposition product from buried organic

material) have accumulated in many areas beneath the shelf

of the Beaufort and Chukchi Seas C13 (see figure 6). These

gas deposits pose structural hazards by inhibiting the normal

consolidation of soils and thereby causing pockets of

abnormally low shear strength. They also can cause blowouts

during drilling operations.
Shallow gas

No gas

Figure 6. Areas o-f Shallow Gas Deposits and Hydt a' ed G < > 1 1

12
Gas hydrates, which are gases caged in the interstices of an

expanded ice crystal lattice, occur in the relict permafrost

deposits at greater depths, as seen in figure 6. If encountered, and there is no evidence that this would happen

landward of the 60m isobath, they could pose a hazard if

changes in the thermal regime cause the permafrost to thaw.

( 3 ) Ice Gouges

Ice gouging is discussed in References Cl, 6, and 73 which in

turn reference many studies. The process of ice gouging can

be understood in terms of the ice regime, which fluctuates

seasonally. With the progress of winter, the area inside the

barrier islands freezes and forms a zone of fast ice (very

little or no ice motion). Part of this fast ice at a shallow

water depth extends to the full water depth ("bottom fast

ice"), while the other part at greater depth is floating. At

the seaward boundary of the fast ice is the polar ice pack,

which consists of ice floes, many of them multi year ice,

which move generally in a westerly direction. The polar ice

pack intrudes on the zone of fast ice and exerts pressure on

it. As a result, a zone of first year and multi year ice

ridges (the "Stamuki " zone) forms. These ridges build up to a

great thickness and have deep keels which are frequently

grounded and generally form in water depths from 15 to

45m. Figure 7 shows, a schematic sketch of the ice regime

during early spring before the onset of thawing. Figure 8

shows the approximate limits of the various ice zones. In

spring Arctic rivers flood the fast ice canopy which

eventually breaks up and melts, while the polar icepack

retreats. Maximum open water generally occurs in September

and early October, however grounded remnants of the Stamuki

zone may persist during the open water season.

Gouging is caused by the keels of grounded ice ridges which

are dragged along the sea floor. The most intensive gouging

13
Figure 7. Schematic Sketch o-f Ice Regime in tar 1 y Spring 163.

14
Figure 8. Ice Zonation in the Alaskan Beau-fort and Eastern

Chukchi Seas C6D.

15
in the Beau-fort Sea occurs in the Stamuki zone between the 15

and 45m isobaths. In accordance with Barnes et al.C6], a

"typical" maximum gouge per 1km tracking length, embodying

the mean value of extensive survey data, would occur in 18m

deep water , form a furrow 0.56m below the mean sea floor and

a ridge 0.47m above the mean sea floor (total relief of

1.03m) , and have a width at the seafloor level of 7.8m. There

would be 70 gouges per km22 with an average tracking length of

1km in each km32

. The dominant orientation of the gouges is

E~W, as would be expected with an ice movement in a Westerly

direction. An idealized sketch of a gouge and a gouge

multiplet (typically caused by first year ice ridges) is

shown in figure 9. Figure 10 shows a mapping of observed

gouge intensities, where gouge intensity is the product of

maximum gouge depth in m, maximum gouge width in m, and gouge

density in No. of gouges per km2

, and is used as a statistical measure. Maximum gouge dimensions observed in the

survey in Ref. [611 were 67m width, 4m depth (with a single

value of 5.5m) and 5m height of flanking ridges. Maximum

density observed was 500 gouges per km2 .

The engineering implications of these gouges are very

serious, particularly when it is planned to use pre fabricated mats resting directly on the ocean f

1

oor , in which case

the gouges would result in only partial contact between the

mat and the ocean floor. The effect of backfilling these?

gouges in the case of underfill and gravel berms would also

be uneven contact stresses, since it is not feasible to

construct compacted fills.

16
Figure 9. Idealized Sketch o-f Ice Gouge.

17
3 (100-999) Medium

4 (1000) High

Figure 10. Ice Gouge Intensities in

Eastern Chukchi Seas CSI.

the Alaskan Beoufort and

18
2.3 Engineering Char acter i st i cs of Unconsol i dated Deposits

2.3.1 Soil Types Encountered

Figure 11 shows a distribution of bottom sediments in the

Beaufort Sea shelf, taken from Reference C83. Soil c 1 assi f i cat i on

is by "mean sediment size", which is interpreted as the particle

diameter below which 507. of any sample by weight is smaller

( D«s© ) . It can be seen that particle sizes ranging from gravels to

clays were identified, with silts generally predominating at

smaller depths and clays at greater depths. A partial map of sand

and gravel resources in the Beaufort Sea is shown in Figure

12. More comprehensive information on sand and gravel resources

is presented in figure 21 in Reference [13 and in Ref.[8a3.

These data show surface deposits and do not necessarily reflect

soil conditions important for foundation design, since the

Holocene deposits shown in the figures could be quite shallow.

Some specific information is shown in the soil profiles in figure

13, which were taken in a notherly direction in Prudhoe Bay. An

idealized schematic profile is shown in figure 14. In general the

data indicate that conditions tend to be quite variable and

cannot be predicted in the absence of site specific data and that

pockets of soft clayey silts and relict permafrost may be

encountered within layers of more competent material. The spacing

of exploratory borings must therefore be close enough to detect

pockets of soft soil. As previously noted and shown in figure 14,

there is a general trend for the thickness of the Holocene

deposits, which cover the more competent Pleistocene deposits, to

increase with increasing water depth.

2.3.2. Strength Character i st i cs of Soils

It is in general dangerous, and perhaps undesirable to try to

generalize available information on subsurface conditions, since

the soils in each individual location have their own unique

19
Figure 11. Distribution o-f Bottom Sediments on the Beaufoet Sea

20

Shel-f C8II.
ISO* 141 *

Gravel and sand (Undiffarantlated)

ffflll Predominantly sand

[ 1 1 Predominantly gravel

Figure 12. Sand and Gravel Resources on the Beau-fort Sea Shelf

Cl].

21
OEPTH

BELOW

8EA

LEVEL

(m) sand and * lit

Soft clayey allt

0 2 4 • 8 10 12

DISTANCE FROM 8HORE (km)

21 22 23 1 24 26 26

PB-6

Proba holas 20

PB-7

0

Qra

Bottom of

exploration

(PB) Proba borings

I I I I

Soft ciayay sUt

W/zT/// So,ld

Qravelly *y]L i bonded

permafrost

Figure 13. Soil Pro-files in the Prudhoe Bay Area till.
big ure 14. Idealised schematic profile of Unconsolidated Deposits

in the Beaufort Sea Shelf C11D.

23
characteri st i cs and geological history which need to be explored

and assessed, and the soils on the Beau-fort and Eastern Chukchi

shelves tend to be quite variable. However -for the purpose o-f this report the trend o-f available data can give some insight

into unique conditions existing in the region and problems that

may be encountered with gravity structures.

Wang et al . C93 studied a wide variety o-f silts which seem

to predominate in that region. Actually the soils they studied

include soils which by the Unified Soil Classification (ASTM

D2487) II 101 would be classified as clays because of their high

Plasticity Index, however the authors point out correctly that

the engineering properties of these soils resemble those of

silts, rather than clays. Figure 15 shows upper and lower bounds

for shear strength and excess pore water pressures as a function

of shear strains obtained in consolidated undrained triaxial

compression tests. In the figure, ° % = major principal stress,

= minor principal stress, a'c= = effective confining pressure

at onset of test (consol i dat i on pressure) and u = excess pore

water pressure. The upper bound of shear strength is for the

densest samples. It can be seen that the soil is dilative and

shear strength increases with shear strain. The lower bound is

for the loosest samples which are somewhat contractive but their

shear strength is stable and does not decrease with increasing

shear def ormati ons. Figure 16 shows the range of undrained

shear strengths with depth for Beaufort Sea silts and figure 1/

shows overconsol i dat i on ratios as a function of depth obtained

from odometer tests. In figure 16, TXUU and TXCU are undrained

unconsol i dated , and undrained consolidated triaxial tests,

respect i vel y

The data in the figures should be viwed with caution, because

they are from laboratory tests and are probably affected by

sample disturbance. Sample disturbance effects are particularl

great for heavily overconsol i dated samples. Nevertheless, r

24
D/n

?

qz/

I- i gure

o’

Lb. Normalised Shear Strength and Excess Pore WaterPressures vs. Shear Strain Obtained in Undrained Consolidated Triaxial Tests of Beaufort Sea Silts C93.
UJz

=_J Q

D

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0

10

20

$ 30

O

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40

50

60

UNDRAINED SHEAR STRENGTH (KSF)

1 2 3 4 5

1 « ©

s A

T T

A •

A

m s

• • i

A

A

• TXUU data

Interpolated from TXCU data

A Laboratory vane data

i i l

Figure 16= Undrained

Silts C9II.

Shear Strength Profile for Beaufort >ea

26
OVER CONSOLIDATION RATIO

0 2 4 6 8 10 12

Figure 17. Overconsol i dat i on Ratio Pro-file -for Beau-fort Sea Silts

C 93 .

27
trends emerge from the data; (1) the soils are heavily overconsolidated, particularly at a shallow depth, with a considerable

scatter of the overconsol i dat i on ratio at a shall ow depth. The

authors attribute this overconsolidation to the effects of

past freezing and call it "apparent over consol i dat

on " . The

origin of the overconsol i dat i on may be responsible for the

variability of the data. However some of that variability may

also be caused by sample disturbance.

(2). While there is considerable scatter in the shear strength at

a shallow depth, with some very high values, there is a trend,

for depths greater than about 10 ft, for the shear strength to

increase with depth. There is little doubt that the high strength

values at shallow depths and their scatter correspond to the high

and scattered values of over consol i dat i on ratios discussed under

( 1 ) . If this is the case, strength values should not be relied on

in design, because of their variability, and also because

overconsol i dat

on effects could be eliminated as a result of

disturbances and large deformations which can cause a reduction

in confining pressures. (3). The normalized shear strength values

shown in figure 15 are for strains of up to 7‘/., and for the A

curve strains up to 3“/.„ The tests were not carried to a large

enough strain to achieve a steady state strength, and probably

could not have been carried to a large enough strain by triaxial

testing. It may be that at larger strains the A curve would drop

to a steady state (residual) shear strength which is considerably

lower, particularly in over consol i dat ed samples. The tests also

provide no information on cumulative d i sp 1 acemen t s that may

result from many strain cycles in one direction. Such information

is of crucial importance for the assessment of the long term

response of offshore structures subjected to ice loading.

Since it is difficult to obtain truly undisturbed samples and to

measure the shear strength in situ, it is important to provide

information that will enable us to estimate the she>r strength

from soil parameters that can be measured in situ. Figure in

28
gives the range o-f correlations between the unconsolidated

undrained shear strength and the natural water content Wn , which

is a measure of in place density and is not as sensitive to

sample diturbance as tri axial test results. Figure 19 provides

information on the range of shear strengths as a function of

depth which would be obtained for normally consolidated deposits.

This would be a conservative lower bound, since the deposits are

overconsol idated.

An example of soil profiles in Harrison Bay presented by Bea

Cl 13 is shown in figure 20. Note that there is a variety of

sedimentary deposits, considerable variation of undrained shear

strength with depth and location, and that the shear strength of

the silts tends to be higher near the surface.

The preceeding discussion applies to soil deposits below the

depth which is affected by seasonal disturbances resulting from

ice gouging and wave action. The bulk of the observed gouges are

less than lm deep C63. However, there are deeper gouges, and it

also must be assumed that the ocean floor below the bottom of the

gouges is disturbed by shear failures associated with gouging. In

the near shore area, the gouge marks are frequently obliterated

by wave action, but the sea floor is disturbed all the same. Thus

it is reasonable to assume that down to the 50m isobath the

seafloor deposts have been disturbed by gouging to a depth of 1,

and possibly 2m. As a result the soils are likely to be loose (or

soft) and their density non uniform. This factor must be taken

into consi derat i on , in addition to the gouges which are physically present.

2.3.3 Summary of Engineering Character i sti cs

In the preceding sections of chapter 2, the characteristics of

the unconsolidated deposits on the Alaskan Beaufort and Eastern

Chukchi shelves have been discussed. The engineering implications

29
Ll_ co*

CD

z

LU

cch

CO

CC. <

LUX

COQ

LU<

X

Q

Z3

Figure 18. Water Content vs. Undrained Shear Strength

Beau-fort Sea Silts C9H.

30

f or
DEPTH

BELOW

MUDLINE

(ft)

UNDRAINED SHEAR STRENGTH (KSF)

0 1 2 3 4 5

Figure 19. Pro-file of Expected Range of Undrained Shear Strengths

for Normally Consolidated Alaska DCS Silts 119].

31
DEPTH

(ft)

DEPTH

(ft)

SITE A SITE B

SH (KSP) SHEAR STRENGTH (KSP> oits oias

SITE C SITE D

SHEAR STRENGTH (KSF)

0 12 3

SHEAR STRENGTH (KSP)

0 12 2

Figure 20. Soil Pro-files in Harrison Bay Clll
of these characteristics are summarised hereafters

(1) Evidence of Instability (2.2.2).

While the youthful landslide terranes tend to be close to the

shelf break, and thus beyond the depth range considered in

this report , their absence should not be taken for granted.

If creep is taking place along a weak layer which is

located at some depth, it may not be possible or desirable to

prevent the gradual diplacement of structures supported on

mat foundations. Flexibility or adjustments may have to be

provided to accommodate anticipated displacement. Scour and

erosion in coastal areas, including a rapidly receeding

coastline, strudel scour near river deltas, and landward

migration of barrier islands may also pose significant

engineering problems in shallow water depths.

(2) Seismicity (2.2.3)

Except for the limited area outlined in figure 5, there seems

to be no significant seismic risk. For the proposed Magnitude

6.5 design earthquake, liquefaction of granular deposits may

pose a risk, particularly for artificial berms and Flolocene

sand deposits which may be naturally loose and difficult to

compact. High plasticity silts are probably not liquefiable,

but low plasticity silts may pose some risk.

(3) Relict Permafrost (2.2.4)

Relict permafrost, where present, poses a troublesome and

unique engineering problem. It can be seen from figure 21

(from Ref . C 1

1

1 ) , that the problem occurs over a wide area.

Since offshore platforms and exploration and production

wells are likely to change the thermal regime underneath the

mat foundation, at least partial thawing may be unavoidable.

The effect would be large settlements and loss of subsoil

shear strength. Mitigation or prevention of these effects

must be a design consi der at i on

.

(4) Shallow Gas and Gas Hydrates (2.2.4)

The pressure exerted by shallow gas deposits may considerably

weaken the shear strength of soil deposits. The presence of
1 Continuous Ice bound bonded sediments

2 Discontinuous ice bonded sediments

3 No ice bonding observed

{Relict ice-bearing permafrost generally

I below 50m

Relict ice-bearing permafrost widespread

land extremely variable in depth

Ice-bearing permafrost

generally present at 1-20m

Figure 21. Provisional Map of Subsea Permafrost Distribution

C11D.

34
gas also causes sample disturbances which make it difficult

to assess in situ soil character i sti cs. In addition, shallow

gas can cause problems during drilling operations which may

have to be considered. Gas hyderates may cause serious

problems when changes in the thermal regime cause thawing of

the enclosing permafrost. In accordance with available data,

gas hydrates may not be encountered landward of the 60m

i sobath

.

(5) Ice Gouges (2.2.4)

Ice gouges, and the associated irregularity of the geometry

and strength character i sti cs of the seafloor deposits must be

considered in the design of mat foundations. Resulting

engineering problems are partial contact of pref abr i cated mat

foundations and uneven stress distribution even when underfilling or low berms are used. Gouges are likely to be

encountered landward of the 50m isobath and are most severe

between the 15 and 45m isobaths (the Stamuki zone). The

dominant orientation of the gouges is parallel to the

i sobaths

.

(6) Shear Strength of Soils (2.3.2)

The Holocene deposits are generally loose and quite variable

in characteristics. The Pleistocene silts, which underly much

of the area considered in this report are by and large

competent. The denser silts tend to dilate under moderate

shear deformation and their shear strength increases as they

dilate. To a depth of 10 - 12 ft the silts are heavily

overconsol .i dated , probably due to effects of past freezing.

The soil conditions are quite irregular and lenses of much

softer material within the silt layers are not uncommon.

(7) Sand and Gravel Resources (2.3.1)

There are some sand and gravel deposits that were mapped. In

general , the rate of sediment deposition in the river deltas

is quite low and there are many areas with no convenient

access to sand and gravel deposits.
.
DESIGN LOADS

3. 1 General

The arctic o-ffshore environment produces unique environmental

loads, Including the thrust exerted by ice and dynamic loads

generated by collisions with large bodies of -Floating ice, and

loads that can result -from thawing o-f relict permafrost or from

cycles of freezing and thawing. The optimal engineering solution

is not always construction of an immovable object that will be

strong enough to resist the loads. Sometimes it may be possible

to reduce loads by permitting the structure to yield or move, to

avoid loads by protective measures, or to provide the means to

augment the resistance of structures if such measures should be

needed in the future.

Not all loading conditions are relevant to the design of mat

foundations. For instance local stresses induced by ice loads are

much higher than the average stress on a large area which

produces the global ice force. Wind and wave impact loads may

be very significant in the design of equipment and structural

components, but they do not produce critical loads for foundation

desi gn

.

3.2. Loads Acting on the Structure

( 1 ) Gravity Loads;

Gravity loads include the weight of the structure, any ballast

applied by flooding of cells or by other means to increase the

gravity load, operational loads associated with drilling and

storage, and any ice or snow loads that may accumulate.

In the design of mat foundations, The minimum gravity load is

generally of interest. If minimum loads are determined, consideration should be given to the probability that gravity loads may

be reduced , and to the possible effect of tidal uplift forces

37
exerted on ad -frozen ice.

( 2 ) Win d 1 pads:

Basic wind speeds near the Alaskan Beaufort and Eastern Chukchi

coast as specified in ANSI A58. 1-1982 C133 are between 90 and 100

miles per hour. For foundation design, wind load effects on the

structure per se are not significant. However, wind loads acting

on ice floes or rubble piles which bear against the structure can

exert significant forces and are to a significant extent responsible for the fluctuation of ice forces with time. The maximur

global ice forces used in the design of the foundations, as

calculated by the present state of the art, depend on the

strength, failure mode and sometimes on the velocity. of movement

of the ice, rather than the magnitude of the wind forces acting

on the ice. Even though wind forces are not used in these

calculations, they may be ultimately responsible for the movement

of ice floes which produce critical collision forces. The

dominant wind direction during late summer and early fall is

Northeast, but high Westerly winds occur during storms Cll.

( 3 > Waves. Tides, Currents, and Storm Surges;

Surface waves, which are restricted to the open water season are

generally small because of the limited fetch resulting from

offshore sea ice. They have generally 2 to 3s periods and heights

less than lm. The maximum wave heights reported in the Beaufort

Sea were over 9m and at F'ointe Barrow 6m and occurred during

summer storms [143. Lunar tides along the Beaufort coast of

Alaska are of the order of 0.5m, but low barometric pressures

and high Westerly winds which prevail during exceptional storms

can cause storm surges which are 3m above mean sea level Cl J. At

Barrow and in the Chukchi Sea these surges can reach 3.5m.

Current patterns and velocities are shown in figure 22.

< 4 ) Seismic Forces:

As noted in chapter 2, much of the area under consi derat i on i .

38
t.s Speed in knots «— Surface current pattern

Beaufort Gyre (near surface) Bottom current pattern —i wind dependent near shore currents <-o Alaska coastal current

igure 22. Current Patterns Recorded in the Beau-fort and Eastern

Chukchi Seas (-from References Cl] and C14D).

39
considered aseismic. Lateral forces generated by seismic loads

would probably not exceed the ice forces, however there are two

instances which under certain circumstances could be critical for

designs interaction between the structure and the surrounding

fast ice during a seismic event; and liquefaction of soil

deposits supporting the structure.

( 5 ) Ice Loads;

The most critical loading condition for mat foundations for

Arctic offshore structures is a combination of minimum vertical

and maximum lateral load. The maximum lateral load, in turn, is

likely to be the global force exerted by the most critical ice

loading condition. The conditions determining the magnitude of

ice loads are difficult to define and predict with any degree of

certainty. Thi is due to the difficulties in predicting seasonal

ice regimes, the problems encountered in defining the strength of

ice, and the many different ice features which can cause critical

1 oadi ng

Ice regimes and features were qualitatively discussed in Section

2.2.4 (3). API Bulletin 2N C153, defines the following ice

morphology features: 1. Sheet ice is the fast ice in protected

bays and lagoons; 2. Rafted ice are two or more sheets of stacl ed

ice which rapidly consolidate into a single unit and may appear

as smooth-surf ace floes on either side of first year ridges:

3. First year ridges are ice ridges formed by compressive forces

in a single season; 4. Rubble features (rubble piles) are

grounded ice ridges, and can also form around offshore struc

tures; 5. Multiyear floes are bodies of floating ice which

survived more than one season. They have weathered rounded and

consolidated sails and relatively solid keels about 3-4 times the

sail height. They can incorporate ridges and rubble piles and in

many instances have length and width dimensions in excess of

300m. 6. Arctic Ice pack is the floating pack ice normally

present seaward of the 60m isobath which may occasionally invade

40
offshore structure sites during the summer season; 7. Natural ice

islands are tabular icebergs calved off the Ward-Hunt ice shelf,

which can be greater than 250 km3 and can have a keel thickness

of 45m. They tend to drift with the Baufort Gyre and fragments of

the islands occasionally invade shallower waters. The last

recorded invasion was in 1973 C16I1

Some quantitative data for the Beaufort Sea were presented by

Vaudrey C 1 7 j s Ice Coverage:

Ice Thickness:

Sheet Ice

Rafted Ice

First-Year Ridge

Extreme

Multiyear Ridge

Multiyear Floe

Early October to mid July

Rubb 1 e Piles:

2m (6-7 ft) Consol id. Thick. 3~4m

5-

6m Max. Height 11- 13m

8-9m Max. Length 300m

> 30m

13—20m

6—

8m

Ice Movement (max)

<10m Isobath

10—20m Isobath

>20m isobath

Open Water

3-6m/h

30— 1 50m /h

300m /h

2-4 km/h

The mechanical properties of sea ice are complex and will not be

discussed in detail. Unconfined compressive strength as obtained

in laboratory tests depends on strain rate, crystal orientation

and crystal grain shape and size. Confined compressive strength

is substantially greater than unconfined strength. Tensile

strength, flexural strong th , shear strength, and Young's Modulus

vary with the brine content, v to , which is defined by the following equation:

vb = S (0.532 - 49.185/T) ... (eq 1)

where: S = ice salinity in parts per thousand (ppt)

T = ice temperature in “C

vto = brine volume in ppt

41
Typical data obtained -from Ref [16] are shown in figure 23. Ice

strength and other physical char acter i st i cs change in a complex

way with temperature, with the strength increasing with decreasing temperature, but also brittle behavior becoming more

pronounced as temperatures fall below -10° C. This affects the

correlation of ice strength to the size and shape of the contact

area between ice and the walls of an offshore structure in a complex way. This problem was discussed by Bruen et al.[19], Correlations between the contact area and pressures from first and

multiyear floas developed in Reference [19] are shown in figure

24. Note that these curves are derived from a specific set of

assumptions. The correlations may be different if any of the

assumed parameters of the problem are changed.

Design forces are stipulated in Reference C15]. Watt [18]

discusses two types of offshore structures which fall within the

scope of this reports Wide indenters, which are wide structures

with vertical walls, and cone structures. Wide indenters are

defined as structures with waterline diameter to first year ice

thickness ratios of 20 : 1 or more, and with sides which are

either vertical or deviate from verticality by no more tha 20°.

Critical global ice pressures are generally controlled by ice

failure and depend on the char ac ter

i

st i cs of the ice feature

exerting the pressure. The critical load on a wide indenter could

be controlled by the "breakout" from an ice sheet, or the

crushing into a floe in which case the strain rate would have a

critical effect, or the kinetic energy of a colliding ice

feature. Relatively thick ice features such as ridges could tail

by flexure reather than crushing. Cone structures can limit the

ice load by causing the colliding ice feature to ride up on their

side and fail in flexure, rather than by crushing. However

this failure mode can only occur if the adfreeze bond between the

structure and the ice is limited. This can be accomplished by

special coatings or by heating the sides of the structur e. I> e Pressures against conical structures can be increased by the

42
HoH

root*

1

Uncooflned

Co«f>rttt1ve

Strength

(

PSI

)

Figure 23. Typical Strenrgth Characteristics o-f Sea Ice [15]

43

MPa
PRESSURE

(KSI)

Fi gur

AREA (m 2

)

1 10 100

3 24. Correlations Between Ice Pressures and Contact Art *

C 1 9 3 .

44

PRESSURE

MPa
rideup of ice against the sides of the structure.

The implications of the effect o-f di-f-ferent ice environments on

global design loads are illustrated by an example presented

by Bea et al . C203. Figure 25 shows a plot o-f global ice -forces

against a large indenter as a -function of the mean recurrence

interval of the loads. The vertical scale is not labeled since

the magnitude of the actual force would depend on the diameter of

the structure. Note that the extreme event of a collision with an

ice island is judged to have a low probability of ocurrence, and

that actual observations at Hans Island [21] indicated that

estimated loads associated with this event may be much too

conservat i ve . Figures 26 and 27, which are also taken from

Ref [203, show design loads estimated for 120 to 180m wide

exploration and production structures as a function of water

depth. In the figures UB is an upper bound estimate and LB is a

lower bound. A 25 year return interval is taken for the exploration structure in figure 26, and a 100 year return interval for

the production structure in figure 27. Note that the estimated

global ice loads for the upper bound estimate increase with

depth. These loads are calculated for collisions with multi year

floes and are sensitive to the strain rate. As the water depth

increases, the velocities of the ice movement, and thus the

strain rates increase.

While the global loads previously dicussed are extreme loads

which control the required load capacity of foundations, It is

also important to consider long term and cyclic loads which

could cause cumulative displacements, and the dynamic characteristics of loads which could cause excess pore water pressures.

Figures 28 to 30 are taken from Ref C21D and show a winter

season ice pressure record from the Beaufort Sea, a calculated

ice impact load signature, and a typical winter ice loading

histogram. The cyclic loads in figure 28 primarily reflect storm

wind effects on the ice canopy (low frequencies) and local ice

45
GLOBAL

ICE

LOAD

Figure 25. Global Ice Forces as a Function of Mean Recurrern •* Interval C20D.

46
TOTAL

LATERAL

FORCE

(KIPxlO

3)

WATER DEPTH (m)

0 20 40 60 80 100

6

5

4

3

2

1

0

coO

IDo

DCo

<

DCID

<

t-o

Figure 26. Estimated Global Ice horces Against a 120— 180m Wide

Exploration Structure as a Function of Water Depth C20H.

47
WATER DEPTH (m)

0 20 40 60 80 100

Figure 27. Estimated Global Ice Forces Against a Production Structure as a Function o-f Water

120-180m Wide

Depth C20D.

48

TOTAL

LATERAL

FORCE

(MN-10
ICE

PRESSURE

Figure 28. Winter Season Ice Pressure Record -for

Beaufort Sea [2 ID.

the Canadian

49
Figure 29. Calculated Force-Time Histories for Large Ice f . -tur»>

Impact C21 3

.

50
PERCENT

OF

MAXIMUM

FORCE

F i gure

100

75

60

26

1 cycle

3

cycles

- 5

cycles

" 10

cycles

Nov Jan Mar May

6

cycles

30. Typical Winter Ice Loading Histogram C211.

51
•fracturing (high frequencies) . It is important to note, that the

di rect i onal i ty o-f the ice -forces discussed above is not random,

since it is a-f-fected by prevailing wind and current directions. If

residual displacements from ice load cycles are anticipated, it

is likely that these displacements will result in seasonal , or

even cumulative annual lateral drifting of the structure. These

lateral displacements must be considered in the design, particularl

for permanent drilling and production structures. It should also

be noted that the high frequency cyclic forces could lead to

liquefaction or cyclic mobility of the underlying or core infill

soils, particularly in the case of low density, hydraulically

placed cohesion less soils.

( 6 ) Special Environmental Loads:

Some special structural loads could be generated by the arctic

environments 1. Settlements caused by the thawing of permafrost

or oil and gas removal could create voids under large, relatively

stiff mat foundations, resulting in shear forces and bending

moments in the mat. 2. Release of trapped or hydrated gases could

in extreme cases induce uplift forces, followed by settlement.

Trapped gases can also reduce the shear strength of supporting

soils. 3. Soil erosion could be caused by wave and current action

and by strudel scour (erosion caused by the rapid flow of melt

water through openings in the ice sheet) and can result in soil

removal at the fringes of mat foundations or gravel berms,

causing high local stresses in the mat foundation or local

stability failures. 4. Freezing could occur under a mat founda

tion as a result of changes in the thermal regime or as a seasonal phenomenon. Freezing will cause uplift forces, and

subsequent thawing could leave voids under the mat foundation.
4. DESIGN LIMIT STATES

4. 1 General

In accordance with accepted usage of the term, a "limit state"

occurs when a structure fails to fulfill the functions or meet

the conditions for which it is designed. In present design

practice, two types of limit states are generally defined; "ultimate" limit states, which are associated with loss of life or

major damage, and "servi ceabi 1 i ty " limit states which are

associated with transient loss of function or minor damage. The

objective of limit states design is then to assure that ultimate

limit states, which are generally associated with strength, have

a suitably low probability of occurrence, and that ser vi ceab i 1 i ty

limit states, which tend to be associated with stiffness, not

occur so often as to impair the functionality of the structure or

unduly increase its maintenance cost.

The previously discussed approach uses two general categories

of limit states, ultimate and ser vi ceab i 1 i ty , and generally

provides more or less uniform safety margins against the ocurrence of these limit states, either by probabi 1 i st i c code

formates, or by more traditional design schemes (working stress

or factored loads and resi stances) . This approach is too simplistic for the complex environment and the innovative technology

associated with offshore structures in arctic regions.

By definition a limit state is a failure mode, and since each

structural scheme tends to have its own failure modes, limit

states are scheme dependent and can not always be generalized.

Offshore technology in arctic regions is evolving rapidly and

many different concepts are under consi der at i on , some of which

are designed to avoid, rather than resist loads. Furthermore,

some of the loads that could be encountered during summer

invasions of ice islands are of such a magnitude that it would be
in many instances unfeasible to design adequate foundations. To

provide adequate guidance for structural design , which would not

unduly restrict the introduction of new, innovative emgineering

concepts, criteria for required structural performance need to be

stated in more fundamental terms.

4.2 Requirements for Safety and Servi ceab i 1 i ty

The requirements for safety and servi ceabi 1 i ty of the structure

in place can be expressed as performance criteria:

1. Protection against loss of life and severe injuries.

2. Protection against major environmental damage

3. Protection against major property damage

4. Protection against minor environmental damage

5. Maintenance of efficient operation

6. Acceptable maintenance and replacement costs

Another set. of criteria is needed to insure contructabi 1 l ty in

the Arctic environment. Criteria 1 to 3 would be normally lumped

together and satisfied by the safety margins against the ultimate

limit states. However in this instance this is not necessarily an

axpedient way to address the problem. A better way to look at the

problem is by stating that Criteria 1, 2 and 4 concern not only

the owner of the facility, but are also a concern of government

at various levels, adjacent property owners, residents in the

general area, and possibly other countries. Criteria 3, 5, and 6

primarily concern the owner of the facility and could be addres

sed in purely economic terms. This distinction is important,

because protection against loss of life and major environment 4

damage could conceivably be accomplished even if the structure i permitted to fail under, or designed to avoid extreme lateral

forces, such as those resulting from a collision with an n - island.

The six performance criteria listed above must be stated in tor.ru

of acceptable failure pr obab i 1 i t

i

es . The question there* nro

54
arises, how such probab i 1 i t i es can be determined. Figure 31,

which was presented by Whitman in the seventeenth Terzaghi

lecture [22], is based on evaluated risks to structures and other

engineering projects, and gives some indication o-f accepted risks

in present engineering practice. The ratio between the cost

o-f the loss and the accepted risk, obtained from a cost benefit

analysis for Arctic offshore structures, may differ from the

trend shown in the figure because of the unique problems associated with the Arctic region. Nevertheless, the figure gives some

indication of what may be acceptable to our society in conjunction with the "ultimate" limit states.

In the figure, an annual failure probability of the order of 10~ 3

is shown for fixed drill rigs. However, this probability primarily addresses the rigs in the Gulf of Mexico, where the cost per

drill rig is lower and the consequences of environmental damage

are more manageable and less severe. For production platforms in

the Arctic region, an annual probability of catastrophic failure

of the order of lO-4 or less would probably be more in line with

the financial and environmental consequences. Examination

of figures 25 and 27, which give some insight into design

criteria actually used indicates, that a design for the 100 year

ice load cannot provide this level of reliability in locations

where ice island invasions can occur, unless some mechanism is

provided by which extreme loads can be either avoided or prevented from occurring. An additional factor which must be taken into

consideration is the limitation imposed by the lateral load

capacity of foundations which will be discussed later.

In the case where economic consi derat i ons govern, such as for

instance for Criterion 3, a cost benefit analysis would be more

appropriate than the imposition of some arbitrary failure

probability. The methodology used in such an analysis is i 1 1 ustrated in figure 32, which is taken from Ref C23]. In view of the

fact that the cost of an offshore production facility could
ANNUAL

PROBABILITY

OF

"FAILURE

COST IN $ 1m 10m 100m 1b 10b

CONSEQUENCE OF FAILURE

Figure 31. Risk -for Selected Engineering Projects (tal cn

Ret. C 22 1 )

.

4 rom

56
TOTAL

EXPECTED

COST

(millions

of

dollars)

EXAMPLE - VALUE ANALYSIS

PROBABILITY OF FAILURE (%/year)

Figure 32. Probability of Failure Obtained from a Cost Benefit

Analysis (taken from Ref. C23]).

57
exceed $ 1 billion, a cost benefit analysis may also result in a low -failure probability. Thus the design capacity against lateral

loads may be limited by our ability to build lateral -load

resisting structures and foundations, and the design solution may

have to be supplemented by strategies which would allow us to

avoid or prevent exposure to extreme ice loads and the consequences

associated with undesirable -foundation performance.

Acceptable risks -for exploration structures would be considerably

higher than those for production structures, because these

structures are less costly, and the envi ronmental impact of a

catastrophic failure would not be as high as in the case of

production structures. The period of exposure of these structures

in any one location and their useful life are also much shorter

than those of production structures. The data in figure 31

indicate a failure probability in the order of 10-2 for existing

mobile drill rigs, and that this failure probability seems to

exceed even marginally accepted reliability in present engineering practice. Again it appears that for Arctic installations

this failure probability would have to be lower, even if dictated

by purely economical considerations. As these structures are

usually designed to be mobile, it may not be very problematic to

plan on avoiding extreme ice loads which tend to occur during the

summer open water season. Annual failure probabilities of the

order of 10—3 may be reasonably compatible with accepted engi

neering practice.

Failure pr obab i 1 i t i es associated with criteria 4, 5, and 6 would

tend to be dictated by cost benefit considerations, except that

in the case of criterion 4 there may be instances where even

minor en vi ronmental damage may not be acceptable. The difficult

i es with these criteria may be in the prediction of environmental

effects which would cause failures, such as cumulative displ

ments, erosion, local force concentrations, and possibly dynamic

effects.

58
4.3 Structural Concepts

4 . 3 . 1 Bac k g r oun d Inf or mat i on

Since the technology of Arctic drilling and production is in a stage o-f rapid evolution, many di-f-ferent concepts are under

consi derat i on or in different stages of design and construction.

For instance Buslov and Krahl [24] discuss 51 new concepts.

There is a record of past successful experience with gravity

structures in hostile offshore environments. Concrete lighthouses

have been used in the St Lawrence Seaway and the Baltic Sea

since the turn of the century, and concrete gravity platforms

have been installed in the hostile waters of the North Sea since

the early 1970's C25]. Past offshore exploration in the Arctic

was generally confined to shallow waters and was carried out

either from the ice or from artificial islands constructed from

locally dredged materials.

More recently different exploration structures for use in deeper

waters were built by LJ.S. and Canadian companies, including

caisson retained islands, mobile gravity structures and floating

vessels. Many other concepts for exploration and production

structures are in various stages of planning and design.

4.3.2 Gravity Structures Used or Considered.

This report deals specifically with design limit states for

foundations. Thus only those features of the various structural types which are relevant for foundation design are discussed.

There are essentially three types of foundations associated with

the structures under consi der at i on : foundations for earth

structures, for artificial islands (though the island also

incorporates a variety of earth supported structures) ; foundations for caisson-retained earth structures; and mat founda—

59
tions supporting rigid structures. Since the design limit states

for these foundations do not only depend on the foundation type,

but also on the geometry, flexibility, and vul nerab i 1

i

t y of the

supported structure, it is necessary to discuss the characteristics of these structures in some detail.

1. General Character i st i cs: the offshore structures discussed in

this report are either exploration or production structures.

Exploration structures are either temporary in nature, such as

dredged artificial islands which are used in the winter sason

and later erode, or mobile structures which are re-used after

completion of the exploration. Exploration structures need a working area of about 100m diameter and are often designed for

loads with a 25 year mean recurrence interval (even though

this philosophy may be questionable for re-useable structures

[253). Production Structures are normally built for a service

life of 25 to 30 years and require a working area of

approx i matel y 200m diameter (unless they are multi level). Non

Arctic offshore production platforms are normally designed for

loads with a 100 year recurrence interval.

2. Artificial Islands? Most of the artificial islands in the

Alaskan Beaufort have been built in water depths of about

7m. The gratest depth is 13m C26J. The materials used are

mostly gravels. Construction methods include barge haul .md

winter haul over ice roads. On the Canadian side, there is

also a considerable amount of sand pumping by dredges and

bottom dumping from barges. Artificial islands become uneco

nomical in depths greater than 20m. Design slopes depend on

the grain size of the construction material, the method of

placement, and whether the slope is in the upper zone (+ m to

top)

,

the wave action zone, or the bed level zone < 'm and

deeper). Since erosion and wave runup are major problem.,

sacrificial berms or other means of slope protection have to

be used. In the Alaskan Beaufort, slope protection generally

60
consists of 2-4 cubic yard sand bags. Figure 33 which is taken

from Ref. C26] shows different slope treatments used. In the

Canadian Beaufort, where dredging is extensively used, slopes

are flatter (typically 1:15) and erosion protection is often

provided by sacrificial berms.

Caisson Retained Islands (CRI): These structures have a sand

core which is contained by peripheral caissons. Some of these

caisson structures are sectional, such as the Tarsiut Island

structure shown in figure 34 (Ref .27) , which has a sand core,

contained by sectional concrete caissons and rests on a sand

berm in a water depth of 21m. Others have a continuous caisson

shell around a sand core, such as the Gulf Mobile Arctic

Caisson (MAC) which has an annular steel caisson and is

designed for water depths of 15 to 40m. This structure is

discussed in Ref. [28] and is shown in figure 35. Other

structures of this type are the Esso stressed steel CRI

and a planned Sohio concrete CRI [29]. The concept is also

under consi der at i on for production facilities [30].

4. Rigid Base Structures: Under this heading many different types

of offshore structures are lumped together because they rest

on a rigid structural mat, which either rests on a prepared

berm or is placed directly on the ocean floor. These structures have either vertical sides which are designed to resist

ice pressures as wide indenters, or sloped sides which are

designed to induce flexural failure in the ice. Several types

of gravity structures are included in this category: (1) Water

ballasted caisson structures. Included in this category is the

vertical sided Super CIDS, built by Global Marine for Exxon,

which is shown in figure 36 (taken from Ref. [31]). The latter

structure was installed with a protective ring of grounded ice

to provide protection against extreme ice forces. The BWACS

cellular floating caisson is also vertical sided and was

developed by Brian Watts Associates and commissioned by Zapata

61
SHEET PILE OVERTOPPING PROTECTION

(b)

SHEET PILE

I- i gure Typical Shallow Dredged Island Cross Sections

•from Ref . [26 3 ) (taken

62
All dimensions

in meters

Berm Sand; , oxClay A Silt

Bonded Sand

- 4.5

- 0.5

r

JT

-55

-64

Silt A Clay

SECTION

TARSIUT ISLAND DESIGN LAYOUT

(PRE-CONSTRUCTION)

Figure 34. Tarsiut Island C273.

63
Figure 35. The Gulf Mobile Arctic Cai sson (MAC) C 28 1

.

64
l-igure 36. Ex x on '

s

Super C I DS E 3 1 ]

.

65
Offshore Services and is shown in -figure 37 (-from Ref.C313)

Sloped sided water ballasted caisson structures include the

Arctic Cone Exploration Structure C33H which is designed for

heavy ice and for water depths from 15 to 33m. The latter

structure is designed to be placed on the unprepared ocean

floor. It has short skirts and a sand undergrout system that

permits injection of sand into the voids under the base. This

sand can then be further consolidated by ballasting. The

structure is shown in figure 38. Another, "second generation" concept in conical caissons, the SONAT hybrid arctic

drilling structure is discussed in Ref. [34]. The structure has

a steel base to provide some foundation flexibility and a concrete mid section, and was designed with special attention to minimising ice loads and ice runup effects.

(2) A second category of rigid based structures is equipped

with a wide rigid base to maximize soil resistance and a relatively narrow stem to decrease the width exposed to ice

forces. These structures are often referred to as "monopode"

structures. The stem is normally equipped with an icebreaker

cone. An example of such a structure is the adjustable

monocone rig which was developed by Esso Resources of Canada

and is shown in figure 39 C35D. Various models of these

structures were designed to operate in water depths ranging

from 10m to 60m. (3) Miscellaneous other structural concepts

have been developed. These include: 1. The Sohio Petroleum

Arctic Mobile Structure (SAMS) which is shown in figure 40 and

discussed in Ref. [361. This structure is an octagonal water

ballasted concrete barge which is equipped with a system

capable of inserting and retrieving 7m Diameter steel pipe

caissons (spuds) to augment the lateral load resisting

capacity of the mat foundation. Sophi st i cated monitoring

systems keep track of the vertical and lateral loads and

def ormations. It is planned to insert the spuds only if the

need arises. 2. A proposed detachable monocone production

structure C373, which consists of a steel cone, supported b> <

66
BWACS

Figure 37. The BWACS Floating Caisson L31].

67
Figure 38. The Arctic Cone Exploration Structure (ACES) [531

68
250 FT

STEEL

160 FT

CONCRETE

350 FT DIA

Figure 39. Adjustable Monocone Rig C353.

69
n

i l

i

i

i

i

l

'

l

i

igure 40. Ihe Sohio Arctic Mobile Structure (SAMS) [361.

70
steel or concrete 130m diameter circular mat -foundation. The

cone is designed to disconnect -from the base, so it can be

removed during a rare ice event. 3. A stacked annular caisson

system -for yearround drilling in 4 to 20m water. The stacked

sectional caissons have a center core which can be -filled

with sand. The sand core is dewatered to increase the shear

strength o-f the sand. 4. Some concepts for structures capable

of resisting ice island impact were also proposed C39, 403.

4.4 Design Limit States for Foundations

4.4.1 Definition of Limit States

Two categories of limit states are defined for the foundations.

Ultimate limit states are failure modes causing major damage to

the supported structure (either total loss or long term interruption of service). Included in this category are not only

stability and truncation failures, but also settlements and

d i sp 1 acements which would precipitate a major structural failure.

Serviceability limit states are failure modes which cause

short term interruption of service or damage which requires

repairs or replacements which can be accomplished without

replacing a major part of the structural system. Included in this

category are local stability failures, settlements and displacements which require remedial measures ,and erosion.

It is necessary to distinguish between the two categories of

limit states because of the much greater safety margins necessary

to protect against ultimate limit states. However in practice the

dividing line is not necessarily clear cut. For instance settlements caused by thawing of relict permafrost may be an ultimate

limit state when no remedial measures are available to mitigate

the effects on the supported structure, and a ser vi ceab i 1 i ty

limit state when provisions are made to mitigate the settlement

effects by underfilling.

71
4.4.2 Practical Limits to Load Resistance

Overal 1 requirements for safety and ser vi ceabi 1

ty were discussed

in Section 4.2 and it was recognized that safety margins which

would be consistent with present engineering practice cannot be

realized because of physical limitations to the lateral load

resistance of gravity structures. The limitations to the load

resistance that can be provided by mat foundations are illustrated by the following example: Let us assume that it is desired to

locate a cylindrical structure of diameter D on a stiff mat

foundation of diameter D + 50 ft., that the design ice load is 700

kip/ft and that the strength of the foundation is determined

using a load factor of 1.3, and a resistance factor of 0.8. Then

the ratio of resistance to load, R/S, would have to be 1.6. Let

us further assume that the structure should be able to resist

the extreme ice load of 1000 kip/ft. with a load factor of 1 and

a resistance factor of 1 = 1). These assumptions are

reasonably consistent with present engineering practice. Let us

now assume that the soil conditions are similar to those shown in

figure 20. Even in the most favorable location, the soil resistance would not exceed 2 kip/ft 2 . Thus the assumption that

the lateral load resistance cannot exceed 2 kip x the area of the

mat foundation would not be overly conser vat i ve . The implication

of this assumption is shown in figure 41, where R/S and Ru /Su are

plotted against the required diameter of the mat foundation. It

can be seen that even a 600 ft. (200m) wide mat foundation could

not satisfy the design criteria. But as the foundation width

increases, so does the required thickness of the mat and the

logistic difficulty of building, transporting and erecting the

structure. There are various methods by which the soil resistance

can be increased, ranging from skirts and spud piles to compac

tion grouting. But these methods, which will be discussed later

in this section, also have economic and physical limitations.

Thus various strategies were adopted to limit or avoid e:: posures to extreme loads, or to mitigate failure consequences

72
R/S

Ru/Su

D, ft

Figure 41. Load Resistance of Mat Foundation

73
4.4.3 Design Strategies

Possible design strategies tall into three categories:

(1) Brute strength. The difficulties with this approach are

obvious and have been discussed.

(2) Hazard Mitigation. In this strategy, the performance criteria

in Section 4.2 are satisfied by mitigating hazards by methods

other than structural load resistance. Several such strategies

have been used: 1. Personnel Protection

2. Well Protection

3. Structure Protection

4. Ice force Management

These strategies can be used separately or in combination to get

the desired level of reliability. A good example of using

strategies 1. and 2. is Tarsiut Island. It is realized that

extreme ice forces could cause a catastrophic failure. Rather

than trying to provide load resistance against these extreme

forces, it was decided to mitigate the consequences of potential

failures [413. For this purpose, four alert levels were instituted, ranging from normal operations to total evacuation after a safe shutdown of the wells. An example of Strategy 4., ice force

management is the protective ring of grounded ice installed

around Exxon's Super Cids. Other measures, such as splitting

dangerous ice features before they impact the structure are under

consi der at i on . Examples of 3., structure protection ire the

sacrificial berms in the artificial islands, and also the

capability of the floating caissons to be moved from a particular

location during the critical summer season. Another fora- of

structure protection are energy absorbing devices developed to

protect offshore structures from the effects of floating ice

bergs. After the collision, these devices can be restored tu

their original position C 39 3 . Weak ductile links in the structui e

or yielding foundations could also protect the structure.

74
3. Adaptive Controls. These strategies involve the design o-f the

structure to a reasonable nominal strength level, with provision

for a capability to increase load resistance in the field in

accordance with sensed or observed structural performance. A good

example of adaptive controls is provided by the Sohio SAMS

structure shown in figure 40. The structure is instrumented with

sophisticated monitoring systems and it is planned to install

the spuds only if ice forces are expected to exceed certain

critical levels. The spuds can be installed and retracted

relatively rapidly. A second adaptive system on the SAMS is

associated with expected settlements due to thawing of relict

permafrost. The remedial measure in this instance is sand

underfilling. Both the instrumentation and the projected settlements,

which are discussed in References C423 and [433, are shown in

figures 42 and 43. Adaptive controls could conceivably also be

used in other ways to modify foundation resistance, such as

ballasting and grout injection.

4.4.4 Failure Probabi 1 i t i es Associated With Limit States

Design procedures for gravity structures presently used generally

follow the FIR r ecommendat i ons C443, which address themselves

primarily to the North Sea conditions. In this document "extreme"

loading conditions, to be used with ultimate limit states, have a

mean recurrence interval of 50 to 100 years, and "normal"

environmental conditions have a mean recurrence interval of 1 month. Neither of these provisions is applicable to ice loads in

the Arctic, however a 100 year mean recurrence interval for

production structures and a 25 or 50 year mean recurrence

interval for exploration structures could be used if appropriate

design strategies are available to deal with specified extreme

ice features.

75
FOUNDATION INSTRUMENTATION

^ ACCELERATIONS, > DISPLACEMENTS,

CONTACT PRESSURES,

a INCLINATIONS . V OF MAT

INCLINOMETERS,

STRAIN GAGES,

8 THERMISTORS

IN SPUDS

( INCLINOMETERS IN

V CAISSONS

Figure 42. Instrumentation Used to Monitor

SAMS.

the Per f ormance of

76
Figure 3—7

THAW SETTLEMENT PROFILES CENTRAL WELL BAY CONFIGURATION

Figure 43. Settlement Control Planned -for SAMS.

77
The assumption can then be made, that if the appropriate probabi 1 i sti

parameters -for the loadings (demands) and resistances (capabi 1 i t i es

can be determined, -failure probabi 1 i t i es , conditional on the

ability to implement design strategies associated with extreme

ice features, can be calculated.

The following load and material factors are cited in the FIR

state of the art report on North Sea gravity structures 1453:

Ref erence Load Factor

for Env. Load

Mater i al

Cohesion or

Undrained str.

Factor

Ef f ecti ve

Friction

FIR (1977) 1 .

1 .

1.2

DnV (1977) 1.3 1.3 1.2

NF'D (1977) 1.3 1 .

1.2

These factors correspond to the European "semi probab i 1 i st i c

design approach. The "material factors" correspond to the inverse

of the capacity reduction factor used in U.S. practice. The

cohesion and undrained strength column mainly applies to silty

and clayey soils and to sandy soils subjected to dynamic loading,

while the effective friction column applies to sandy soils. It

can be seen that there is reasonable consensus on safety margins

and that global nominal safety factors range from 1.82 for

cohesive soils to 1.56 for sandy soils. Even though there is no

definitive credible set of data on which assumptions with respect

to the statistical charac ter i st i cs of the load and resistence

parameters can be based, it is possible to asses failure probabi

lities associated with the above mentioned global safety factors

when some engineering judgment is made with respect to load and

strength var i abi 1 i t i es.

The following closed form solution can be used to calculate

notional failure probability (P*) as a function of the global

safety factor, if the loading and resistance distributions arr

78
idealized as lognormally distributed variables CC46D (as modified

and supplemented) >:

F'-f = 1— 0{ (In CR/S (1 + V®22

) / (1+Vr2 > 3

1 ' 2 ) / C 1 n ( 1 - VBZ ) (1 + VR

2 )]^2 }

21 1 - <J> [In (R/S)/ (Vr2 + Vs52

) 1 ' 22

:! for VR , VQ < 0.3 ...(Eq.4.1)

where: <J> CXD = tabulated normal cumulative probability of the

standard variate, X (expression in brackets)

R = mean value of resistance

S = mean value of load

Vr = coefficient of variation of the resistance

Vs = coefficient of variation of loading

While it is not anticipated that credible data on the statistical

parameters in the above equation will be available in the

forseeable future, estimated values may give us some indication

of the order of magnitude of failure probab i 1 i t i es . For instance,

the following values are calculated:

R/S ! VR

i

i Vs 1

1 P*

1.82 : 0.2 ii 0. 2

1

1 .016

1.82 ! 0.3 ii 0. 2

11 . 053

1 . 69 : 0.2 it 0.2 11 . 03

1 . 56 : o. i

i

i 0.2 11 . 02

P«= in the table would be the failure probability for a time

period equal to the mean recurrence interval of the design

load. If it is assumed that for the North Sea a 100 year mean

recurrence interval was used for determining the design loads on

production structures, the calculated failure probabilities would

range from 1.6 x 10“^ to 5.3 ;< 10"“*.

Another factor that should be taken into consi der at i on is the

conservatism inherent in traditional geotechnical engineering

practice. Thus design values for undrained strength or the

79
angle of shearing resistance are likely to be somewhat lower than

the true mean values that should be used to calculate P* .

The approach used in the preceding calculations -for North Sea

structures cannot be applied to off shore structures in arctic

regions where the coe-f f i ci ent of variation o-f the load is o-f the

order o-f 100"/. (v„ 21 1 ^ - In that instance the annual P* could be

approximated as the reciprocal o-f the mean annual recurrence

interval o-f the load required to develop the ultimate capacity

(resistance) of the structure.

The foregoing discussion is intended to illustrate how probabilistic reliability approaches can be used, as an aid in determining

design loadings and load capacities. The computed notional P* '

should be interpreted as indices that are indicative of actual

P-p ' s . They should not be confused with the true safety or

reliability character i st i cs of a given structure. The notional

P-p

/ s have value in quantitative comparisons of design criteria

alternatives C76, 773. Extensive literature has been developed

during the last 20 years on the topic of reliability based design

criteria for offshore structures C78 through 833.

4.4.5 Design Limit States for Artificial Islands

Artificial islands have been constructed by three methods: Summer

construction by dredging (the fill material is taken from a nearby site and deposited directly by dredges); summer construction by bottom dumping from barges; and winter construction by

dump trucks supported by ice. The strength and grain sire

character i sti cs of the fill material, as well as the sideslopeof the island depend to a large extent on the construction

method used.

Dredged material normally ranges from fine silty sand to gravel.

Bottom dumping is only possible when D=0 > 0.28mm where Dn.-> i

80
the grain size below which 50°/ of the material by weight is

smaller, and when the silt content [material passing a #200

(0.075mm mesh) sieve] is less than 57. [47] . Thus a material

meeting these minimum requirements would have to be poorly

graded , or else the material would have to be courser. Trucked

material is gravel from on shore borrow sites.

Typical si deslopes used are summarized in the following table

which is taken from Ref.C26].

Mater i al ! Elevation ! Placement ! Side Slopes

Onshore

Gravel

! + 10 ft. to top !

! Wave Action Zone !

! Bed Level to -10ft!

Truck Haul I 1 : 1.5

! 1 : 8 to 1 s 10

! 1:3

Offshore

Sand and

Gravel

! + 10 ft. to Top 1

! Wave Action Zone '

! Bed Level to -10ft

1

Dredge ! 1 s 1.5

! 1 : 12 to 1 s 15

! 1:8

Offshore

Sand and

Si It

i + 10 ft. to Top 1

! Wave Action Zone !

! Bed level to —10ft!

Dredge ! 1:3

! 1 : 20

! 1:12 tO 1:15

Note that slopes are dictated not. only by strength and stability

requirement but also by construction feasibility.

Six generic categories of limit states need to be considered:

Major stability failure; local stability failure; excessive

displacements; erosion; wave overtopping; and runup of ice.

Whether these limit states are ultimate or serviceability would

entirely depend on the severity of the consequences of failure

and should not be prescribed a priori.

Stability Failures: Two types of loading can cause stability

failures: gravity loads, and ice push. Stability failures under

gravity loads will tend to occur toward the end of construction,

when the maximum gravity load is applied while the foundation

soil may not have had sufficient time to consolidate. The problem

occurs when the soils supporting the island are cohesive. Failure

could occur at the mudline where bottom sediments tend to be the

81
weakest, or at some depth below the mudline because consolidation

at that depth will require more time. To analyze this problem,

soil strength as a -function of time as well as depth must be

considered. Because of the peculiarities of frozen fills, slope

stability problems can also arise in steep slopes during winter

fill construction L48D.

Potential stability failure slip planes associated with ice push

are shown in figure 44 which is taken from Ref.C26D. The broken

line shows the depth of freezing. The frozen cap of the island is

likely to act as a rigid body, thus facilitating a major stability failure of this type. The slip plane could also be located at

some depth below the mudline. Watt C 483 points out that stability

problems should be analyzed as 3-D problems so that account is

taken of the island geometry. Slip surface 2 in figure 44 is

a truncation failure through the granular fill of the island,

which has relatively low density and therefore is sensitive to

even small cyclic loads such as those induced by wave action.

These cyclic loads can cause excess pore water pressures and thu c reduce the effective stress, and consequently the shear strength

on the slip plane. This possibility should be cosidered if ice

forces during the open water season could cause a threat. A

parametric study by Kontras et al.C491 in which various assump

tions were made with respect to the location of the bottom of the

freeze front indicates that the failure plain is likely to b<- located at the freeze front, even if it extends into the bottom

sediments. During the first winter, local, rather than global ice

failures are likely to occur (slip plane 1 in figure 44) since

sea ice is stronger than the unfrozen fill. In subsequent

easons the global stability failures pose a credible threat,

even though localized failure mechanisms may reduce the overall

ice forces.

82
FAILURE PLANES/MODES UNDER ICE LOADING

1. SLOPE FAILURE/EDGE FAILURE

2. TRUNCATION FAILURE

3. BOTTOM SLIDING FAILURE

Potential Fa i 1 Lit e Nodes uf Fill Islands

Re-f . C26D ) . Taken Figure 44 from
Another -failure mode not shown in -figure 44 is liquefaction. This

type of failure could occur in seismically active zones. It also

could be precipitated by other dynamic or rapidly applied loads.

However the latter possibility seems remote. Artificial islands

and other hydraulically placed soils are very vulnerable to

earthquake induced liquefaction because the soils are granular

and have low relative densities.

Excessive Di pi acementss Excessive settlements could occur as a result of the weight of the fill, oil or gas withdrawal, or

thawing of relict permafrost. Damage caused by settlements would

be to structures and well casings. Settlements could also lower

the top elevation of the island and increase its vul nerabi 1 i ty to

truncation failures, wave topping and ice runup. Excessive

lateral displacements could also occur since the frozen upper

portion of the island could act like a rigid body. Mitigation of

displacement effects would be difficult, however flexibility

could be provided to accomodate reasonable amounts of displacement. Horizontal displacements may be very difficult to predict.

Erosions Erosion could be caused by wave and current action, ice

gouging and strudel or other scour. Slope protection was dis

cussed in Section 4.3.2 and was shown in figure 33. Erosion is a servi ceabi 1 i ty limit state, but it could precipitate stability

failures if unchecked. Erosion, particularly during the summer

season, is likely to require periodic maintenance work. Experience with 4 cubic yard sandbags at Seal Island C 50 1 indicates

that erosion protection is expensive, and the bags tend to be

damaged by ice. Other forms of protection, such as concrete armor units (i.e. D0L0S) have been suggested but not tried. In dredq.'d

islands, large sacrificial berms are sometimes used instead of

slope protection.

Ice Runups Ice runup is not strictly a foundation limit <•<'• . Structural type protection normally provided is illustrated in

84
•figure 33. Another method of protection may be construction of

sacrificial berms, which cause formation of rubble fields away

from the island. Such grounded rubble piles may also act to

reduce ice forces, but this effect is offset by an increase of

the width of the area exposed to ice forces C25]. If unchecked,

ice could completely overrun an artificial island.

Wave Overtopping: Wave overtopping, and also spray may be a problem. Present designs seem to accept a 57. overtopping rate for

exploration facilities and a 27. rate for production facilities.

These criteria would require 25 to 30 ft freeboards (from data

presented in 11481).

4.4.6 Design Limit States for Caisson Retained Islands

Caisson retained islands are normally placed on a sand or gravel

berm, but they could conceivably be placed directly on the ocean

floor if conditions are suitable. They have a hydr aul i cal 1

y

placed sand core which is partially dewatered to increase the

effective stresses at the base of the caisson. Their stability

depends to a large degree on the strength properties of the

supporting sand berm and the enclosed sand core.

Seven categories of limit states are identified:

failure; local stability failure; structural

caissons; excessive displacements; erosion;

excessive construction loads. Some of these

dependent on the strength of the supporting berm.

Major stability

failure of the

ice runup; and

limit states are

Stability Failures: Figure 45 shows potential stability failure

modes for Tarsuit Island, identified in Ref.C27D. Other possible

failure modes for similar structures would be a bearing capacity

failure under the caisson, possibly aided by seepage pressures

caused by the partial dewatering of the core, and a deeper

seated sliding surface if the foundation soil under the berm is

85
Surface Loads

FAILURE MECHANISMS

1. Passive failure, single caisson

2. Decapitation

3. Beaded failure, entire structure

4. Active failure, single caisson

5. Rotational failure, single caisson

6. Beaded failure, single caisson

Figure 45. Failure Nodes Identified for Tarsiut Island (taifn

from Ref . [27]

86
weak. The accuracy of an assessment of the safety margins against

these failure modes would depend on the accuracy with which the

shear strength of the supporting soils can be estimated. The

shear strength of the berm and the core depends to a large extent

on the grain size and method of placement of these materials. The

purpose of the berm subcut shown in figure 45 was to interface with stronger materials, and also to provide a shear key for

the berm. This is predicated on the assumption that the berm is

stronger than the supporting deposits. Some additional failure

modes were identified in Ref.CISD and are shown in figure

46. These include active slope failures behind the caisson

precipitated by tensile forces associated with an ice breakout,

and structural caisson failures associated with the uneven

distribution of ice forces acting on the caissons.

Ring type caisson structures would tend to act differently from

segmented caissons. Active or passive stability failures at the

caisson would have to be associated with an overall structural

failure of the caisson, and truncation failures could occur even

when the sand core is not frozen. These ring type structures

could also experience a major bearing capacity failure, aided by

seepage forces or precipitated by erosion of the berm. As in the

case of artificial islands, effects of cyclic loading in the

shear strength along slip planes needs to be considered.

Liquefaction is another failure mode which could occur during

seismic events. The consequence of liquefaction would be spreading of the supporting berm and a buoyancy type bearing failure

of the caisson, causing major settlement and possibly rotational

tilting. Dredged or bottom dumped sand berms would be extremely

vulnerable to liquefaction, unless the construction procedure

includes compaction or compaction grouting.

Structural Failures A structural failure would only be a foundation limit state if the structure relies on the support derived

87
(b) Local poulve failure

Ice pr

(e) Tensile Ice breakout load

with active toll pressure

(d)Rotatlonol Instability in plan

Figure 46. Failure modes for caisson retained islands identified

in Ref . C 18]

.

88
from the sand core or the supporting foundation soil, particularly when the pressure on the structure is non uniform. The amount

of support derived from the soil, in turn, would depend on the

method of placement and compaction.

Excessive Displacements: Potential causes of cumulative vertical

and lateral displacements were already discussed. Cumulative

lateral di spl acements would be more likely in the case of ring

type caisson structures, which can translate as a rigid body.

Erosions In general CRI's are designed so that the caissons are

placed in the wave action zone. However the shelf of the sand

berm could be subjected to gouging by grounded ice, associated

with rubble piles which are likely to form in the vicinity of the

island. Scour effects at the base of the caissons and on the sand

berm slope may also be significant. The type and effects of

erosion likely to occur would depend on the depth of the berm

below mean sea level.

Ice Runup: Ice runup is more likely to occur when the sides of

the caisson structure are sloped. The consequence could be

overtopping of the caissons.

Excessive Construction Loads: Two loading conditions may be

critical: lateral loads and uplift on the caissons associated

with the placement of the hydraulic core; and pressures on the

underside of the caissons during installation (up to 1500 kPa

483 ) .

4.4.7 Design Limit States for Rigid Mat Foundations

Included in rigid mat foundations are those for the floating

caissons and for monopod structures. Even though these mat

foundations are called rigid, attempts have actually been made to

build in some structural "flexibility" (or ductility) by using

89
steel rather than concrete -for the mat L34D. The -foundation could

be placed on a prepared berm, or directly on the ocean -floor with

a minimum o-f preparation. In some instances it is planned to

place the structures directly on the ocean -floor and subsequently

back-fill any voids with sand C 33 D . In North Sea gravity structures concrete grout, rather than sand has been used to fill

these voids. In the Arctic it is preferred to have sand fill, so

that it will be possible at some future time to fill voids

created by vertical settlements. Mat foundations usually incorporate skirts to enhance load capacity and facilitate underfilling

or grouting and dowels for accurate positioning. They can be kept

vertical by differential ballasting, and the space below the mat

is vented to permit. control of porewater pressures. The underside of mats can be level or can consist of dome shaped surfaces

to improve structural efficiency.

Many of the structures with rigid mat foundations can be repositioned. This capability not only enables the operator to

avoid extreme and rare loading conditions, but perhaps more

importantly permits correction in positioning. This may enable

the operator to accomodate major lateral displacements. The rigid

foundation also permits remedial actions for settlements, by

filling the voids created under the mat.

The load capacity of mat foundations can be augmented by ballasting, grout injection, spud piles and skirts. Short skirts (about

7m long) can also be used to ensure full development of the soil

shear resistance near the soil-foundation interface.

Six categories of limit states are identified: Major stability

failure; structural failure of the mats; failure of skirts dur ing

penetration; local failure at skirts by piping; excessive

displacement or tilt; and scour. The failure modes associated

with these limit states depend to a large degree on the at u t • . of the structure. Some of the critical conditions occur during

construct i on . 90
Stability Failures; Different types of local and global stability

failures, identified for North Sea gravity structures C451 are

shown in figure 47. For Arctic structures, the ratio of horizontal to vertical loads is much larger. Thus sliding failures

are more likely to occur than are bearing capacity failures. However, for structures with sloped sides, the ice forces have a

substantial vertical load component and bearing capacity failures

are possible. Similarly, when the water depth is great the overturning moment is substantial and can lead to rotational

sliding. The shear resistance along these potential sliding

surfaces depends on many parameters and will be discussed in the

next section. The presence of ice gouges can reduce sliding

resistance because of incomplete contact between the structure

and the soil and non-uniformity in the density and shear strength

of the upper soil layer.

Structural Failure of the Mat: While structural failure of the

mat is a structural design limit state which is outside the scope

of this report , it is precipitated by excessive soil pressures.

The critical load could occur during construe t i on , after set-down

and before underfilling or grouting. However, critical contact

stresses may also develop at the ridges of ice gouges. The effect

of ice gouges in sandy soils (which produce the most critical

contact stresses) is illustrated in figure 48. Figure 49 shows

some measured data on contact pressures on domed foundation

sections for very dense sands in the North Sea C51D.

Failure of Skirts During Penetration: Skirt penetration resistance has been considerable in North Sea structures and could

conceivably cause a structural failure if the structural design

is inadequate. Thus it is important to determine the maximum

possible skirt resistance that the soil could develop.

Local failure at skirts by piping: While the structure is being

91
SHALLOW BEARING

CAPACITY FAILURE

/ /////// . / / / / / /.//./ 7K

/

LOCAL FAILURE ALONG SKIRTS

LATERAL FAILURE MODES

I

VERTICAL FAILURE MODES

Figure 47. Failure Modes of Gravity Structures (From ReF.C451)

92
CONTACT

STRESS

(KSF)

Fi gure.

Base slab

ZZZZZZZ2ZZ2ZZZZZZZ

; ^width «

| ; Sand ridge

—

Q 1 , = — Y BN

v

A 2

Y

48. Contact Pressures at the Ridges of Ice Gouges.

93
PENETRATION

DEPTH,

Figure 49. Contact pressures on Mat Foundations in the North S«a

(•from Re-f . L 5 1 D >

94
lowered and the skirts are penetrating some distance into the

soil, high hydrostatic pressures develop under the foundation.

This could cause a blowout around the skirts which could erode

the area around the skirt and also transport considerable

quantities of loose soil under the mat. Such failures occurred in

the North Sea and prompted designers to provide for venting or

pressure regulation of the space enclosed by the skirts. Such

pressure regulation has also been used to control the vertical

position of the platform during installation [51].

Excessive Displacements: Excessive static displacements or

rotations can occur by the cumulative effect of ice loads, but

also as a result of dislocation caused by a single ice impact.

The determination of these displacements requires a realistic

estimate of the number, magnitude and di rect i onal i ty of significant load cycles as a function of time, where a significant load

cycle is associated with a load large enough to cause a permanent residual displacement. This assessment is difficult to

make. For sandy soils, the threshold shear strain for a permanent

displacement is about 1 0~z7. . Excessive dynamic deflections could

conceivably occur in deep water as a result of wave and wind

forces, however it is unlikely that these forces would have a

significant effect on structures which are designed for the much

larger ice forces. Tolerable lateral di spl acements and settlements must be defined in terms of the tolerance of the structural

and mechanical systems. These systems, in turn, should be

designed to permit reasonable amounts of displacements, or to

permit adjustments for displacements. For instance Andersen et

al . C50] determined vertical settlements which would cause

yielding of the well casings to be of the order of 100 - 150mm.

Displacements under cyclic loads, rather than global stability

failures could under certain conditions govern the design.

95
Scour: Scour could occur at the edges of mat foundations or on

the slopes of sand berms. In general scour is not likely to be a problem in great water depths, but it could occur as a result of

cur rents.

4.4.8 Design Limit States for Sand and Gravel Berms.

Sand or gravel berms are either dredged or dumped from barges.

Dredged material can be fine (silty sand) but the angle of repose

would be very flat. Dumped material has to be courser, as noted

in Section 4.4.5. The angle of repose of dredged material placed

under water is 1:15 to 1:20, however the material can be placed

between bonds (dikes prepared by other placing methods) which

have a steeper slope angle C47].

Sand berms or gravel pads that are constructed at a location

prior to the placement of an Arctic gravity structure have been

developed and applied in the Canadian Beaufort Sea. They have

been found economically attractive for the following reasons:

1. A relatively long open water season (about 100 days)

2. An abundance of nearby good borrow material (coarse

grained sands).

3. Several high capacity dredge systems are mobilized,

capable of dredging, tr anspor t i ng , and placing fill.

4. Proposed exploratory drilling sites could be determined

one or two years in advance of drilling.

5. Soil conditions are conductive to placing berms.

6. Berms have the advantage of acting as ice defense beaches

that can ground large ice features and develop protective

rubble pads around the platform.

7. Berms were a part of the evolution of offshore systems

from gravel islands to caisson retained islands.

In the U.S. part of the Beaufort this site preparation has

limited applications for the following reasons:

1. A short open water season (generally less than 30 days).

96
2. Good nearby borrow material is not abundant.

3. No high capacity dredge systems are mobilised, and their

availability is restricted by legislation.

4. Development leases are such that at this stage it is

unusual that an exploratory drilling site can be determined -far enough in advance to allow construction of a

berm (over several working seasons) without significant

disruption of exploratory drilling. Berms at several

candidate sites would be too expensive.

The limit states for berms were discussed in previous sections.

However, the question arises, whether it is possible to density

the hydraul i cal 1 y placed material, which has a very low relative

density, and therefore is vulnerable to liquefaction and excess

pore water pressure buildup by even small cyclic loads, and

generally develops low shear strengths. Soil improvement methods

will be discussed in Section 4.5, but at this point it should be

noted that they are both time consuming and expensive. Some

options for soil compaction are discussed below.

Silts and clays can be generally compacted by surcharge, provided that their consol i dat i on time is reasonably short. However

sands and gravels can not be significantly densified by surcharging. Thus some mechanical compaction method must be used.

Three methods which are reasonably successful on shore are

vi br of 1 ot at i on , use of explosive charges, and dynamic compaction. Vi brof 1 otat i on is quite expensive and is generally performed to a depth of + 23m. Equipment for working under water has

apparently not been developed. Blasting seems reasonably effective, but may be difficult and environmental ly objectionable

under water. A comparison of blasting and vi br of 1 ot at i on in an

offshore project is made in Ref.C533. Dynamic compaction seems

more promising and appropriate. The latter method has been

successfully used offshore [54,55,563. It seems that it is

necessary to provide a rock blanket of about 1-1. 5m thickness on

97
top of the fill. Drop height of the weight under water is 10m,

effective compaction was achieved to a depth of about 5-6m.

Weights used varied from 13 to 40 metric tons. The weights should

be fluted to minimise water resistance. In situ monitoring to

determine the density achieved, such as cone, dilatometer, or

standard penetration tests would be necessary. Whether the

logistics of transporting and lifting the heavy weights, building

a rock blanket on top of the fill, and performing the in situ

tests would justify the potential benefits derived from the

compaction has not been determined. Gravel islands could conceivably be effectively compacted from a curface above the water

line, which would eliminate the need for a rock blanket, increase

the efficiency of the operation, and permit conventional in situ

moni tor i ng

4.5 Resistance to Loads and Displacements

4.5.1 General

It has been noted in Sections 2 and 3 that there is considerable

variation and uncertainty in subsurface and loading conditions.

Specifically, ice loads will vary from location to location, in

addition, there are smaller cyclic or variable loads caused by

wind forces on the ice canopy , multiple ice impacts or wave

forces which may have a significant effect on the foundation

resistance and displacements and which at this point in time

cannot be very accurately characterized. Soils are also quite

variable, in part because they were and still are partially

frozen. Lenses or layers of frozen soil will not consol i date

together with the adjacent unfrozen soil, and after they thaw

they form pockets of soft soil or dicontinuities in soil profiles. Presence of past permafrost and many cycles ot freezinq

and thawing probably also account for the irregular overconsol

- dation pattern shown in figure 17, which makes it difficult to

accurately characterize strength properties at a shallow depth.

98
Analytical procedures and models by which foundation resistance associated with the previously discussed stability limit

states can be calculated are generally available. However

considerable uncertainty is associated with the prediction of

cumulative residual di spl acements associated with smaller load

cycles and with the measurement of the soil strength and stiffness parameters used to define the constitutive laws associated

with analytical models.

4.5.2 Characterization of Soils

Soil properties that generally need to be defined are; shear

strength (static, cyclic, residual, steady state); compressibility; permeability (coefficient of consolidation); maximum past

pressure (overconsol i dat i on ratio); stiffness (shear modulus);

liquefaction potential; sensitivity (in clays); grain size

distribution; Atterberg linits; natural moisture content; void

ratio; and density. In addition we need information on soil

fabric, anisotropy, fissures, and trapped gas and permafrost

lenses which may cause us to misinterpret laboratory test

resul ts

.

To measure these soil properties we have two types of methods at

our disposals in situ measurements and laboratory testing. In

situ measurements have the advantage of measuring soil properties

in place, without the disturbances caused by sampling and release

of confining pressures, and without the scale effect associated

with small laboratory samples. On the other hand, most of the in

situ tests also disturb the soil, and our in situ testing

capabilities are very limited since we cannot controll the

boundary conditions which dictate the state of stress and

fail ure mode

.

In Situ Soil Exploration: methods most commonly used off shore

are: geophysical exploration; static cone penetration (CF'T)

99
tests; vane shear tests; pr essur emeter tests; and gamma logging

tests. Other teste which could yield useful results are standard

penetration (SF'T) tests, dilatometer tests, self boring pressuremeter tests, and plate loading tests.

Geophysical tests generally provide information on the velocity

of compression and shear wave propagation and on soil stratification. They also permit us to locate pockets of trapped gas and

relict permafrost. The compression wave propagation velocity in

saturated soil is generally controlled by the elastic properties

of the pore water and therefore provides little information on

the soil skeleton. The shear wave propagation velocity, however,

provides information on the shear modulus associated with very

small strains (of the order of ID-7’), Gm*« , which in turn has

been correlated with shear moduli at larger strains for various

soils (for instance C57II). Thus it can be said that geophysical

measurements provide information on the stiffness of the undisturbed soil deposits. This information could be particularly

useful in the deter mi nat i on of the susceptibility of granular

soils to excess pore water pressure buidup under cyclic loading.

Cone penetrometer (CF'T) tests expand a cylindrical cavity in the

soil as the penetrometer advances. Tip and side resistance are

measured separately. The test gives excellent information on the

strati f i cat i on of the deposits. Even though the penetration

resistance, particularly in granular soils, is a complex function

of strength and deformation properties, very useful information

for preliminary design purposes can be derived. Application o< the test in North Sea foundation practice is summarized in

Ref.C583. Undrained shear strength of clays can be reasonably

estimated, particularly if the plasticity index and the over consol i dat i on ratio (OCR) are known. The angle of shearing

resistance of sands can also be reasonably estimated if ths OCR

is known. Flowever , the latter parameter may be difficult to

determine, unless intervening clay layers are present

which the OCR can be determined in odometer tests. Schmer tmai n

100
[59] also developed correlations -for determining the relative

density of sands. The CRT does not retrieve a soil sample which

is a major drawback. However, an indication o-f the grain size o-f the material can be derived -from the "-friction Ratio" (the ratio

between the side and tip resistance). CRT test results can also

be used to assess 1 i que-f act i on potential [60], however a reasonably reliable assessment would require information on DBo. This

information can then be used to estimate the the SPT blowcount

which would be obtained. The SRT blowcount, in turn, has been

correlated with liquefaction potential on the basis of extensive

data [61]. CRT-SRT correlations are shown in figure 50. Figure 51

shows correlations between SPT blowcount and liquefaction

potential. CRT tests have also been used very successfully to

predict soil resistance to skirt penetration and pile capacity.

Prediction of skirt penetration resistance is illustrated in

figure 52 from Ref. [58]. Similarly, resistance to spud pile

penetration could be predicted.

Vane shear tests give useful information on the undrained sheafstrength of clays and cohesive soils that were in an undrained

mode during shear. Extensive experience has been gained with this

test

.

Pressure meter tests have been extensively used and correlations

with soil strength and stiffness were established. These tests

provide more reliable information on in situ confining pressures

than other in situ tests. However confining pressures in sandy

soils are difficult to obtain because of the sensitivity of the

measurements to disturbances. In the latter case, self boring

type pressure meters would have a better chance of yielding

sucessful measurements.

Gamma logging tests are used in offshore exploration to log soil

strat i f i cat i on through casings. However attemps to use the soil

emission for density logging were not very successful.

101
q

e

. bars ; N , blows/foot ( I bar lOOkPa ) CLAYEY SILTS SANDY SILT

0.001 0.01 0.1 1.0 MEAN GRAIN SiZE,D50 ,mm

Figure 50. Data on CPT-SPT Correlations (-from Re-f.C603)

102
0.6

• Data on Correlations Between SF'T Blow Count and

Liquefaction Potential of Sands and Silty Sands
q c , AVERAGE CONE TIP RESISTANCE AND STANDARD DEVIATION SHEAR STRENGTH APPLIED TO SKIRT PENETRATION ANALYSIS PROPOSED METHOD

MEASURED AVERAGE PENETRATION PROPOSED OF ANALYSIS PENETRATION FORCE

Figure 52. Predicted and Back Calculated Skirt Resi stance

re-f . [48D ) .

( from

104
Standard penetration tests are well known and need not be

discussed. Except in the case of assessing liquefaction potential, the CRT yields more reliable and repeatable results than

the SPT. However the SPT will retrieve disturbed soil samples

which yield much valuable information. The SPT is generally more

reliable in granular soils.

Dilatometer measurements are of relatively recent origin but they

show promise in measurements related to soil stiffness and in

situ confining pressures. The test is much more inexpensive than

the pressuremeter test and is lately extensively used in the

monitoring of dynamic compaction.

Plate loading tests have been used in the North Sea to a minor

extent, primarily to predict contact pressures. The plate used

had a 300mm diameter, thus requiring a substantial downward

thrust. Another plate loading test that could be used is the

screw plate test. Large plate test may be of interest, since it

has been determined as a result of a great number of tests with a 865mm diameter plate, that calculations based on laboratory test

results for clays tend to under est i mate stiffness (settlements)

and overestimate strength C623. While these conclusions may not

apply to the particular soils investigated, insitu plate loading

tests could help calibrate information obtained from laboratory

or other in situ tests.

Laboratory tests: Laboratory tests permit us to apply controlled

boundary conditions and separate the many variables involved in

soil strength and stiffness by changing one parameter at a

time. However they also have severe drawbacks, these include:

1. Scale effects. Not only do we have to upscale the resuls by

several orders of magnitude as we go from a laboratory sized

sample to the full scale problem, but there is also a size effect

on strength and stiffness, particularly when the soils are

f i ssured

.

105
2. Sample disturbance. It is difficult, and in many instances

impossible, to take undisturbed samples. Not only is there a distortion from the center to the periphery of the sample, but we

also remove the confining pressures in the process of sampling

(and as we pointed out, these confining pressures are very

difficult, and frequently impossible, to determine). Undisturbed

samples of cohesionless soils are very difficult to obtain, even

though some freezing techniques were recently successfully

appl i ed

.

3. Boundary conditions. While we attempt to create a determinate

state of stress in the soil volume under test, there are stress

and pore water pressure gradients within the sample. The problem

is particularly severe in the simple shear test, where stress

gradients develop at the end plates.

4. Anisotropy and direction of principal stresses. Soils are

generally consolidated under gravity load and thus are not

isotropic. While in the simple shear test the directions of the

principal stresses resemble those in the field, this is normally

not the case in the triaxial test. Soil strength may change as we

rotate the axis of the principal stresses.

5. Load history simulation. Soil strength is stress history

dependent. However it is virtually impossible to exactly simulate

the actual loading sequence anticipated in the field.

Sampling: The difficulties in obtaining undisturbed samples have

been already discussed. In the offshore, samples are obtained by

driving or jacking. Drive samples are generally disturbed. Frequently samples are also disturbed by internal gas pressures, in

order to insure the quality of offshore samples it is desirable

to use X ray techniques to examine the samples while the explora

tion is in progress.

Determi nat i on of shear strength: Generally the undrained shear

strength has to be determined, however the point in time after

the application of the full vertical loads when the cnt it »1

106
shear stresses are expected to occur may be an important consideration when the tests are conducted. Generally five different

types of test are used: unconfined compression tests; Undrained,

consolidated triaxiar shear tests, which can be compression tests

(the major principal stress is increased while the minor principal stress Ccofining pressure! is held constant), extension tests

(the major principal stress is held constant while the confining

pressure is decreased), or torsional shear tests (the vertical

and horizontal stresses are held constant while torsion is

applied); and undrained consolidated direct simple shear tests.

Unconfined compression tests are generally performed for clays

and can give a good indication of undrained shear strength if

good undisturbed samples are available. Figure 53 which is taken

from Ref.CSl! shows the location on a hypotetical slip plane for

which tri axial compression, direct simple shear, and tri axial

extension tests are applicable. The torsional triaxial test

could be used instead of the direct simple shear (DSS) because of

the similarity between the state of stress and the direction of

c^pplied stresses. For failures generated by a large horizontal

load and a relatively moderate vertical load the state of stress

would be most closely approximated by direct simple shear or

torsional triaxial tests.

The effect of the deformation modes represented by the above

discussed tests on the shear strength is illustrated by figure

54, which is taken from Ref. [63! and was obtained for Santa

Barbara silt, a soil somewhat similar to the Beaufort silt. Note

that triaxial compression produces the largest shear strength,

triaxial extension the small es, and simple shear falls in

between

.

Another important parameter that must be considered in these

tests is the effect of the consol i dat i on pressure applied in a

drained condition before the undrained shear test is performed.

This consolidation pressure corresponds to the prevailing

107
Figure 53. Deformation Modes associated With Shear

(taken from Ref.CSID).

ail ure

108
DEPTH

BELOW

MUDLINE

(ft)

UNDRAINED SHEAR STRENGTH (KSF)

Figure 54. E-ffect o-f De-formation Modes on

Re-f . C633)

Shear Strength (-from

109
confining (or overburden) pressure in the field, and the additional pressure exerted by the structure. Wether or not this

additional pressure is applied in the test will depend on whether

the structure had time to consolidate before the critical

horizontal load occurred. The effect of this parameter is

illustrated in figure 55 from Ref.L63D.

A third parameter affecting shear strength is the OCR, or the

apparent maximum past pressure. The magnitude of this effect in

Santa Barbara silts is shown in figure 56 (also from Ref.C633).

This effect is of interest because the Beaufort offshore deposits

are heavily overconsol i dated at a shallow depth. Since the

pattern of overconsol i dat i on is rather irregular, and overconsolidated soils frequently have a tendency to form fissures

and loose some of their shear strength when horizontal confining

pressures are reduced [64], great care must be exercized when the

shear strength of these soils is evaluated.

Effect of Cyclic Loading on Shear Strength: When relatively small

cyclic loads are applied to soils in an undrained condition (or

within a time frame where pressure dissipation by drainage does

not have a significant effect) they will cause a buildup of

excess porewater pressures. This phenomenon will occur , even if

the soils dilate under large shear strains. Somewhat of an

explanation for this porewater pressure buildup can be derived

from an examination of figure 15. Even soil A, which is dilative

under large axial strain exhibits an initial increase of excess

porewater pressure (u) . As excess porewater pressures increase,

effective stresses, and thus shear strength, decrease. In

relatively pervious soils, the buildup of porewater pressures is

frequently checked by drainage and the pressures dissipate soon after the cyclic load ceases. In soils of low per meab i 1

i

t y

,

however, the excess porewater pressures may affect the she \r

strength for a long time after the cyclic loads were applied.

An extreme example of the potential effect of cyclic loads on the

1 10
DEPTH

BELOW

MUDLINE

(ft)

Figure

I I

10

20

30

40

t =4 months

Drainage barrier

73*357^

1 - initial strength profile (simple shear,

a'vm =4 KSF)

2 - One-way one dimensional-consolidation,

t = 4 months

3 - Vertical drains, t = 4 months (spacing = 15 ft, drain diameter = 2.4 in)

1.0 2.0

UNDRAINED SHEAR STRENGTH (KSF)

55. Effect of

Structure on

Consol i dat i on

Shear Strength

Under the Weight

(from Ref . [63] )

.

of the

1

1

1
DEPTH

BELOW

MUDLINE

(ft)

Figure 56 Effect o-f Apparent Over consol i dat 1 on on Shear Strength

Profiles (from Ref.C63H)

1 12
static shear strength is shown in -figure 57, from Ref.[49]» The

data in the figure were obtained from large diameter plate

tests. This effect is more severe in overconsol i dated soils than

in normally consolidated soils. The effect is very significant in

North Sea structures, where it is generally assumed that the

maximum wave is going to occurr at the end of a 6 hour storm, and

the 6 hour storm is then simulated by a cyclic loading sequence,

whose effect. on the failure load is evaluated. In the Beaufort

Sea the effect of cyclic loads is probably less pronounced , since

the cyclic loads are much smaller in relation to the maximum load

than in the North Sea loading. However we cannot afford to

diregard this parameter , particularly for soils of low permeability.

Another case that, needs consideration is cyclic shear strength

(the number of cycles to failure under a given stress (or strain)

level). A typical set of data for North Sea clays is shown in

figure 58 from Ref.C493. It can be seen from the figure that

over consol i dated soils are more sensitive to strength reduction

by cyclic loading than normally consolidated soils. This undrained cyclic load effect should not be confused with the case

where there is sufficient time for the excess pore water pressures to dissipate. In the latter case, prior load cycles are

likely to increase, rather than decrease the shear strength.

Determination of Compressibility; Compressibility can be determined by odometer tests or drained tri axial tests, which do not

need to be further discussed. However, in the Beaufort Sea, a

careful evaluation needs to be made whether settlements could be

affected by enclosed permafrost or entrapped gases. Predictions

based on odometer tests are reasonably reliable with respect to

magnitude of settlements, however the estimates of the time

element invoved in consolidation settlements is usually not very

reliable. Most soils are or tot r op ic with respect to their

permeability because of the manner of their deposition and pre
Figure 57. Effect of Cyclic Loads on Monotonic Shear Strength in

Undrained Condition (from Ref.C513)

114
T hey

T hf

NUMBER OF CYCLES TO FAILURE

(YC y~± 10%)

Figure 58. Undraind Cyclic Shear Strength o-f North Sea Clays

(•from Re-f .[513).

1 15
leading, and permeability in the horizontal direction tends to be

much greater than that in the vertical direction. This -fact leads

to a tendency to overestimate the settlement time, which is an

error on the conservative side. Analytical methods -for considering anisotropy and the geometry of boundary conditions are

available (for instance [65]). As previously noted, it is

important to estimate the consolidation time because of its

effect on the shear strength of the supporting soil.

Liquefaction, Cyclic mobility and Steady State Strength: Liquefaction is a matter of concern when dealing with sands and

silts. However as previously noted, cyclic strength may be a

matter of concern with most soils. Seismic loads are the primary

concern , but dynamic loads associated with multiple ice impacts

may also have to be considered. As previously noted, excess

porewater pressures will buid up in most saturated soils subjected to undrained cyclic loading. When the cyclic loads continue

for a sufficient number of cycles and are repeated rapidly

enough the effective confining pressure will approach zero, and

soils with no cohesive strength will experience initial liquefaction. In that stage soils lose their shear strength and act

like heavy liquids. Whether or not this condition can lead to

major stability failures depends on whether the soils will dilate

when subjected to further shear deformation. In general , str uc tures on loose sands or silty sands will experience catastrophic

failures when initial liquefaction occurs, while structures on

dense soils will experience limited displacements. Some indica

tion of the effect of liquefaction on sands of various retd i ve

densities can be derived from figure 59 which shows shake table

data presented by Seed [66]. It can be seen that sands with

relative densities below 407. liquify and cause catastrophic

failures and sands with relative densities of less than 6u7. are

subject to large shear deformations. Whether or not initial

liquefaction is likely to occur as a result of a given cv> li<

load can be determined from a set of cyclic triaxial test s slu h

1 16
Figure 59. Shake Table Data on Cyclic Mobility (taken

Ref . [66] )

.

117

from
as those shown in figure 60 which is taken from Ref. [673. The

trouble is that it is very difficult to obtain undisturbed

samples of granular soil, and there is no assurance that reconstituted laboratory samples are r epr esent at i ve of field conditions. Thus at the present time empirical curves such as the one

shown in figure 51 are most frequently used to estimate susceptibility to 1 i quef act i on

.

There are two concepts which should receive consideration when

evaluating susceptibility to liquefaction: The threshold strain,

and the steady state strength. The threshold strain concept is

based on data obtaind for a variety of sands, but has not yet

been extended to silts. It is based on the hypothesis that there

must be intergranular sliding if the volumetric strain which

causes pore water pressure buidup is to occur. It has been

observed exper i mental 1 y and shown theor et

i

cal 1 y that below the

threshold strain of 10~s 7. (0.0001) no porewater pressure buildup

in quarts sands is possible. Thus when it can be shown that this

strain will not be exceeded there is no need to worry about

liquefaction. Whether or not this strain is likely to be exceeded

can be determied from an analysis using shear modulus values

derived on the basis of geophysical measurements. The concept can

probably also be extrapolated to predict porewater pressure

buidup above the threshold strain, but not enough research has

been done to accomplish this end. Figure 61 shows threshold

strain data C683. Figure 62 shows trends of porewater pressure

buidup above the threshold strain from experimental data C673.

The threshold strain increases with an increase in the OCR C68a3.

Following the hypothesis that porewater pressure buildup l primaril y related to the shear modulus, some attempts have been

made to correlate shear wave velocity measurements with lique

faction potential C693. The results of this work are shown i r» figure 63. Unfortunately there are not enough data at thi - time. The use of geophysical measurements for liquefaction

1 18
CYCLIC

SHEAR

STRESS.

Kg/cm 2 MEDIUM MONTEREY SAND

INITIAL VOID RATIO, ej = 0.68 (Dr = 50%)

IMITIAI CCCCTTU/r PCI M CIM IM P DDCCC 1 1 D C \

<7 V

= 2.0 kg per

uuivnmivu rntoo

sq cm

unt,

\

0

1 10 100 1,000 10,000

NUMBER OF CYCLES TO CAUSE FAILURE, n

Figure 60. Stress Controlled Cyclic Triaxial Tests o-f Reconstituted Sand Sample (taken from Ref.C67D).

119
RESIDUAL

PORE

PRESSURE

RATIO,

Au

,/a

' 3

SHEAR STRAIN,7, PERCENT

l-igure 61. Observed Threshold Strain (taken from Ref.C6Sl)

120
PORE

PRESSURE

RATIO,

Au

/o

' 3

u — — — 1——-———— —r

o Crystal silica sand <73 = 2000 psf, fresh samples, dry vibration

1.0 a Crystal silica sand <73 = 2000 psf, fresh samples, wet redding - a Crystal silica sand a 3

= 2000 psf, staged testing, dry vibration 0

0.8 Sand no. 1 (73 = 2800 psf, staged testing, moist tamping * V Sand no. 1 0 3 = 1400 psf

,

staged testing, moist tamping

Monterey no. 0 sand 0

\

= 2000 psf, fresh samples, moist tamping

0.6 0/ A -

/a

0.4 &) -

0.2 n = 10 cycles D r = 60%

10-3 10 2 10 1 SHEAR STRAIN, 7, PERCENT

Figure 62. Correlations Between Cyclic Strain and Pore Water

Pressure Buildup (from Re-f.C67II).

121
Shear

Wave

Velocity

of

Liquefiable

Sand

Layer

(fps)

F i gure 63 . Correlation Between Shearwave Velocity and Liquefaction Potential of Sands (from Ref.C69D).

122
eveluation is an attractive option, particularly in the preliminary stages of a project.

The steady state strength concept C70] is based on the observation that saturated soils, subjected to large shear deformations,

will reach a stage of ideal platic flow, where further shear

deformation is not associated with any volume change. In this

stage the effective confining pressures remain constant. The

steady state strength is a unique function of density and grain

size characteristics, and independent of soil fabric or initial

confining pressures. A typical steady state strength curve is

shown in figure 64. The steady state strength could be approached

from the left side of the curve in figure 64 by volume expansion

if the material is dense and the initial confining pressures are

relatively small, or from the right side of the curve by volume

contraction if the material is loose and the initial confining

pressures are relatively high. If the shear strength required to

support the structure is smaller than the steady state strength,

a catastrophic failure cannot develop. Conversely, if the

required shear strength is higher than the steady state strength

a catastrophic failure will result from initial liquefaction.

The steady state strength approach is attractive because the

parameters of the problem, steady state strength and in situ

density, can be evaluated. The steady state strength line can be

determined in the laboratory from disturbed samples (even

though the authors recommend to also test undisturbed samples).

In situ density must be measured very accurately. The reason for

this requirement is that the steady state line, particularly for

soils with subrounded or rounded grains, is rather flat (a small

change in density is associated with a large change in steady

state strength). The steady state strength approach is also an

effective tool for evaluating the resistance of a foundation

after large shear displacements occurred and therefore could

be used to estimate displacements caused by large ice features.

123
VOID

RATIO.

MAXIMUM SHEAR STRESS AT THE STEADY STATE . q t , Isf ( I Isf S 100 k Pa )

Figure 64. Steady State Strength (taken -from Ref. £703)

124

ORYUNIT

WEIGHT,

y

,

pcf(tpcf*16

Vq/m'
Some analytical models where developed which are designed to

predict pore water pressure buildup and cyclic mobility. One such

model developed by Arulanandan et al.C713 is correlated with

electrical in situ measurements. Whether these measurements can

bu used ott shore has not been established.

Field and laboratory tests associated with the previously

dicussed methods of, liquefaction evaluation would be SF’T and CRT

tests tor empirical approach, stress controlled triaxial tests ot

reconst i tuted samples tor the traditional approach, and shear

wave velocity measurements tor the threshold strain approach. For

the steady state strength approach very accurate in situ density

measurements are needed (piston samples are recommended). To

establish the steady state line tor the material, shear strains

have to be very large. This can be in some instances accomplished

by triaxial tests with special provisions tor sample expansion

near the end platens. Larger strains can be achieved in torsional

triaxil tests and in simple shear tests. For clays laboratory

vane shear tests may have to be pertormed.

The State Parameter: The concept ot the steady state strength has

been used more recently in conjunction with the stability problem

associated with the placement ot hydraulic tills tor sand berms

(dredged or bottom dumped). Large excess hydrostatic pressures

develop in these tills C723, leading sometimes to stability

tai lures. The concept ot the steady state strength is ideally

suited to deal with this problem. It the sand exhibits contractive behavior when subjected to shear stress, a shear tai lure

can be precipitated as excess porewater pressures increase. Conversely, it the sand expands under shear stress the till will be

stable. Expansive sands will be to the lett ot the steady state

line in tigure 64, while contractive sands will be to the right

ot the line. The position ot the in situ material with respect to

the steady state line is detined by two parameters: the void

ratio, and the in situ state ot stress (contining pressures)

.

125
A single parameter has been proposed by Been et al . C731 to

define this in situ state of the sand, the "state parameter".

The state parameter, vp

, was defined by the authors as the

difference between the void ratio of the in situ sand and the

void ratio the same sand, subjected to the same confining

pressures, expressed in terms of the first stress invariant, 1^

=

C (

a

j. + a 2 + 'CT.3 >/ 33, would have to have in a steady state of

deformation. This definition is illustrated in figure 65.

tp would be positive if the in situ sand is contractive, and

negative if it is expansive. Thus a sand with a negative state

parameter would be inherently stable, and a sand with a positive

parameter inherently unstable, regardless of its relative

density. Therefore the state parameter describes the in situ

state of the sand in a more fundamental way than the relative

densi ty

.

In reference E73D the state parameter is correlated with several

sand characteristics. An example is shown in figure 66. As

previously noted, the utility of the concept of steady state

strength goes beyond the liquefaction problem. In terms of

measurement of soil properties, definition of the state parameter would require deter mi nat i on of in situ confining pressure.

It is suggested that another parameter, namely the ratio of the

in situ confining pressure to the confining pressure at the sam*

>

void ratio at the steady state of deformation would be very

useful for design purposes.

Work now in progress [721 includes study of the use of th> CF1

test in the definition of the state parameter. A proposal at the

end of this report to develop an instrument which can combine the

CF'T measurement with a measurement of the in situ confining

pressure could solve this problem.

126
Void

ratio

Fi gur Definition of the State Parameter (from Ref. [7

127
State parameter +

Figure 66. Correlation Between the State Parameter and the Angl»’

of Shearing Resistance (from Ref.C73J).

128
Determination of stiffness parameter©? Soil stiffness is generally affected by several parameters. These includes density;

con-fining pressure; magnitude of strain; rate of strain; and

OCR, and there is a considerable body of knowledge by which

various in situ measurements can be correlated with shear

modulus. In the field, geophysical measurements can be used to

determine Gm«,< as previously dicussed. In the laboratory,

triaxial, as well as compressive and torsional resonant column

tests can be used, provided reasonably undisturbed samples

are available. Stiffness data are important for mathematical

modeling of static and dynamic response to loads.

Use of Index Properties: In many instances, when there is already

substantial information on the soil deposit, index properties can

be effectively used to predict strength and stiffness parameters. For instance values of su /p ' , where su is the undrained

shear strength and p' is the effective overburden pressure have

been correlated with the plasticity index of normally consolidated clays, and and the s^/p ' value for normally consolidated

clays, in turn, can be adjusted for the overconsol i dat i on

ratio. At least this latter adjustment may have to be used even

when the shear strength is determined from laboratory tests

because of the erratic variation of overconsol i dat i on ratios in

shallow Beaufort offshore deposits (the test values may have to

reduced to correspond to some reasonable OCR value). The use of

index properties may become very effective when calibrated

against a body of available data, and may be used to detect test

results which substantially deviate from the mean values of

available data.

4.5.3 Augmentation of Soil Resistance

As noted in Section 4.4.2, there are practical limitations to the

load resistance of gravity mat foundations. The load capacity can

sometimes be increased by modi f i cat i ons of the foundation or the

129
supporting soil. It should however be realized, that the plan

dimension of these foundations is very large, and that strengthening of the soil to a moderate depth may sometimes merely change

the location of the slip plane without a substantial change in

the load resistance. There are three types of methods by which

the load resistance can be augmented:

(1) Compaction or consol i dat i on to increse density.

(2) Grouting, admixture or freezing, to change the material

character! sties.

(3) Addition of load resisting elements such as spud piles and

skirts.

Increase of Soil Density: Cohesive soils are can generally be

densified by surcharging and consolidation. The time element

required for the consolidation settlement may be a problem if

critical loads could occur soon after installation of the

structure. Consolidation settlements can be increased by vertical

sand or wick drains. Granular soils generally can not be substantially compacted by surcharging. Mechanical means of compacting

sands and gravels were discussed in Section 4.4.8.

Compaction Grouting, admixtures and freezing: These methods will

generally increase the soil strength by at least an order of

magnitude. They tend to be expensive, but they may be effective

when a relatively shallow layer of weak soil is underlain by much

stronger soil. The feasibility and type of grouting which can be

effectively used depends on the soil type and permeabi

1

i ty . A

newly developed mixing procedure C74] has also been used success

fully for seabed strengthening and is being tried in the Arctic.

In that latter procedure, cement is mixed in place with in situ

soils. An interesting possibility would be the use of a grouted

soil 1 ayar instead of a pre fabricated mat. This possibility has

apparently not yet been explored.

1 30
Soil -freezing is also being explored. This method could rely on

a passive system, using heat exchange with the environment, an

active system which would require energy input, or a combination

of both. On a much smaller scale, -freezing has been successfully

used in the -foundations o-f the Trans Alaska Pipeline.

4.6 Reliability o-f Predictions

Any analytical model can only be as accurate as the parameters

used to characterize the soil. There is presently no organized

body o-f in-formation to assess the accuracy o-f our in situ and

laboratory tests. Thus this estimate must be le-ft to the judgment

o-f the engineer. In general, the confidence in our estimates can

be much improved when the strength and stiffness parameters are

determined by one method and then verified by other methods. The

topic of gravity foundation reliability has been addressed in a

recent presentation [75].

The use of partial safety factors, like those in European

practice, may be reasonable, since it permits the engineer to

assign a numerical value to the partial safety factor for

resistance which corresponds to his confidence in the results of

in situ and laboratory measurements. In conjunction with this

approach, a target value could be assigned to the safety index

(the argument in equation 4.1). A value of 2.3 would not be

unreasonabl e , when used in conjunction with the 100 and 25 year

mean recurrence intervals for design loads, presently used for

production, and exploration facilities, respectively.

131
-
SUMMARY

Subsurface Conditions:

Subsurface conditions in the Alaskan Beaufort and Eastern Chukchi

areas pose complex and unique problems. These include the

presence of relict permafrost, shallow gas deposits, ice gouges,

and erratic patterns of shear strength and overconsol i dati on

.

There are some areas of instability and some seismically active

areas. Sand and gravel resources are relatively scarce.

Surface deposits are pr edomi nant 1 y over consol i dated silty clays

and clayey silts of Holocene origin. These are underlain by

Pleistocene deposits of sands, silts and clays which are frequently fully or partially ice bonded. Pockets of soft soils, and

irregular overconsol i dat i on patterns are common.

Loads:

Ice loads develop large horizontal thrusts. Design wind speeds

with a 50 year mean recurrence interval are between 90 and

100 miles per hour. Maximum wave heights are about 10m and storm

surges rise to about 3m above mean sea level. Currents and

strudel scours can cause erosion problems. Seismicity is limited

to a small area in the vicinity of Barter Island.

Ice loads are are much greater than any other envi r onmental loads

and controll the design of foundations. However, potential

effects of other en vi r onmental loads should be taken into

account. Rare and exceptional ice loads (e.g. due to ice islands)

are of a magnitude that cannot be resisted by feasible gravity

structures on mat foundations. These loads must be avoided, or

their effects must be mitigated by means other than load resistance. Global ice lods used for foundation design increase with

water depth and depend on the exposed width of the structure and
the slope of its sides.

Foundation Types:

Foundation types considered in this report are foundations for

sand and gravel islands, foundations for caisson retained islands

which are rigid circumferential structures with a sand core,

resting on sand or gravel berms, and rigid mat foundations for

gravity structures which are usually placed on the ocean floor

with relatively little preparation, but could be placed on

prepared berms. The type of structures and foundation used is to

a large extent dictated by contruction feasibility during the

short open water season of the U.S. Beaufort Sea (about 30 days).

Design Limit States:

When failure probabi 1 i t i es , associated with design limit states

are considered, account must be taken of the fact that rare and

exceptional ice loads, which can not be realistically resisted by

gravity structures, have probab i 1 i t i es of occurrence which may

not be acceptable for property damage or extensive environmental

damage, and would probably also not be acceptable for the total

loss of the installation. Thus strategies must be considered to

avoid such ice loads or mitigate their consequences

.

Design limit states are identified for the various types of

foundations. Whether these limit states (which are failure modes)

should be defined as ultimate or sevi ceab i 1 i t y limit states

should not be determined in accordance with rigid design rules,

since this deter mi nat i on depends on the consequences of the

failure mode associated with the limit state. Many offshore

structures have the capability to mitigate the consequences oi

certain types of limit states, such as excessive displacements or local instability.

134
Load resistance:

While analytical models to determine foundation response to loads

are generally available, the accuracy of the prediction is only

as good as our ability to determine the relevant soil and loading

parameters. Accurate characteri sat i on of soils is difficult

because of the many parameters which influence the strength and

stiffness of soils in a given state of stress. These parameters

are discussed in this report. In particular, it is difficult to

determine in situ geostatic stresses and to obtain undisturbed

samples of granular soils for laboratory testing.

Prediction of cummulative di spl acements and degradation in soil

strength and stiffness resulting from seasonal cyclic ice loads

are problems that warrant significant attention.

Various methods of soil strengthening, including consolidation,

mechanical compaction, grouting, admixtures and freezing are

briefly discussed. It is noted thet compacion to a shallow depth

may not significantly improve load capacity because of the large

dimension of the foundation. The

mat foundation by a layer of

grouting or mixing with cement i s of interest.

possibility of replacing rigid

soil which is strengthened by

135

6. NEEDED STUDIES AND RESEARCH

(1)

. Our capability of measuring soil properties in situ should

be improved, with particular emphasis on in situ confining pressures, strength and stiffness. (2)

. It is recommended to conduct combined analytical and

experimental studies to improve our ability to predict horizontal

displacements under sustained ice loads and cycles of seasonal

ice loads.

(3)

. Generic information should be generated on the density and

strength character i st i cs of hydr aul i cal 1 y placed granular

material used in berms, as a function of placement method and

grain size characteristics.

(4)

. Density and strength char acter i st i cs of sand underfill

placed after seating of the structure should be studied.

(3). More generic information should be generated on the geotechnical properties and character i st i cs of Beaufort shelf deposits.

This information should include cyclic and steady state strength,

effect of cyclic loads on static strength, effects of overconsolidation, and correlations between in situ and laboratory data.

1 37

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D.S. GOVERNHEMT PRIKT1NC OFFICE : 1987- 1 « 1 -0 7 . / f. 0 : I :
NBS-114A (REV. 2-80

U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See instructions)

1. PUBLICATION OR

REPORT NO.

NBSIR 86-3419

Performing Organ. Report NoJ 3.

4. TITLE AND SUBTITLE

Mat Foundations for Offshore Structures in Arctic Regions

Publication Date

FEBRUARY 1987

5. AUTHOR(S)

Felix Y. Yokel, Robert G. Bea

6. PERFORMING ORGANIZATION (If joint or other than NBS, see instructions) 7. Contract/Grant No. national bureau of standards

DEPARTMENT OF COMMERCE 8. Type of Report & Period Covered

WASHINGTON, D.C. 20234

9.

SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street, City, State, ZIP)

10.

SUPPLEMENTARY NOTES

| | Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11.

ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here)

Design limit states for offshore gravity structures in the Alaskan Beaufort and

Eastern Chukchi continental shelves are discussed. The report contains a description of geological conditions, design loads, and type of structures used.

Three foundation types are considered: foundations for artificial islands;

foundations for caisson retained island with sand cor^s; and rigid foundations

for various types of gravity structures which are positioned on the ocean floor

with a minimum of preparation. Design limit states for these foundations are identified and the required reliability against ;the occurrence of these limit states

is discussed. Our ability to determine foundation resistance is assessed.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolons)

Artificial islands; geotechnical engineering; ice forces; mat foundations; ocean

engineering; offshore platforms; oil production; sand and gravel berms; soil exploration; soil testing.

13. AVAILABILITY 14. NO. OF

j_)(J Unlimited

PRINTED PAGES

| | For Official Distribution, Do Not Release to NTIS

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~1 Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C.

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15. Price

[XJ Order From National Technical Information Service (NTIS), Springfield,' VA. 22161 $18.95

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