1.Chapter 6_SPE-116633-PA_An Integrated Geomechanical and Passive Sand Control Approach

8/18/2019 1.Chapter 6_SPE-116633-PA_An Integrated Geomechanical and Passive Sand Control Approach

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z z x y r w xy A P P= − −( ) − − −2 2 4 2cos ( ) sin ,

. . . . . . . . . . . . . . . . . . . . . . . . (4)

z xz yz= − +( )2 sin cos , . . . . . . . . . . . . . . . . . . . . . . . .(5)

x H h v= +( ) +cos sin cos sin2 2 2 2 , . . . . . . . . . . (6)

y H h= +

sin cos2 2 , . . . . . . . . . . . . . . . . . . . . . . . . . . (7)

z H h v= +( ) +cos sin sin cos2 2 2 2 , . . . . . . . . . . (8)

yz H h= − −( )0 5 2. sin sin , . . . . . . . . . . . . . . . . . . . . . (9)

xz H h v

= + −( )0 5 22 2. cos sin sin , . . . . . . . . . . . . (10)

xy H h= − −( )0 5 2. sin cos . . . . . . . . . . . . . . . . . . . . . (11)

In Eqs. 6 through 11, v, H , and h are the vertical, maximumhorizontal, and minimum horizontal in-situ stresses, respectively; is the hole inclination with respect to the vertical; and is theangle between the hole direction (projection on plane) and H direction (counterclockwise positive from H direction). In Eqs. 3and 4, Pr is the far-field reservoir pressure, and A is a poroelasticconstant given by:

A =−−

( )

( )

1 2

1

, . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (12)

in which is Poisson’s ratio and α is Biot’s constant. Biot’s con-stant is defined as follows:

= −1 C C

r b

, . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (13)

where C r and C b are the grain and bulk compressibilities, respec-tively.

The borehole direction (well azimuth, W az, in degrees) and theminimum-horizontal-stress direction ( haz in degrees) are definedclockwise positive from the north direction. The well deviation, , (in degrees) then can be related as

= = = + −0 0 90,if ,otherwise; haz azW . . . . . . . . . . . . (14)

The effective formation strength, U, in Eq. 1 is not measureddirectly in the laboratory. In relation to sanding study, the so-calledthick-wall-cylinder (TWC) strength is measured in the laboratory.The standard dimensions for TWC samples used are usually 1½-in.outside diameter (OD) × ½-in. inside diameter (ID) × 3-in. length.The standard OD/ID ratio = 3 for samples does not replicate thehole collapse under the in-situ condition where the OD/ID ratiotends to be infinity. By varying the sample OD/ID ratio in a seriesof TWC tests, it has been established that in-situ TWC strength isapproximately 1.55 times the laboratory measured TWC strengthwith the standard sample dimensions (Willson et al. 2002). Thus,theoretically, the effective formation strength, U, could be equatedto 1.55 × TWC strength. However, because of the conservativenature of the linear elastic theory on which the preceding equa-tions are based, the satisfaction of Eq. 1 with a value of U = 1.55 × TWC strength does not necessarily cause instant sand failure of asignificant volume caused by residual strength of the failed rockand arching effects after initial rock failure. This effect is usu-

ally taken into account pragmatically by boosting the formationstrength further, as follows:

U b TWC f = × ×1 55. , . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (15)

where b f is the boost factor and TWC is TWC strength. The value ofthe boost factor should be adjusted to match sand-failure evidencein a well under certain drawdown conditions during productiontests or actual production in a field. In the absence of any suchdata, a conservative default value of 2 is often used for b f , whichis usually the source for U = 3.1 × TWC strength in many refer-ences for cased-and-perforated wells (Willson et al. 2002). For anopenhole completion, a slightly more conservative boost factor of1.6 is usually recommended.

For openhole completions, Eq. 1 is directly applicable tocalculate LF (i.e., max e / U ) for a given drawdown or BHFP (Pw)at a certain depth using the well’s inclination and azimuth. Forcased-and-perforated completions, the inclination of perforationsfrom a deviated, or a vertical, well can be defined by the followingequation as a function of well trajectory ( , ):

perf perf

= −90 cos . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (16)

For a horizontal well ( = 90),

perf perf = , . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (17)

in which perf is the perforation orientation in degrees from the top ofthe wellbore in deviated wells. The value of perf is positive clockwiseand negative counterclockwise from the top of the well. For a devi-ated well, the perforation azimuth can be calculated as follows:

perf az perf W = + 90sin . . . . . . . . . . . . . . . . . . . . . . . . . . . . (18)

For a vertical well ( = 0), the perforation azimuth is defined as

perf haz perf

= + . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (19)

For a horizontal well ( = 90), the perforation azimuth is definedas

perf perf

az perf

az

W

W

= == − <

= +

0 0

90 0

90

;

;

;

if

if

if f perf > 0

, . . . . . . . . . . . . (20)

where perf is the perforation azimuth in degrees. Having the perfo-ration azimuth defined by using the appropriate equations amongEqs. 18 through 20, the other perforation-trajectory parameter, perf , can be calculated as follows:

perf perf

ha

= ==0 0if ;

z z perf + −90 ;otherwise

. . . . . . . . . . . . . (21)

Sand failure at a particular perforation can be assessed using Eq.1 by replacing and with perf and perf , respectively. The criticalBHFP (CBHFP) or critical drawdown (CDD) (CDD = Pr – CBHFP)can be calculated by numerically solving Eq. 1 with LF = 1.

The effect of reservoir-pressure decline caused by productioncan be accounted for in the preceding computation by updating thein-situ stresses. For a laterally large reservoir compared to its thick-ness, the change in vertical stress, v , is considered negligible, andtherefore it is usually kept constant. The maximum and minimumhorizontal stresses are updated as follows, respectively:

H H H P= + × , . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (22)


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h h H P= + × . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (23)

In Eqs. 22 and 23,

H =

−−

1 2

1. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (24)

and

P P Pc i= − , . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (25)

where Pc is the current reservoir pressure and Pi is the initial res-ervoir pressure. While using updated H and h stresses in Eqs. 1through 11, it is also important to use current reservoir pressurePc in place of Pr . In Eq. 24, H

is defined as the stress-path factorexpressing the change of horizontal stresses with reservoir-pressurevariations in a passive and normal-faulting stress regime. Different

equations have been suggested for the determination of stress-pathfactors in other stress regimes and tectonically active areas (Addis1997). Nevertheless, published literature and the worldwide dataset indicate a range of 0.5–0.9 for the stress-path factor.

Sanding-Evaluation Workflow

The final application of Eq. 1 to assess sand failure requires anextensive geomechanical workflow, presented in Fig. 1. The centralcomponent is the geomechanical model that consists of the magni-tudes of in-situ stresses, v, H , and h; direction of H or h stressand magnitude of pore pressure, P p; and rock mechanical propertiesincluding rock strengths and elastic moduli. Zoback et al. (2003)have provided a comprehensive methodology for determination ofmagnitudes and orientation of stresses in hydrocarbon reservoirs.

The magnitude of vertical stress is most preferably calculatedby integrating the density log ( b) over the depth. In the absence ofa direct density log, a pseudodensity log can be created by usingother logs such as sonic or VSP. If no log is available, lithologicaland regional information can be used to approximate v.

The magnitude of least principal stress ( h in normal and strike/ slip-faulting stress regimes) is usually estimated from leakoff tests(LOTs), extended LOTs (ELOTs), or minifrac tests. ELOTs/LOTsare usually carried out as part of a drilling program. A minifrac testis conducted with the particular objective to measure the minimumhorizontal stress in a field. It is a controlled fracturing of a well atlimited scale with detailed recording of injection-pressure data. Whiledetermining h by interpreting LOT, ELOT, and minifrac data, screen-ing out poor quality data is an important part of quality control. IfLOT, ELOT, or minifrac data are not available, h can be calculated

from the vertical stress by using Poisson’s relationship. This approach,however, is not reliable in tectonically active areas. Use of regional

information and empirical models for the magnitude of least principalstress is also useful where reliable field data are not available.

The orientation of horizontal stresses is best derived fromobservation of borehole breakouts and drilling-induced tensilefractures in high-resolution image logs and/or multiarm caliperlogs (for borehole breakouts). Anisotropy of shear-wave velocitydetermined from fast and slow shear-wave-velocity data measuredby acoustic logs in cross-pole mode can also be used to derivestress orientation. This is possible when the stress-induced veloc-ity anisotropy can be distinguished from other sources of velocityanisotropy, such as anisotropy caused by rock fabric, bedding andlamination, and natural fractures. In the absence of high-resolutionimage- and multiarm-caliper-log data or shear-wave-anisotropydata, other sources of information for stress orientation includerecent tectonic activities in the area of interest and regional stressdata such as the world stress map (Reinecker et al. 2005).

There is no direct way to measure or determine the value of

maximum horizontal stress, H . However, it could be modeled fairlyaccurately using linear elastic theory and rock-failure criteria for agiven borehole condition (e.g., presence or absence of hole break-outs and/or induced tensile fractures observed in high-resolutionimage logs and multiarm caliper logs). A systematic graphical tech-nique known as the “stress polygon” method developed by Moosand Zoback (1990) uses frictional equilibrium theory and borehole-failure observation to estimate the magnitude of H consistent within-situ stresses, rock strength, mud weight, and pore pressure.

Rock-strength parameters, including unconfined compressivestrength, peak cohesion, and friction angle, are best determined fromtriaxial tests on representative rock samples. Laboratory-derivedrock-strength values are then correlated with well logs to derive acontinuous strength profile as a function of depth. Published rock-strength models (Chang et al. 2006; Khaksar et al. 2009) and analogrock-strength data sets are often used for rock-strength profilingwhere limited or no core rock-strength data are available.

Pore pressure is determined from direct measurements in perme-able and reservoir zones (e.g., repeat formation tester and drillstemtest or other well tests). Using the modeled points as a basis, in-situ stress profiles vs. depth are produced by using the effective-stress-ratio principle.

Once the initial geomechanical model is built, the model isvalidated, and revised if necessary, to support drilling experiencesand hole conditions at other depth sections in the study wells andacross the field. The final geomechanical model, the rock-strengthprofile, and initial reservoir-pressure and pressure-decline data(from the production plan) are used for sanding evaluation.

Sanding-evaluation results are expressed in terms of “sand-

free operating envelope” and “sanding-evaluation- log” plots. Thesand-free operating envelopes show a series of curves representing

Drilling and Production Datamud weights/ECD, PWD, survey,

drilling history and events, LOT/XLOT,

RFT/MDT, DST, production info

Geomechanical Modelstress magnitudes and orientation,

pore pressure and rock strength

Core DataRoutine &SCAL

UCS, TWC, PSD, thin section,

SEM, XRD dispersion, chemical

Core Data

,

Geolo. Geophys. and Petrophy.

Seismic, Tectonic history,

sediment., analogs, etc.

Sanding Evaluation

Well LogsCaliper, GR, Rhob, Sonic,

image, dipmeter, MWD/ LWD

Well Logs

Sanding Evaluation Log Perforation Optimization Operating Envelope

Validate the model

with drillingexperience and hole data

Fig. 1—Sanding-evaluation workflow.



specific rock strengths at a given depth, whereas the “sanding-evaluation logs” are depth plots for a production zone with knownrock-strength profile under a specific reservoir-pressure condition.In sand-free operating envelopes the critical BHFPs are plottedfor a series of drawdown and depletion conditions, including thecurrent drawdown and the planned final abandonment pressures.These form a family of curves separating the sand-free zone (nosand) from the sand-prone zone (sanding). Where the BHFP andreservoir pressure for any well flowing condition cross a rock-strength curve, sand failure is predicted and the onset of sandproduction is assumed from rocks with the specified strength.

In the sanding-evaluation log, the CDD depth profile is plot-

ted over the entire production zone for a given reservoir pressure.Where the critical drawdown (Pr – BHFP) is less than the planneddrawdown, sand failure is predicted and the onset of sand produc-tion is assumed. Alternatively, a sanding-evaluation depth plotknown as the load-factor plot can be generated for any pair ofreservoir and drawdown pressures. A load-factor plot comparesthe maximum effective borehole stress acting around a hole orperforation to the effective formation strength U (Eq. 1). If LF isless than 1, then the formation will not fail and sand will not beproduced. However , if LF > 1, the converse is true.

Sanding-evaluation logs and sand-free operating envelopes areproduced for a given reservoir pressure, well trajectory, and perfo-ration orientation or openhole condition. These graphical outputsclearly show the weak (sand-prone) intervals in the reservoir section

and the maximum drawdown that can be achieved in these intervalswithout producing sand with the reservoir pressure at a particular pro-duction stage. This information assists production engineers to makestrategic sand-management decisions, such as minimizing sandingrisk by optimizing well trajectories and perforation orientations,planning for selective perforation from the beginning or shutting offperforations in weak intervals at a certain stage of production, oradapting the best active sand-control measures on the basis of theseverity and the likely timing of sanding in sand-prone intervals.

A Case Study: Application to Gas Field Y

The preceding sanding-evaluation methodology has been appliedto a gas field in the south Asia region. Because of the confiden-tiality agreement, a hypothetical name Y is used for the gas field

in this paper. The field had been on production since 1999, and9 gas producers were in operation at the time of this study (late2006). These producers are: Y-2, -3, -7, -8, -11, -13, -16, -17, and-18, which all have cased-and-perforated completions with no sandcontrol installed. The field production-history data have indicated avery minor amount of sand production (i.e., only approximately 10kg of sand was found in the production separator in late 2005). Thissand might have been produced during the initial beanup and/orcleanup of perforation debris in the wells. The Y-12 well, whichwas completed with an expandable sand screen (ESS) was shutdown in 2006 because of unexpected water production.

In 2003, an initial sanding-evaluation study was conducted forthe field using a limited data set. The initial sanding study indicatedthat there was a moderate risk of sanding at the early stage of fieldlife and recommended a selective-perforation and controlled-beanup

strategy as the means of minimizing sanding risk under no- or limited-pressure-depletion conditions. The field came to the hands of a newoperator in 2004, and additional data were acquired from the fieldby drilling six new wells within the field and adjacent areas. A newfield-development plan came into play, and the field life was plannedto extend to 2029–2030. The initial reservoir pressure within the targetsandstones was approximately 3,100 psi, and the current reservoirpressure (as of 2006) has decreased to 2,400–2,700 psi as a result ofapproximately 7 years of gas production. The final reservoir pressureis expected to drop to 400–700 psi. In light of the new field plan andadditional new well data including image and multiarm caliper logs,ELOTs, core rock-strength data, and 7 years of production history inthe field, an augmented sanding study was conducted to

• Update the geomechanical model (rock strength and in-situ

stresses) in the field for the overburden and reservoir rocks usingavailable data from the most recent wells.

• Build a sand-failure prediction model using the updated geo-mechanical model, verified and calibrated with production data.

• Assess sanding risks for the specific reservoir sandstones inexisting wells over their producing life, with a more detailed reviewwhen the flowing tubinghead pressure (FTHP) drops to 100 psi,for which compression would be required.

• Assess the risk of sand production for planned infill wells tobe drilled in the near future.

Building the Geomechanical Model. First, daily drilling reportsand end-of-well reports for all post-2004 wells were reviewed todocument drilling experiences and to extract relevant information

for building the geomechanical model (see Fig. 2 as an example;there is discussion later on its use).

The vertical stress ( v) was calculated from the integration ofdensity-log data in a well with the longest density-log coverageamong the study wells. The vertical-stress profile indicates a stressgradient of approximately 1.01 psi/ft at 6,850 ft true vertical depthsubsea (TVDSS), which is the depth of the target reservoir sand-stones. The values of least principal stress (or h in vertical hole)were estimated from ELOTs conducted at the casing shoes in studywells. Assessing the availability and quality of reported pressure/vol-ume data in all study wells, 7 ELOTs from three wells were usedto constrain the magnitude of the least principal stress in the field(see Fig. 3 as an example).

The pore-pressure (P p or Pr ) information was obtained from

direct measurements in the reservoir sections. Available image-logdata in two of the study wells were analyzed for stress-inducedborehole breakouts and tensile fractures. The hole condition overthe interval with image-log coverage was mostly in gauge withsporadic overgauge sections in shale and sandstone intervals(Fig. 4). No clear borehole elongation and/or tensile fracturescould be detected from image data in the reservoir and sectionsimmediately above the reservoir in both wells; therefore, theazimuth of H could not be determined from image data in thisstudy. Analysis of four-arm caliper (dipmeter) data in two pre-2004 vertical wells drilled in the nearby area indicates consis-tent borehole elongation in northwest/southeast direction in theoverburden section at depths between 4,000 and 5,500 ft TVD,which indicates a northeast/southwest direction for the azimuth ofmaximum horizontal stress. The stress orientation inferred from

dipmeter data is consistent with the published regional stress data(Reinecker et al. 2005).

Rock-Strength Data and Modeling. Multistage triaxial compres-sive-strength tests were performed on five sandstone samples fromthe Y-11 core, one of which was from the reservoir interval of con-cern. Individual samples were tested at different confining pres-sures: 300, 1,000, 2,050, and 3,100 psi. As a sample is confined,it structurally reinforces the sample so that the axial stress requiredto cause failure increases in accordance with Mohr-Coulomb theory.The multistage data from the five samples were combined to pro-vide a set of composite Mohr-Coulomb parameters for each sam-ple. The unconfined-compressive-strength (UCS) values derivedfrom the multistage tests scatter between approximately 2,300 and3,660 psi, except for one sample from the nontarget reservoir with

a UCS value of 1,754 psi. During inspection of the Y-11 core, itwas found that this sample had been taken from a highly laminatedsection of the reservoir, and therefore, the test results may notrepresent mechanical properties of the entire reservoir sandstonesin the study area.

As discussed earlier, TWC rock strength is necessary forsanding evaluation. TWC tests were conducted on 13 horizontalplugs from the Y-11 well. The measured TWC values rangedbetween 1,450 and 5,300 and between 2,200 and 5,770 psi forinternal failure initiation and external wall catastrophic failure,respectively. Typically, there is a distinction between internal andexternal wall failures. The external wall catastrophic-failure pres-sure corresponds to the perforation-failure condition that couldcause continuous and possibly catastrophic sand production. The

internal-wall-failure pressures normally correspond to the onsetof transient sanding, which is often assumed to be manageable.



Similar to the multistage triaxial test results, the TWC-test resultsindicate intermediate rock strengths that are reasonable, consider-ing the age and depth of the target sandstones in the field.

TWC and multistage triaxial test results along with well-log

data in a cored well were used to derive empirical relationships

for UCS and TWC rock strength for reservoir sandstones in theY field. The results are shown in Figs. 5 and 6, for UCS andTWC relationships, respectively. The UCS values derived frommultistage triaxial tests correlate well with corresponding sonic

travel times ( Dtc) from the sonic log. A similar strong correlation

0

200

400

600

800

1000

1200

1400

1600

1800

2000

2200

2400

2600

2800

3000

3200

3400

3600

3800

4000

7.0 8.0 9.0 10.0 11.0 12.0 13.0 14.0

Mud Weight (ppg)

M e a s u r e d D e p t h [ M D , m ] Mud Weight

Formations

Casing Shoes

Formation

KCL/PHPA

LTSBM

133 /8 in.

95 /8 in.

26 in.

185 /8 in.

7 in.

Seawater with HVIS Sweep

0

200

400

600

800

1000

1200

1400

1600

1800

2000

2200

2400

2600

2800

3000

3200

3400

3600

3800

4000

Drilling and Logging Experiences

D e p t h [ R T D , m ]

Tight HoleReamFormationOverpull/Drag/TorqueLossesPipe/Tool Stuck

133 /8 in.

185 /8 in.

26 in.

7 in.

95 /8 in.

TD 3933 m MD

12900 ft MD

Fig. 2—Mud weight and drilling experiences in Well Y-16.

Fig. 3—Interpretation of ELOT in Y-8 well (13 3 ⁄ 8-in. casing shoe) to determine least principal stress ( 3 or hmin).



between UCS and sonic log data has been reported for other rocks(McNally 1987; Horsrud 2001). The UCS/ Dtc correlation for thetarget sandstones is in the form of

UCS Dtc= × −40 847 0 0268, exp( . ) , . . . . . . . . . . . . . . . . . . . (26)

where UCS is in psi and Dtc is sonic transit time in µs/ft.A very good correlation was also found between the core-

measured TWC strength and log-derived dynamic compressionalmodulus ( M ). The dynamic compressional modulus ( M ) is definedby the following:

M D

bc

tc

= × ×1 34 1010

2.

, . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (27)

where bc is bulk density in g/cm3 from corrected density log. TheTWC / M correlation for the target sandstones is in the form of

TWC M = ×−10 8 1 77. , . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . (28)

where TWC is in psi and M is dynamic compressional modulus alsoin psi. Similar strong correlation between TWC and dynamic com-pressional modulus has been reported for other rocks (McPhee et al.2000).

No laboratory strength tests were conducted on the interres-ervoir and overburden shales. The following empirical equation

developed for tertiary shales in the North Sea by Horsrud (2001)was used for the shale sections in the study area:

GR

Callipers

in-gauge, good quality image in sands

in-gauge, good quality image in sands

in-gauge, good quality image in shales

Over-gauge

hole (washedout), but no

clear borehole

elongation

(breakout) can

be detected

Fig. 4—Example of image-log data in a post-2004 well showing good hole quality. Caliper logs indicate overgauge hole condi-tion at 7,335- and 7,350-ft intervals, but no clear breakout can be detected in image data.

(b)

UCS vs Dtc (sonic log)y = 40847e

–0.0268x

R2 = 0.7565

0

500

1000

1500

2000

2500

3000

3500

4000

4500

85 90 95 100 105

Dtc (us/ft)

U C S ( p s i )

UCS

Expon. (UCS)

(a)

Fig. 5—Core-log calibration and UCS profile in Y field based on core and log data in Y-11 well: (a) UCS calibration and (b) log-derived UCS profile.



UCS Dtc

= ×

111 7304 87

2 93

..

.

. . . . . . . . . . . . . . . . . . . . . . . . . (29)

These empirical core-log rock-strength equations derived fromthe Y-11 well were used to create a continuous rock-strength profile

in the target sands of all study wells. Figs. 5 and 6 display the log-derived UCS and TWC profiles, respectively, for reservoir layersin the Y-11 well. The UCS profile for the target sand interval inY-11 indicates a UCS ranging from 3,050 psi (P10) to 4,500 psi(P95) with an average (P50) of 3,570 psi, indicating intermediaterock strength.

Log-derived TWC rock-strength profiles in target sands ofstudy wells indicate the following:

• P10 TWC-strength range between 3,270 (in Y-2) and 3,870psi (in Y-13)

• P50 TWC-strength range between 3,350 (Y-8) and 4,820 psi(Y-13)

• P90 TWC-strength range between 5,200 (Y-17) and 10,100psi (Y-13)

The magnitude of maximum horizontal stress ( H

) was con-strained by modeling the stress and pressure conditions with therock strengths (UCS) that are consistent with drilling experiencesand wellbore failures (or lack of failure) inferred from image-logand caliper data in the Y-11 well. For example, if a tight spot wasexperienced at a certain depth (Fig. 2) and a breakout could beidentified from caliper and image logs (Fig. 4) at the same depth,then the accurate value of H at this depth would be able to simulatethe hole breakout in conjunction with other stresses, pore pressure,rock strength, and mud weight used during drilling at this depth.

Fig. 7 shows the resultant stress and pressure profile in the Yfield. According to the generated geomechanical model, the field isassociated with a normal-faulting stress regime, h < H < S v. Themagnitude of the stresses and pore pressure in the target reservoirsat approximately 6,850 ft TVDSS are v gradient (psi/ft) ≈ 1.01 ±

0.005; H gradient (psi/ft) ≈ 0.83 ± 0.02; h gradient (psi/ft) ≈ 0.79 ±0.02; and original Pr , or P p , gradient (psi/ft) ≈ 0.45 ± 0.02. The

above stress and pressure model corresponds to preproduction con-ditions. The initial reservoir pressure within the target sandstonereservoirs at 6,850 ft TVDSS was approximately 3,100 psi (i.e., ahydrostatic-pressure gradient). The reservoir pressure at the time of

Log-derivedCore data

7850

7900

7950

8000

8050

8100

8150

8200

8250

0 4000 8000 12000 16000

TWC (psi)

M D ( f t )

(b)

TWC = 1E-08 M

1.77

R = 0.85

0

1,000

2,000

3,000

4,000

5,000

6,000

7,000

2.E+06 3.E+06 3.E+06 4.E+06 4.E+06 5.E+06 5.E+06

Dynamic Compressional Modulus (psi)

C o r e T W C ( p s i )

0

1,000

2,000

3,000

4,000

5,000

6,000

7,000

2.E+06 3.E+06 3.E+06 4.E+06 4.E+06 5.E+06 5.E+06

(a)

Fig. 6—Core-log calibration and TWC profile in Y field based on core and log data in Y-11 well: (a) TWC calibration and (b) log-derived TWC profile.

0

1000

2000

σv

Pp hydrostatic

Pp Y 7

Pp Y 11

Pp Y 16

Pp YN 1

Pp YNE 1

MW Y 7

MW YNE1

Reported FIT/LOT

QC'd LOT-XLOT

σh

σH Y 11

σH

Tops Y 7Casings Y 7

σv

Pp hydrostatic

Pp Y 7

Pp Y 11

Pp Y 16

Pp YN 1

Pp YNE 1

MW Y 7

MW YNE1

Reported FIT/LOT

1000 2000 3000 4000 5000 6000 7000 8000

3000

4000

5000

6000

7000

8000

σv

Pp hydrostatic

Pp Y 7

Pp Y 11

Pp Y 16

Pp YN 1

Pp YNE 1

MW Y 7

MW YNE1

Reported FIT/LOT

QC'd LOT-XLOT

σh

σH Y 11

σH

Tops Y 7Casings Y 7

Pressure and Stress (psi)

Pressure and Stress Profile

T V D ( f t s s )

Fig. 7—Stress and pressure profile for Y field.



this study (end 2006) in the target reservoirs decreased, as a resultof several years of gas production, with the depletion rate varyingbetween 50 to 150 psi/yr.

Sanding Evaluation. The values of characterized geomechanicalparameters (magnitudes and orientations of in-situ stresses, porepressure, and TWC strength) were used as input for sand-produc-tion prediction in the Y field. The operator provided the initialreservoir pressure and the production history and forecasts. Theplanned abandonment reservoir pressure was considered to beapproximately 400–700 psi. The depletion rates were reportedto be approximately 50–150 psi/yr since the start of productionin 1999. A Biot’s constant of 1.0 and Poisson’s ratio of 0.25 areassumed for effective-stress and depletion calculations.

In the Y field, there was no evidence of sanding from theproducing wells since the beginning of production in 1999. Allwells had been completed with standard casing and perforationswithout sand control with the exception of the Y-12 well. Y-12was completed with an openhole ESS in 2004. Following waterbreakthrough in 2006, the well had been shut in. Using the defaultvalues of boost factor (Eq. 15), reservoir pressure, and drawdownconditions that were being used during production in 2006, modelpredictions matched field observations of no sanding in all cased-and-perforated study wells but predicted sand production in Y-12before water breakthrough. However, the sand-production predic-tion for Y-12 well could not be verified by field data because theY-12 well had downhole sand exclusion and hence sand could notbe seen at the surface even if it was produced. It therefore seemedreasonable to continue the sanding modeling over the life of thefield conditions assuming the default values for boost factor as2.0 for cased-and-perforated completions, and 1.6 for openholecompletions which are considered to be conservative.

To limit the size of this paper, sanding evaluations for the Y-11

(vertical), Y-12 (deviated openhole), Y-16 (highly deviated), and Y-7(deviated) wells will be discussed here, whereas the findings for

other wells will be summarized briefly. The procedure for sandingevaluation for each well is as follows:

• Construct a continuous TWC rock-strength profile and TWC-strength percentile curve along the completed section for each well

using the developed TWC / M relation (Eq. 28) derived from theY-11 core-log calibration.

• Verify model prediction against the no-sand observation underthe reservoir-pressure and drawdown conditions as of 2006.

• Establish the best and worst perforation orientations for sandproduction.

• Assess the sanding risk for each well with the best and worstperforation orientations for life of well conditions, from currentreservoir pressure to the designated abandonment pressures, usingsand-free envelope plots.

• Identify the sand-prone zones in the entire perforated sectionfor the worst perforation orientation.

• Generate sanding-evaluation logs for reservoir-and draw-down-pressure conditions when the FTHP drops to 100 psi as aspecial case of interest.

Y-11 Well. This well was drilled in 2004, was completed withstandard perforations, and began gas production in August 2004with an initial reservoir pressure of approximately 2,900 psi. Thewell was producing sand free with a pressure depletion of approxi-mately 250 psi and drawdown pressure of 200 psi as of August2006. The log-derived TWC rock-strength profile and percentilecurve over the perforated zones (Figs. 8a and 8b, respectively)indicate that TWC values range between 3,560 (P10) and 7,050 psi(P95) with an average of 4,730 psi (P50). Fig. 9 shows the produc-tion history and forecasts for the well. The well is nearly verticalat the depth of the sandstone (8,100–8,226 ft measured depth).

Note that perforation optimization is not necessary for thiswell because it is known that the H direction is the worst and h direction is the best for perforations in terms of sand production

from a vertical well subject to a normal-faulting stress regime. Thisis because the perforations along the H direction become subject

8098

8106

8114

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8146

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0 2000 4000 6000 8000 10000

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(a)

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Perforated Zones only

0

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60

80

100

0 2000 4000 6000 8000 10000

P10 TWC ~ 3560 psi

P95 TWC ~ 7050 psi

P50 TWC ~ 4730 psi

Perforated Zones only(b)

Fig. 8—Calibrated log-derived TWC profile using data from Y-11 well (a) TWC-strength profile, (b) TWC-strength percentiles.



to the highest deviatoric stress ( v – h) that is orthogonal to theperforation axis. This causes the maximum principal stress, andhence the highest sanding risk. The reverse is true for perforationsoriented along h. Fig. 10 shows a CDD log plot (sanding-evalu-ation depth profile) with the worst and best perforation directionsunder the reservoir- and drawdown-pressure conditions in August2006 (brown dashed vertical line shows the drawdown). This

figure indicates that the entire perforated interval should haveremained sand free. This prediction is consistent with no-sandingobservation in this well as of August 2006. Also, it is interestingto note the wisdom behind the decision not to perforate the twoweak intervals (shaded blue in the plot). Had these two weakintervals been perforated, sand might have been produced before2006. Fig. 11 shows the modeled CDD pressure in the Y-11 well in October 2016 when the FTHP drops to 100 psi with a reservoirpressure of approximately 1,300 psi (i.e., approximately 1,600-psipressure depletion), which indicates that sanding is expected fromapproximately 40% of the production zone.

Fig. 12 is a sand-free envelope plot for the Y-11 well for boththe best and the worst perforation directions with a series of TWCrock strengths from 3,500 to 5,000 psi. This plot indicates thatthe well should produce completely sand free until the reservoir

pressure drops to approximately 2,400–2,500 psi (i.e., approxi-mately December 2008 through January 2009 according to Y-11production plan; see Fig. 9). At this reservoir pressure, very minorsanding would be expected from perforations along worst direc-tions only because rocks with a TWC strength of approximately3,500 psi or lower (approximately P10) will be sand prone. Fora given rock strength, perforations parallel to h (i.e., northwest/ southeast orientation) would be more stable and allow an additionalapproximately 200-psi pressure depletion before sanding. Approxi-mately 50% of the perforated zone is expected to produce sandat the final reservoir pressure because rocks with TWC strength< 5,000 psi will fail. Sand production is likely to be exacerbatedwith water production (from January 2018 onward). The effect

B l o c k P , F T H P , F B H P , S B H P , Q w t r , Q c o n d

G a s R a t e , M M C F D

4000

3000

2000

1000

0

100

80

60

40

20

0

Jan-99 Jan-09 Jan-19 Jan-29

Y-11 WELL PERFORMANCE

485 MM 1100 psi WHP, 500 psi WHP, and 100 psi WHP

Fig. 9—Production history and forecast for Y-11 well.

M D ( f t )

Critical Drawdown

Fig. 10—CDD log for Y-11 well with reservoir pressure anddepletion in 2006.

Well Sanding Evaluation Log

Fig. 11—CDD log for Y-11 with FTHP = 100 psi condition.

B H P ( p s i )

Reservoir Pressure (psi)

Fig. 12—Sand-free operating envelope for Y-11.



of water production on sanding severity is not considered in thispaper. Water production may exacerbate sand production becauseof chemical reactions between water and sand and changes ofcapillary forces leading to cohesive-strength reduction (Han andDusseault 2002). The effect of capillary-force alteration is usuallyincorporated in the elastoplastic modeling approach (Hawkes andMcLellan, 1996).

Y-12 Well. This well was drilled and brought on production in

2004 from a secondary target above the main sands. It is a highlydeviated well with deviation and azimuth of 74 and 63º, respec-tively, through the reservoir target at 15,400–16,000 ft measureddepth. The well was shut in 2006 because of water production (Fig.13). The values of log-derived TWC strength in the productioninterval vary between 3,070 psi (P10) and 6,730 psi (P95), withan average of 3,960 psi (P50).

To establish the optimum well trajectory for openhole comple-tion in the field, a sanding polar plot showing critical BHFP iscreated (Fig. 14) at 6,800 ft TVDSS with the initial reservoirpressure and TWC strength = 3,070 psi (P10). Y-12 well trajectoryis shown by the green triangle on the polar plot. This plot showsthat a vertical (at the center of polar plot) or a near-vertical wellwould have been the optimum well trajectory for openhole comple-tion from a sanding point of view, giving approximately 1,200-psiCBHFP (i.e., approximately 1,800 psi CDD) at the beginning ofproduction. However, the well may have been drilled with thechosen trajectory for other reasons, and consequently, it is almost

a horizontal well approximately perpendicular to the minimumhorizontal stress. Thus, the well has been subject to the maximumdeviatoric stress and hence almost the maximum sanding risk inthe prevailing stress field. The initial CBHFP for Y-12 well trajec-tory is approximately 3,700 psi. Noting that the initial reservoirpressure is approximately 3,000 psi, no sand-free drawdown wouldexist for the openhole-completed Y-12 well even at the beginningof production. This is also confirmed by the sand-free envelopeplot (Fig. 15) for TWC-strength values between 3,500 and 4,000psi. Sand production could be expected from intervals with rockstrength less than 3,750 psi with 240-psi pressure depletion and a400-psi drawdown (these conditions existed before water break-through in 2006). Therefore, nearly 40% of the production zoneshould have been producing sand at the time of water production.Sand production should have been exacerbated by water produc-tion. Nevertheless, this prediction could not be verified from thefield observation data because this well had downhole sand control(ESS). It is highly likely that though the bottomhole sand failed

from the beginning, it had not been detected at the surface becauseof the presence of an ESS.

Y-16 Well. This well was drilled and gas production started inJuly 2004 with an initial reservoir pressure of approximately 2,900psi. The well was producing sand free with a pressure depletionof approximately 250 psi and drawdown pressure of 200 psi as ofJanuary 2006. The well is deviated 57–59° with a hole azimuth ofapproximately 190° at the target sand depth (12,060– 12,180 ft mea-sured depth). The log-derived TWC rock strength profile over theperforated zones indicated that TWC values range between 3,600(P10) and 5,450 psi (P95), with an average of 4,350 psi (P50).

Fig. 16 shows the effect of perforation orientation on sandingrisk for the Y-16 well trajectory for a rock with TWC strength of3,500 psi at 6,650 ft TVDSS under different reservoir-pressureconditions starting from 2,860 psi. This is indicated by plotting

the CDD pressure vs. perforation angle measured clockwise (posi-tive angle) and counterclockwise (negative angle) from the top ofthe hole. Fig. 16 indicates that all perforation orientations shouldbe stable (sand free) under zero and 500-psi pressure depletion,but at 1,000-psi pressure depletion, perforations with a directionangle between –70° and 60° will remain stable with the planned200-psi drawdown. With further depletion, the risk of sandingwill increase; for example, with 2,000-psi pressure depletion onlyperforations with a –30° and a 10° angle from top will be stable.The worst perforation direction (highest sanding risk) will be theperforations at the side of the well. Fig. 17 is the load factor ( LF )vs. depth plot for perforations along the worst direction under res-ervoir-pressure and drawdown conditions as of January 2006. The LF value less than unity (LF < 1) over the entire production zone

indicates sand-free production. This is consistent with no-sandingobservation in the well as of January 2006.

B l o c k P , F T H P , F B H P , S B H P , Q w t r , Q c o n d

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Y-13 WELL PERFORMANCE

485 MM 1100 psi WHP, 500 psi WHP, and 100 psi WHP

Fig. 13—Production history for Y-12 well.

Fig. 14—Openhole sanding polar plot for Y-12.

Fig. 15—Sand-free operating envelope for openhole Y-12.



Fig. 18 is the sand-free envelope plot for this well with hori-zontal perforations located on the sides of the hole (worst orienta-tions) and for perforations located near to the top and bottom ofthe hole (best orientations) for a series of TWC rock strengthsfrom 3,500 to 5,000 psi. For horizontal perforations located at thewellbore sides, no sanding would be expected until the reservoirpressure drops to approximately 2,200 psi (approximately January2010). At this time, minor sand production may occur from theseside perforations in rocks with a TWC strength less than 3,500psi (approximately P05). In January 2015, when the FTHP dropsto 100 psi (with predicted reservoir pressure of 1,340 psi), sand-ing would be expected from approximately 40% of the horizontalperforations. In contrast, no sanding is expected from perforationsat the top and bottom of the holes even at the final reservoir pres-sure (P

r

≈ 400 psi), while at this pressure, approximately 90% of

the side perforations are expected to produce sand.

Y-7 Well. This well had been producing sand free since May 2004from an initial reservoir pressure of approximately 2,900 psi. In Janu-ary 2006, depletion and drawdown pressures were approximately250 and 200 psi, respectively. The log-derived TWC rock-strengthprofile over the perforated zones indicates that TWC-strength valuesrange between 3,300 (P10) and 5,800 psi (P95), with an average of4,100 psi (P50).

Fig. 19 shows the effect of perforation orientation on sandingrisk with the well trajectory at 6,500 ft TVDSS for rocks with aTWC strength of 4,100 psi (P50) as function of reservoir pressure.The well deviation and azimuth at this depth are 24° and 64°,respectively. Perforations orientated at 25–30° from top and bottomof the hole are the most stable. All perforations will be unstable atthe final reservoir pressure, 550 psi. For a given reservoir pressure,these perforations would have a critical drawdown approximately150 psi greater than those located on the side of the wellbore(worst perforation direction). This relatively small difference forthe CDDs between the worst and best perforation orientations isbecause of the low deviation angle for this well. A comparison ofthe perforation-optimization plots for Y-7 (Fig. 19) and Y-16 (Fig. 16)

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Depl.=0 psi

Depl.=500 psi

Depl.=1000 psi

Depl.=2000 psi

Fig. 16—Perforation optimization for Y-16 well.Fig. 17—Load-factor log for Y-16 well with January 2006 condi-tions.

Fig. 18—Sand-free operating envelope for Y-16 well. Fig. 19—Perforation optimization for Y-7 well.



indicates that the higher the hole deviation angle, the greater thedifference between the CDDs for the best and the worst perfora-tion orientations.

Fig. 20 shows LF vs. depth plots for perforations at sides and topand bottom of the hole under the reservoir-pressure and drawdownconditions as of January 2006, indicating sand-free production forthe perforated zone ( LF < 1). This is consistent with observationof no sand in this well. The sand-prone zones (e.g., 7,000–7,030 ftand 7,065–7,200 ft) were not perforated following a selective per-foration strategy. The results in Fig. 20 supports the wisdom of theselective-perforation strategy (i.e., LF >1 in these intervals impliesthat they would have produced sand if they were perforated).

The sand-free envelope plot for the best (top and bottom) andworst (side) perforation orientations for a series of rock strengthsfrom 3,000 to 5,000 psi revealed that no sanding from side per-forations would be expected until the reservoir pressure drops toapproximately 2,350 psi (approximately mid-year 2008) whenminor sand production may occur in rocks with a TWC strength

less than approximately 3,300 psi (approximately P10). Sandingwould be expected from approximately 50% of the side and fromapproximately 30% of the top and bottom perforations of the holewhen the FTHP drops to 100 psi in January 2014 (with predictedreservoir pressure of 1,450 psi).

Conclusions

The paper has presented a methodology for sand-production pre-diction from arbitrarily oriented wells with openhole and cased-and-perforated completions.

The methodology has been applied to the Y field to assesssanding risks in existing wells and to guide in minimizing sand-ing risk in future infill wells. Sanding-evaluation results using thepredictive model presented in this paper matched the observations

in the Y field of no sanding as of year 2006. Sanding evaluationswere performed over the life of field conditions from current tofinal (abandonment) reservoir pressures and at pressure conditionscorresponding to the decrease of FTHP to 100 psi during the welllife for all existing production wells.

Sanding evaluation indicates that the likelihood of sand pro-duction in the Y field increases with further pressure depletion forall existing wells. Therefore, some type of sand control would beneeded for producing wells in the area. With the planned BHFPand depletion rates, modeling results indicated that some of thewells may start very minor sanding as early as 2008 (e.g., Y-2,-11, -17, and -18) from perforations along their respective worst

directions. Even this predicted minor sanding may be delayedfurther because conservative boost factors have been used in themodel in the absence of credible data (e.g., sand-influx-test data)for model calibration. At the time of this revision (mid-2009), nosanding has been observed at the surface. This is generally consis-tent with the model prediction; however, definitive conclusions stillcould not be made because the reservoir-pressure and drawdowndata were not supplied for further analysis. Also, because of thelack of sophisticated sand-detection tools in the wells, it was notpossible to be certain whether sand has started failing downhole.Sand may be failing, but the solids produced are not being liftedby the well fluids to the surface. So far, any production decline isnot attributed to the well filling with solids. Depending on welltrajectories, rock strength, and perforation orientations, sanding

would start later for other wells (e.g., Y-7, -13, -16). With the givenreservoir-pressure conditions and planned production conditionsfor the field, the sanding risk in all producing wells in the Y fieldcan be summarized in Table 1.

The sanding evaluations summarized in Table 1 are for theleast-stable perforation orientations (i.e., perforations at the sideof the hole in deviated wells and parallel to the H azimuth invertical wells). The exact timing of sand production in existingwells cannot be determined because of variations in rock-strengthprofile and perforation orientation. The effect of water productionon sanding potential has not been accounted for in these analyses.Water production may exacerbate sanding, particularly in weakerrocks and rocks with reactive cementation. The effect of waterproduction on sanding is not considered in this study.

For the current reservoir and drawdown pressures, there are

possible well paths and perforation orientations that may be usedto delay the requirement and installation of sand control in newproduction wells. For openhole (barefoot) completion, verticalwells and deviated wells in the northwest/southeast direction (90°from the H azimuth) can produce sand free longer because thesedeviated wells will be subject to the minimum deviatoric stress( v – H ). For cased-and-perforated completions, deviated wellswith perforations near to the top and bottom of the hole can pro-duce sand free longer. Perforation optimization in terms of criticaldrawdown with two deviated well trajectories has shown that themagnitude of difference in drawdowns between the best perfora-tion orientation and the worst perforation orientation dependssignificantly on the well trajectory. For higher pressure depletionsand/or higher flow rates (drawdowns), sanding would be expected

Fig. 20—Load-factor plots for Y-7 well as of January 2016.

TABLE 1—SUMMARY OF SANDING EVALUATION RESULTS FOR PRODUCING WELLS IN Y FIELD

Well

InitialP r (psi)

Current (late2006) P r (psi)

Pr in psi WhenFTHP=100 psi

(date)

Percentage ofUnstable Perfs When

FTHP=100 psi (%)

Final P r (psi)

Percentage ofUnstable Perfsat Final P r (%)

Onset of WaterProduction

(year)

Y-2 3,120 2,570 1930 (July 2011) 35 700 60 –

Y-7 2,925 2,550 1450 (Jan. 2014) 45 690 75 Mid 2021

Y-8 3,100 2,560 1800 (Jan. 2012) 60 447 75 Mid 2017

Y-11 2,880 2,590 1600 (Jan. 2016) 40 750 50 Mid 2017

Y-13 2,980 2,720 1800 (Apr. 2013) 10 550 55 Mid 2006

Y-16 2,860 2,500 1340 (Jan. 2015) 60 400 90 –

Y-17 3,125 2,450 2000 (Jul. 2011) 40 525 90 Mid 2015

Y-18 3,120 2,490 1800 (Aug. 2011) 50 425 90 –


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for all wells in the Y field. Therefore, some type of sand controlwould be needed for new production wells in the field.

The optimum sand-control mechanism for existing and futurewells in the Y field could vary from well to well, depending onrock-strength profile, timing of water production, well trajectory,and production plan. A combination of passive (such as recomple-tion for selective and oriented perforation) and active sand controlis currently under investigation for the existing wells.

Although the sanding-evaluation results with the default boostfactors are consistent with the no-sanding observation in existingwells, model predictions could be improved further with the resultsof beanup trials and observation of sanding from any of the study

wells. Results from such a model that is calibrated by beanup trialsshould help decide the timing of sand-control installation.

Acknowledgments

The authors acknowledge the contribution of Richard McNaught,senior software engineer, Helix RDS, for the development of com-puter code and graphics outputs. Phil Taylor also deserves thanksfor internal peer review of this paper. Finally, two anonymousreviewers deserve thanks for providing very constructive com-ments to address, which has raised the standard of this paper.

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doi: 10.2118/56017-PA.

Khaksar, A., Taylor, P.G., Fang, Z., Kayes, T., Salazar, A., and Rahman, K.

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Go. Paper SPE 121972 presented at the EUROPEC/EAGE Conference

and Exhibition, Amsterdam, 8–11 June. doi: 10.2118/121972-MS.

McNally, G.H. 1987. Estimation of coal measures rock strength using

sonic and neutron logs. Geoexploration 24 (4–5): 381–395. doi:

10.1016/0016-7142(87)90008-1.

McPhee, C.A., Lemanczyk, Z.R., Helderle, P., Thatchaichawalit, D., and

Gongsakdi, N. 2000. Sand Management in Bongkot Field, Gulf of

Thailand: An Integrated Approach. Paper SPE 64467 presented at the

SPE Asia Pacific Oil and Gas Conference and Exhibition, Brisbane,

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Moos, D. and Zoback, M.D. 1990. Utilization of observations of well bore

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Khalil Rahman is a Principal Geomechanics Specialist withBaker RDS. His teaching, research, and consulting career spansover 22 years. In the past, he held senior lecturer positions atthe University of New South Wales and later at the Universityof Western Australia, where he taught various petroleum engi-neering courses, led curricula development in petroleum engi-neering, and conducted research and consulting activities inhydraulic fracturing optimization, wellbore-stability assessment,and sand-production prediction areas. In his university andindustry positions, Rahman has completed numerous researchand consulting projects in petroleum geomechanics, published65 peer-reviewed journal and conference papers, offeredindustry short courses, and supported technology and soft-ware development activities. He obtained his PhD degree from

the University of Newcastle Upon Tyne, UK, is an active mem-ber of SPE, and serves various SPE committees. Abbas Khaksar is a Principal Geomechanics Specialist with Baker RDS, with 19years of industry and research-and-development experience inpetroleum rock mechanics, petrophysics, and exploration. Hehas a BS in mining engineering from Tehran University (1989), andan MS in petroleum geology and geophysics (1994), and a PhDin rock physics (1998) from the University of Adelaide. Khaksarhas worked extensively in petroleum geomechanics, involved inmore than 100 projects in many areas of geomechanics. Before joining RDS, he was with Geomechanics International (GMI)(2000–2005) as a petrophysicist, geomechanics specialist, andmanager of consulting services in Asia Pacific. Khaksar was apost-doctoral researcher at the University of Adelaide (1999–2000), working on pore-pressure prediction and CO2 sequestra-tion. He is a short course instructor and has about 50 publica-

tions and conference presentations in reservoir geomechanicsand rock physics, and he is a member of SPE and FesAus. TobyKayes is a principal geomechanics specialist and a senior man-ager with Baker RDS based in Kuala Lumpur. He has 16 yearsof industry experience in all areas of petroleum geomechanicsincluding drilling optimization, wellbore stability, pore-pressureprediction, stress sensitive reservoir modeling, hydraulic fractur-ing, and sand management. This is based on sound soil androck mechanics experience Kayes gained as a civil engineerworking for Arup before joining the petroleum industry. He hasalso worked as a wireline engineer, log analyst, research scien-tist, and consultant for Schlumberger, specializing in petroleumgeomechanics. Kayes obtained his BE in civil engineering fromthe University of Bristol and researched sand production physicsat the University of Cambridge with a Schlumberger CambridgeResearch award. He is a chartered petroleum engineer and

chartered European engineer. Kayes is a member of the SPEcommittee for sand management, the external advisory boardfor the Center of Integrated Petroleum Engineering and Geosci-ence at the University of Leeds, and the Energy Institute.

1.Chapter 6_SPE-116633-PA_An Integrated Geomechanical and Passive Sand Control Approach

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