Abstract: The STADYN program was developed for the analysis of driven piles both during installation and in axial loading. Up until now the test cases used were in predominantly cohesive soils. In this paper, the expansion of the program’s use into predominantly cohesionless stratigraphies has required consideration of two important factors. The first is the difference between strain-softening in clays as opposed to sands, and additionally static vs. dynamic strain effects. This requires a review of the whole concept of the “magical radius” for pile elasticity. The second is the effect of dilatancy on the response of the pile to axial loading. Both of these are discussed in this paper, and test cases are presented to illustrate the application of the program to actual driven piles.
Abstract: Although it is widely understood that soils are non-linear materials, it is also common practice to treat them as elastic, elastic-plastic, or another combination of states that includes linear elasticity as part of their deformation. Assuming hyperbolic behavior, a common way of relating the two theories is the use of strain-softened hyperbolic shear moduli. Applying this concept, however, must be done with care, especially with geotechnical structures where large stress and strain gradients take place, as is the case with driven piles. In this paper a homogenized value for strain-softened shear moduli is investigated for both shaft and toe resistance in clays, and its performance in the STADYN static and dynamic analysis program documented. A preliminary value is proposed for this “average” value based upon the results of the program and other considerations.
In our last STADYN post we discussed the addition of factors to take into account adhesion phenomena with cohesive soils. In this post the addition of a more mundane but nevertheless important parameter for impact pile hammer systems is done: consideration of plastic losses in hammer and pile cushions, and interfaces as well.
Most impact pile hammers use some kind of hammer cushion; additionally, concrete piles are almost always driven with pile cushion at the pile head. Cushions of both kinds are subject to significant plastic deformation and generation of heat. There are several possibilities of modelling these elements in a simulation such as STADYN.
The first is to use velocity-dependent (viscous) damping to simulate the dissipation of energy. STADYN in its current form has no velocity-dependent parameters; to add these would involve some major changes in the code, and in any case the testing of cushion material does not produce a result that would indicate such a property.
The second is to use an elastic-purely plastic approach similar to the one used in the soils. The problem with this is that it would “flat-top” the impulse to the pile, and there is no evidence that the cushion material fails in this way.
The third is to use a “coefficient of restitution” approach, where the rebound of the cushion takes place at a different stiffness than the compression. This is illustrated in two variants below.
Cushion Models for Plastic Cushions (from Warrington (1988))
The conventional model dates back to Smith, and is still used in GRLWEAP. The ZWAVE model is described by Warrington (1988). In both cases the energy lost in the cushion is represented by the shaded area.
For STADYN the conventional model was adopted. Implementing this took a little more care in a finite element code than in finite-difference codes like WEAP and GRLWEAP but it was done. To accomplish this, it was necessary to compute the force in the cushion incrementally, as with plasticity the response is now path-dependent. When the cushion rebounded (i.e., the distance between the cushion faces increased from one step to the next) the rebounding stiffness is used. In this way multiple rebounds can be modelled properly.
Since the inverse methods do not model the hammer, the Mondello and Killingsworth case is not considered here. This leaves the other two cases, and these can be summarised very briefly.
The Finno (1989) case had a blow count increase from 15.8 to 17.0 blows/30 cm. For the SE Asia case, the blow count increased from 11.8 to 13.5 blows/30 cm. Additionally for the latter case comparisons with the pile head force and ram velocity vs. time tracks were produced.
Pile Head Force
The pile head force until peak was identical, and then decreased more rapidly afterwards. There was an additional “kick” at 2L/c not present in the previous run.
Ram Point Velocity
The ram (point) velocity is the same until rebound, and then the ram is essentially stationary with the coefficient of restitution until 2L/c, after which the ram velocity in the two cases is very close. The sawtooth effect is mostly due to the “ringing” of the ram, i.e., a stress wave going up and down the ram.
While it is evident that the method of energy transfer is different with the addition of the coefficient of restitution, the actual effect of plasticity on the blow count is not great. This is probably due to two factors: most of the energy transfer takes place during compression of the hammer cushion, and both hammers are using micarta and aluminium, which has a relatively high coefficient of restitution (0.8). Nevertheless cushion losses are greater in materials such as plywood, which is used with concrete piles. It is to this type of pile that STADYN’s development now turns.
One of the things that makes STADYN more complex than either TAMWAVE or most other 1-D solutions is that soils are not considered as purely cohesive or cohesionless. In most analysis of driven piles, soils are either on or the other, or at best alternately layered. In reality the division between the two is not so clear-cut except for either clean sands on the one end or pure clays on the other. STADYN’s soil system envisions soils as a continuum between one and the other; although this adds to the flexibility of the program (especially in the inverse mode) and its modelling of reality, it makes specifying soils a challenge.
As noted earlier, for soils between the purely cohesionless () and cohesive () interpolation is done so that soils have no cohesion in the former case, no friction in the latter, and are interpolatively mixed in between. For example, for a middle case of , the soil would have a reduced cohesion and friction for the same value of and share these properties. In this way any adjustments for adhesion of either type of soil would be made for each.
Cohesionless soils: there are two ways of looking at this problem. We can assume a straight-up Coulomb friction failure between the pile and soil, or we can assume that the pile acts as a “direct shear” tester and thus forces the soil to fail at an apparent angle that is not the same as would be predicted by Mohr-Coulomb failure. As with TAMWAVE, we have assumed the latter; this is explained in some detail here. It is reasonable to assume that a continuum model such as is used by STADYN could predict such a failure; thus, no modification to the elements closest to the pile surface is done for cohesionless soils.
One thing that did change, however, was the way the lateral earth pressure on the pile was computed. In an elastic-purely plastic system, lateral earth pressure varies in the elastic region, and with elastic theory that means with the variance of Poisson’s Ratio. With a Mohr-Coulomb failure criterion, frictional cohesionless soils’ strength is mobilised by vertical effective stress acting laterally. In recent code iterations Jaky’s Equation has been used to estimate Poisson’s Ratio; however, this has been changed to use the method given by Randolph, Dolwin and Beck (1994). Once the lateral earth pressure coefficient is computed using this method, Poisson’s Ratio is determined. At or below the pile toe Jaky’s Equation is used.
Cohesive soils: Mohr-Coulomb theory has no way of taking degradation of cohesion at an adhesion surface into account. To do this the cohesion for the element(s) immediately adjacent to the pile is reduced by an factor as computed by the method of Kolk and van der Velde. This is only done for the element immediately adjacent to the pile shaft surface. This is the way STADYN does a pile-soil “interface.” Doing it in this manner obviates the need for special interface elements between the pile and soil.
Implementing this is a little tricky, because the factor is dependent upon the effective stress. It is necessary to thus generate the layers, compute the mid-point effective stresses in each, and then apply the factor to the cohesion of one set of elements only.
Results: Finno (1989) and Modello and Killingsworth (2014) Comparisons
For the first case, the Davisson capacity changed from 971 kN to 965 kN and the blow count from 17.6 to 15.8 blows/30 cm.
For the second (inverse) case, the Davisson capacity for the case of the Davisson capacity changed from 269 kN to 274 kN and the blow count from 24.6 to 24.4 blows/30 cm. The least squares difference actually increased from 0.00143 to 0.00149.
In both cases the soils were heavily cohesionless (at least that’s the way the pile looked at them) and the reduction in adhesion was minimal in impact.
Results: Notional Southeast Asia Case
Of all the test cases in the original study, the notional Southeast Asia case was the most problematic in the results, especially as they were compared to the GRLWEAP output. The previous phase produced little difference in outcome; it was hoped that applying factors to the adhesion would at least solve the discrepancies of SRD estimates. The results did not disappoint.
Since we have not presented too many results from this case, some graphical output is in order. First, the force-time and velocity*impedance-time curves:
Force and Impedance*Velocity vs. Time Curves
The result above is a classic “offshore” pattern. In the early part of impact () both the actual pile head force and the product of the impedance are virtually identical. This indicates an “infinite pile” condition; the theory behind this is discussed by Warrington (1997). Beyond this the two diverge; first the pile head moves upward in rebound from the pile shaft (indicated by the fact that the rebound takes place before ) and impacts the pile cap, producing a secondary force in the pile head. Beyond the pile head force goes to zero and the velocity oscillates with the reflections from the pile; however, just after that time the compressive “kick” from the toe is evident.
Static Load Test Results
Now we have the result of the static load test. As noted in the original study, static load tests are exceptional offshore, and for actual loading a tension test is probably of just as much interest (if not more) than a compressive one. In the original study doubt was also expressed as to the relevance of Davisson’s criterion to offshore piles; the variation among different interpretation methods, however, were not that great. In any case, the effect of reducing the adhesion of cohesive soils along the surface is evident: the Davisson ultimate load has dropped to 20,600 kN. This is nearly identical to the Dennis and Olson (1983) method result, and below the API RP2A (2002) result. This indicates that the application of the method to the soil elements along the pile shaft results in bringing the static results of STADYN more in line with those of static methods in use.
For the Dennis and Olson (1983) SRD, the GRLWEAP blow count varied from 18.4 blows/30 cm to 21.8 blows/30 cm, depending upon which value of damping was used ( or . STADYN returned a blow count of 11.8 blows/30 cm. This is a significant improvement. There are two possibilities to explain the remaining difference:
STADYN is modelling a lower effective damping value for the soils than is used in GRLWEAP. As noted in the original study, STADYN has a different model for handing dissipative phenomena than GRLWEAP.
The two programs have differing methods for arriving at the blow count.
Before we can make more definitive statements about this, we need to include cushion losses, which is our next step. Nevertheless this result clears up a great deal of the difficulty with this case in the original study.
In the last post we discussed the change in cohesion in the interpolation. Now we apply these to the program to see how they change the results.
In the original study there are three test cases, as noted here. For this stage the first and the third will be considered. The results for the second (after adjustments for changes in the interpolation) are little different from before; we will consider major changes to impact that case in our next round.
That leaves the first and third. For the first–comparison with a static load test–it was necessary to readjust the values for to reflect changes in the meaning of relative to soil properties, as discussed in the last post. The changes made resulted in layering with the following characteristics:
Layer
Bottom y-coordinate, m
1
5.18
-1
0
2
7.32
-1
0
3
15.2
0.5
0
4
30.5
0.5
0
The static load test results can be seen below.
Comparison with the original study show little improvement in the failure load correlation; however, the load-deflection relationship before failure is significantly improved. The system was much softer in the original study and the improvement reflects the better estimate of the elastic modulus. The blow count is very similar to the original study.
Turning to the inverse case that was also discussed in Warrington and Newman (2018), based on the previous results, it was decided to drop consideration of 1-norm results. The results are reproduced in outline below for the three cases run (refer to Warrington and Newman (2018) for details.)
Limiting
Difference Sum
Static Load, kN
Average Shaft
Average Shaft
Toe
Toe
+-1
0.0034
673
0.084395
-0.46825
-0.992
0.758
+-2
0.00327
449
-0.455
-1.3375
-0.57
-0.727
+-3
0.00143
269
-0.26775
-2.1775
0.113
1.12
Comments:
The difference sum for the highest variation was the best match we have obtained to date for this case; the velocity match is shown below. The situation around L/c = 1.5 is still difficult but the rest of the correlation is improved.
The static load decreases with broader variation and a lower difference sum. This is different than Warrington and Newman (2018), where the static load “settled down” to a close range of values.
The plot above does not reflect that, on the whole, the variation of along the shaft was less than experienced in the past, especially with the +-3 run. The stratigraphy of the site suggested relatively uniform soil properties along the shaft, and this is beginning to be seen in the inverse results.
The +- 3 run was very heavily “toe weighted” in resistance.
While both of these cases show progress, the time has come to consider the whole issue of pile-soil interface issues, which will be considered in our next updates.
factors to take into account adhesion phenomena with cohesive soils. In this post the addition of a more mundane but nevertheless important parameter for impact pile hammer systems is done: consideration of plastic losses in hammer and pile cushions, and interfaces as well.
soil property system
) and cohesive (
) interpolation is done so that soils have no cohesion in the former case, no friction in the latter, and are interpolatively mixed in between. For example, for a middle case of 
, the soil would have a reduced cohesion and friction for the same value of 
and share these properties. In this way any adjustments for adhesion of either type of soil would be made for each.
the Davisson capacity changed from 269 kN to 274 kN and the blow count from 24.6 to 24.4 blows/30 cm. The least squares difference actually increased from 0.00143 to 0.00149.
) both the actual pile head force and the product of the impedance are virtually identical. This indicates an “infinite pile” condition; the theory behind this is discussed by 
) and impacts the pile cap, producing a secondary force in the pile head. Beyond 
or 
. STADYN returned a blow count of 11.8 blows/30 cm. This is a significant improvement. There are two possibilities to explain the remaining difference:
to reflect changes in the meaning of 
vulcanhammer.net 