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2
ASSESSMENT OF CONDUCTOR SETTING DEPTN
OTC 671:
3.
The packer is inflated to seal the test sec-
tion, and a wireline dart is lowered to the bit
to measure pressure during the test.
4.
The test section is
pressurised by pumping
fluid into the drillstring at a given flow rate
and the test performed as outlined on Fig. 2.
The test is therefore generally performed in a
pre-drilled section, and is flow controlled. The
measurement made during the test include the fol–
lowing:
1.
The initial fracture pressure.
2.
The steady state pressure
rate.
3. The close up pressure after
4.
The re-fracture pressure.
under a given flow
initial fracture.
Whilst all stages of the HFT
test may be used
to infer geotechnical soil parameters, this paper
concerns
the pressure required to cause
initial
fracture, which is generally adopted as the limit of
allowable
excess fluid pressure during drilling
operations.
Initial fracture pressures from such
tests have been used to check the theoretical method
presented in the following.
THEORETICAL BACKGROUND
Minor Principal Stress Approach
The hydrostatic,
drilling
fluid and in situ
soil pressures during drilling for the first casing
string
are shown
schematically on Fig. 3.
The
drilling fluid pressure which may be expected to
fracture the soil formation has been the subject of
analysis by Bjerrum et al.
(1972).
This approach
was developed following obaervationa of fracture
occurring
whilst
installing push-in piezometers.
The excess pressure, Au, required to cause a ver-
tical crack in the soil as derived by Bierrum et al.
is given by:
Au =
or Au =
where v =
CY=
Po’ =
k=
o
.-
Po’(l/v-l) [(l-~)ko + Pt’/po’] ... (1)
po’(l-v) [(2+&c@ko + pt’/pol] ... (2)
Poisson’s Ratio of the soil
effect of installation on circumferen-
tial stress
effect of installation on radial stress
vertical effective stress in-situ
coefficient of lateral earth pressure
at
tensi
est
e stress sustainable by the soil
Equation 1 represents the case of fracture
occurring prior to blow off, and equation 2 repre-
sents fracture occurring following blow off of the
soil from the piezometer.
This assumption of a “perfect” installation and
a Poisson’s ratio of 0.5 (undrained response) re-
duces equations 1 and 2, respectively, to:
Au = ko.p ‘ + pt’
o
... (3)
or Au = ko.p ‘ + pt’/2
... (4)
o
Bjerrum also notes that there is a possibility
of a horizontal crack forming if the excess head
exceeds p ‘.
In his recommendations on allowable
pressures,”derived from theory and field and labora-
tory observations,
the tensile stress ptt was con-
servatively ignored.
These results then reduce to
the assumption adopted by many oil companiea in
estimating required conductor setting depth, that an
excess pressure equal to the lower of the principal
stresses in the ground should be assumed to cause
hydrofracture, i.e.:
Au=p’
o
(ko>l)
... (5)
Au = ko.p ‘
o
(ko
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ALDRIDGE AND HAIMD
?
tial stress falls to zero or to any given value of
tensile stress,
pt’, as follows:
Au = 2,ko.p ‘ + uh + ptt
o
... (7)
where
‘h
= hydrostatic pressure at the given depth
Equation 7 is consistent with the results given
by Jaeger (1969) for a porous elastic medium, if the
permeability of the medium is set to zero.
It may
therefore be expected that hydrofracture will occur
at the preseure given by equation 7; unless a gen-
eral shear failure of the clay at the wall of the
borehole occurs at a lower pressure.
Examination of
the three principal stresses given on Fig. 4 may be
used to calculate the maximum deviator streea in the
clay
material.
If the maximum deviator stress
exceeds twice the undrained shear strength of the
clay, then a plastic failure of the borehole wall
may be expected to occur.
The deviator stresses
derived from the vertical (v), radial (r) and cir-
cumferential (c) principal etresses are as followe:
or-u
= Au- p ,
v
o
.,. (8)
Ur - IJC
= 2.Au -
ko.po’
... (9)
Uv-u
‘Au+p ’-2,ko.p ’
c
o
0
... l o
Shear
failure will occur when any of these
deviator stresses exceeds twice the undrained com-
preaaive ahear strength of the soil.
Using equa-
tions 81 to 10 it is therefore possible to derive
eqUatiOnS 11 to 13,
respe tively
defining the
eXCeSS
fluid pressure which would cause a shear
failure in the borehole wall:
Au = 2.s
u + Po’ ... (11)
Au= S + k
u
O.PO’ . . . (12)
Au = 2.au + po’(2ko-1)
... (13)
where s
= undrained ehear strength in compression
u
An alternative possible mode of failure to that
considered above is a uniform cavity expansion,
for
which the excees pressure to cause failure is of the
order of
5.7 to 6.3 times
the undrained shear
strength, depending on the overconsolidation ratio
of the clay (Randolph et al. (1979)).
It ia poaaible that a ahear failure, as given
by the lowest of equations 11 to 13, will occur
prior to the tensile failure given by equation 7.
There are assumptions inherent in both theoretical
approaches, however. Observations made during
drilling for the first casing string are therefore
reviewed below to assess the validity of each ap-
proach.
The results of these approaches are also
compared
with
the method traditionally used, as
given by equations 5 and 6.
method.
RESULTS OF FIELD TESTS
A review has been made of the reeults of 34
hydraulic fracture tests ‘(HFT’s) performed in pre-
drilled sections in geotechnical boreholes performed
during platform
site-investigations in
the North
Sea,
The teste were performed at aix sites, at
depths of between 40 and 140 metres below mudline,
in hard clay strata. Of the 34 tests, three resul-
ted in very high fracture pressures close to
those
expected from cavity expaneion theory, as previously
reported by Overy and Dean (1986).
Three resulted
in anomalously low pressures, believed to have been
due to leakage around the packer.
Data from the
remaining 29 tests are reviewed below.
The predicted and meaaured test results are
compared for the six test sites on Fig. 5.
The
dashed line representa the calculated minor princi-
pal stress and the solid line is the lowest of the
preseures
derived from equations 11 to 13.
The
results chow graphically that the “shear failure
approach gives a closer fit to the HFT test data
than the traditional “minor principal stress method
at these sites.
The ratio of measured to calculated fracture
pressure has been plotted for the traditional ap-
proach, i.e.
equations 5 and 6,
on Fig, 6, and for
the pressures given by the lowest of equations 11 to
13 on Fig.
7, for all 29 sites.
The valuee given by
equation 7 are not plotted, since they are alwaYs
higher than the values given by equations 11 to 13,
and the
“shear failure”
mechanism may therefore be
assumed to control.
It may again be seen from Figs.
6 and 7 that the “shear failure ’approach repre-
sented by equations 11 to 13 gives a significantly
better overall fit to the data than the “minor
principal stress” method.
Observations of drilling mud pressures and
returns or lack of returns during offshore drilling
operations are not generally available to the geotech-
nical consultant.
Records obtained during drilling
from a semi-submersible drilling rig at the Draugen
site in the Norwegian Sea are presented on Fig. 8.
This figure shows the mud preaeure actually applied
during drilling and the estimated fracture pressure
based on the
“minor principal stresstt
and
“shear
failure” methods.
Again this data confirms that the
“ahear failure”
approach gives
results which are
more coneiatent with observations,
since although
excess drilling fluid pressures exceeded those given
by the “minor principal stress” approach, no IOS5 of
returns waa encountered.
The results of a statistical analysis of the
data are also presented on Figa. 6 and 7, and show
that the measured test pressures
are on
average
almost
exactly double those given by the “minor
principal strees”
method but only 34 per cent higher
than given by the “shear failure” method.
The
statistical correlation,
as measured by the standard
deviation,
is also better for the ‘rehearfailure”
169
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)TC 6713
A.LDRII GENI
10
2.
3.
4.
5.
6.
Bjerrum, L.,
Nash, J.K.T.L., Kennard, R.M.
and Gibson, R.E. (1972), “HydraulicFractur-
ing in Field Permeability Testing , Geotech-
nique Vol. 22, No. 2, pp. 319-322.
Den Hartog, J.P. (1952), “AdvancedStrength
of Materialsit,McGraw-Hill.
Jaeger, J.C. (1969), “Elasticity, Fracture
and Flow”, Halsted Press.
Randolph, M.F., Carter, J.P. and Wroth,
C.P. (1979),
llDri~enpilee in Clay - ‘he
Effects of
Installation and
Subsequent
Consolidation”,
Geotechnique Vol. 29, No.
4, pp. 361-393.
Overy, R.F. and Dean, A.R. (1986), “Hydrau-
lic Fracture Testing of Cohesive soil”.
Proc.
Offshore
Technolozv
Conference.
Paper No. OTC 5226.
Poulos, H.G.
and
“Elastic
Solutions
Mechanicst .
Series
John Wiley and Sons.
-.
Davis, E.H.
(1974),
for
Soil and
Rock
in Soil Engineering,
.
—SIGNAL CABLE
—~ F T DART
_SLllJING VALVE
FOR PACKER
—PRESSURE DROP VALVE
—PRESSURE SENSOR
—ROUGH HOLE PACKER
-OPEN BIT
iil
,.
Fig.1 HydraulicFractureTestEquipment
I
I
1
I I
I
I I
I I
I I
I
,,
TJUF
II
RF
STEAOY STATE
PREsSUBE
1
I
1’
CLOSE UP
PRESSURE
800
I ‘i
k
I
/ “
4
600
I
\
? w P
‘
\ <
400
T
200
PUMP OFF -
+
PUMP O J
b +
PUMP OFF
o
0
4
a
12
16
20
24
28
32
36
40
TIME AFTER START OF TEST (rein)
Fig. 2 Hydraulic Fracture Test Procedure
171
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.
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TOTAL STRESSESIN-SITU
koPo’+ Uh
TOTAL STRESSES DURING ORILL ING
UhdJ
Fig.3 Hydrostatic,Mud-umand
InsituSoilStresses
Fig.4 Changes inTotalSoilStresses
DuetoDriUingOperations
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EXCESS FRACTURE PRESSURE (MPal
40
60
80
100
120
140
0
1
z
‘Y\
\
1
\
\
\
o
i
2
50
I 1
I
70
I
I
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‘*
I
90
1 ,.
1
x
110
0
i
2
00
\
\
100
\
\
\
120
\
w
\
\
D
140
EXCESS FRACTURE PRESSURE (MPa)
o
2
40
60
80
100
120
140
1-
1
I
I
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)
1
I /
I
I
1
I
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m
60
80
100
120
0
2 4
\
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t
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A
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\
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\
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I
\
120
\
\ o
.
(B
140
HFT RESULTS
‘-’-- MINOR PRINCIPAL STRESS PREDICTION
— SHEAR FAILURE PREDICTION
Fig.5 Comparison ofHFI’Reaults withPredictions
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40
60
100
120
140
RATIO MEASURED/CALCULATED
o
1 2 3 4
5 6
-1 SD MEAN
+1 SD
I
I
[
I MEAN =
1.99
I
I
I
S.D. =
4
0.52
I
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1
b
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1: ~
10
I
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1
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1°
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q
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Fig. 6 Ratio of Measured to Calculated Fracture Presaure-
Minor Principal Stress Method
RATIO MEASURED/CALCULATED
o i
2
3
4
5 6
I
-1
SD MEAN
+1 SD
I
I
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1
I
L
I
MEAN =
1.34
-1
I
S.D. = 0.32
60
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Fig. 7 Ratio of Measuxwl to Calculated Fracture Pres.sure-
Shear-Failure Method
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